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CONNECTOR DEVELOPMENT FOR HYBRID MASONRY SEISMIC STRUCTURAL SYSTEMS Seth R. Goodnight Gaur P. Johnson and Ian N. Robertson Research Report UHM/CEE/11-03 May 2011
CONNECTOR DEVELOPMENT FOR
HYBRID MASONRY SEISMIC STRUCTURAL SYSTEMS
Seth R. Goodnight Gaur P. Johnson
and Ian N. Robertson
Research Report UHM/CEE/11-03 May 2011
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ABSTRACT
Hybrid masonry construction has recently been developed as a means to provide
a more efficient use of masonry infill walls in steel frame structures by designing these
walls to function as a part of a structure’s lateral force resisting system. This system
has current applications in low seismic regions (Seismic Design Categories A, B, and
C). In regions of higher seismicity (Seismic Design Categories D and greater), there is a
need for a clear understanding of the mechanics of the interaction between the steel
frame and the masonry shear wall, as well as the development of specific detailing
requirements to provide for a ductile and somewhat predictable response in a seismic
event of significant magnitude.
The scope of this report is the investigation of the connection between the steel
beam and the top of the masonry wall for a Type I hybrid masonry system. The
proposed steel connector plates are through-bolted to the masonry wall bond beam and
welded to the bottom flange of the steel beam above. Preliminary guidelines and
recommendations are provided for the development of energy dissipating hybrid
masonry connector plates. The first portion of this study focused on the development of
connector plate designs to determine which provided the most stable, ductile cyclic
response. The second portion of this study investigated the strength and limit states of
the through-bolted connection to the masonry wall bond beam under shear loading.
Based on these experiments, recommendations are provided for the most suitable
connector plate designs as well as a general evaluation of the connection to the
masonry wall and potential limit states to consider in future design applications. For a
“rigid” connection between the steel frame and CMU wall, link connectors with a
thickness greater than 0.5 inches are recommended, whereas fuse Type T was proven
to be superior for a ductile, energy dissipating connector plate. ACI 530-08 design
values for bolt shear yielding and masonry shear failure were determined to be the most
appropriate limit states for determining an approximate design value for the through-
bolted CMU wall connections.
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ACKNOWLEDGEMENTS
This report is based on an MS thesis prepared by Seth Goodnight under the
direction of Drs. Gaur Johnson and Ian Robertson. The authors would like to thank Drs.
Ronald Riggs and David Ma for their contributions as part of the Thesis Committee for
the project.
Special thanks to Tileco Inc. and Bonded Materials Co. for their generous
donation of materials used in this project, as well as to Mitch Pinkerton and Miles
Wagner of the University of Hawai‘i at Mānoa’s Structural Testing Laboratory for their
assistance throughout the project.
This research was supported by the National Science Foundation under Grant
No. CMMI 0936464 as part of the George E. Brown, Jr. Network for Earthquake
Engineering Simulation. This support is gratefully acknowledged.
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TABLE OF CONTENTS
ABSTRACT ..................................................................................................................... i
ACKNOWLEDGEMENTS ............................................................................................... ii
1 INTRODUCTION ....................................................................................................... 1
1.1 INTRODUCTION ...................................................................................................... 1 1.2 OBJECTIVE ............................................................................................................ 1
2 LITERATURE REVIEW ............................................................................................. 2
2.1 HYBRID MASONRY LATERAL LOAD RESISTING SYSTEM ............................................. 2 2.2 CONNECTOR PLATES ............................................................................................. 4 2.3 CMU THRU-BOLT PUSH-OUT TESTS ....................................................................... 5
3 APPROACH .............................................................................................................. 7
3.1 CONNECTOR PLATE TESTS ..................................................................................... 7 3.1.1 Experimental Setup .................................................................................................................. 7 3.1.2 Test Specimens ........................................................................................................................ 8 3.1.3 Procedure ............................................................................................................................... 10
3.2 CMU THRU-BOLT PUSH-OUT TESTS ..................................................................... 12 3.2.1 Experimental Setup ................................................................................................................ 12 3.2.2 Test Specimens ...................................................................................................................... 15 3.2.3 Procedure ............................................................................................................................... 16
3.3 FULL-SCALE STEEL-MASONRY SUBASSEMBLIES ..................................................... 16
4 RESULTS ................................................................................................................ 17
4.1 SHEAR CONNECTOR TESTS .................................................................................. 17 4.1.1 Test Results Summary ........................................................................................................... 17 4.1.2 Specimen Behavior Narratives ............................................................................................... 20
4.2 CMU THRU-BOLT PUSH-OUT TESTS ..................................................................... 27 4.2.1 Specimen Behavior and Failure Mechanisms ........................................................................ 27 4.2.2 Test Results ............................................................................................................................ 30
5 ANALYSIS .............................................................................................................. 33
5.1 COMPARISON OF CONNECTOR PLATES (LINK AND FUSE TYPES A, S, AND T) ............ 33 5.2 ATC-24 CONNECTOR PLATE ANALYSIS ................................................................. 34 5.3 CMU THRU-BOLT PUSH-OUT TESTS ..................................................................... 44
5.3.1 Connection Failure Mechanisms ............................................................................................ 45 5.3.2 Comparison of Wall Specimens (Partially vs. Fully Grouted) ................................................. 48
6 CONCLUSIONS / RECOMMENDATIONS .............................................................. 49
7 REFERENCES ........................................................................................................ 51
APPENDICES (A, B, C, D)
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TABLE OF FIGURES
Figure 1. Three proposed types (Type I, Type II, and Type III, shown clockwise from top left) of hybrid masonry systems, classified according to types of loads transferred to masonry walls [Biggs, 2011]. ..................................................................................... 2
Figure 2. Current state of connector plate detailing requirements as provided by the International Masonry Institute. .................................................................................. 5
Figure 3. Connector plate test setup with specimen highlighted. Lateral load is applied to the steel tube section at the top of the image. ....................................................... 7
Figure 4. Typical link and fuse connector details (in = 25.4 mm). .................................... 9
Figure 5. Graphic representation of the proposed full-scale test structure to be constructed at UIUC. ................................................................................................ 10
Figure 6. Modified ATC-24 cyclic loading routine for steel connector plates, with three cycles at each level of lateral displacement. ............................................................ 12
Figure 7. CMU bolt push-out test concept. .................................................................... 13
Figure 8. CMU bolt push-out test setup. ........................................................................ 14
Figure 9. Schematic drawing showing the dimensions and layout of the test assembly for loading of the CMU wall specimens. ................................................................... 14
Figure 10. (a) CMU block configuration "1", (b) CMU block configuration "2". .............. 15
Figure 11. (a) Non-Ductile Weld Rupture: P4_T4-01; (b) Rupture of Base Metal: P4_T4-02. ............................................................................................................................ 18
Figure 12. (a) Yielding then buckling: P4_T2-01; (b) Yielding/Fatigue rupture: P4_T4_FT3-01. ........................................................................................................ 19
Figure 13. Bent plate behavior, ductile response followed by C-shaped weld rupture: P4_T4_FT3-B-01. .................................................................................................... 19
Figure 14. P4_T2-01, showing typical buckling failure. ................................................. 20
Figure 15. P4_T4-02 showing weld failure and base plate rupture. .............................. 21
Figure 16. Specimens P4_T4_FA1-01, _FA2-01, and _FA3-01 showing typical fuse base rupture (shown from left to right). .................................................................... 22
Figure 17. Fuse roughness and early rupture of P4_T4_FA2-P-01 due to cutting with a hand-held plasma torch. .......................................................................................... 23
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Figure 18. P4_T4-FS-01 showing significant buckling at ±2.00” displacement. ............ 23
Figure 19. P4_T4-FT2-01 displaying a combination of lateral-torsional buckling and rupture due to fatigue. .............................................................................................. 24
Figure 20. P4_T4-FT3-01 displaying uniform distribution of yielding along entire fuse section. .................................................................................................................... 24
Figure 21. Specimens P4_T4_FT2-01 and _FT3-01, showing buckling and tapered fuse base rupture (shown from left to right). .................................................................... 25
Figure 22. P4_T4_FT3-B-01 showing rupture of C-shaped fillet weld on bent leg. ....... 25
Figure 23. P4_T6_FA2-01 showing typical fuse base rupture. ...................................... 26
Figure 24. (a) Final cracking patterns for specimen FG-1A; (b) Typical cracks around rod bearing area as seen in specimen FG-1A. ........................................................ 27
Figure 25. (a) Final cracking patterns for specimen FG-2A; (b) Typical threaded rod shear deformation as seen in specimen FG-2A. Load values shown on specimens were rounded to the nearest whole number. ............................................................ 28
Figure 26. Specimen PG-1B showing mortar joint separation along the entire height of the right side of the middle blocks, typical of partially grouted specimens. .............. 29
Figure 27. Final cracking patterns of specimen PG-2A, with vertical mortar joint separation visible above the location of the threaded rod. ....................................... 30
Figure 28. Load vs. Displacement for all CMU specimen "A" tests. .............................. 32
Figure 29. Load vs. Displacement for all CMU specimen "B" tests. .............................. 32
Figure 30. Assumed projected shear area for masonry breakout failure (in = 25.4 mm).47
Figure 31. Tapered fuse design dimensions. ................................................................ 66
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TABLE OF TABLES
Table 1. Seismic Design Values for Reinforced Masonry Shear Walls. .......................... 3
Table 2. Connector plate specimen naming convention (in = 25.4mm). .......................... 9
Table 3. Actuator displacement summary. .................................................................... 11
Table 4. Critical Points of Hysteretic Response (in = 25.4 mm, kip = 4.45 kN, kip-in = 113 J). ...................................................................................................................... 17
Table 5. ASTM E8 tension test results for connector plate steel bar (inch = 25.4mm, ksi = 6.895 MPa). .......................................................................................................... 18
Table 6. CMU bolt push-out maximum load values and corresponding downward deflection (in = 25.4 mm, kip = 4.45 kN). ................................................................. 31
Table 7. Summary of results from CMU prism tests. ..................................................... 31
Table 8. Specimen Cumulative Displacement and Total Energy Dissipation Values (in = 25.4 mm, kip-in = 113 J). ......................................................................................... 33
Table 12. Summary of Applicable ACI 530-08 Design Values and Measured Test Results (kip = 4.45 kN)............................................................................................. 47
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1 Introduction
1.1 Introduction
There currently exist a wide variety of structural systems designed to resist lateral seismic and wind load, each system presenting its own unique set of advantages and disadvantages. Some of the most common systems in use today are steel moment frames, shear walls, and braced frames. In steel frame construction, architectural masonry infill walls are often present to enclose the building envelope but do not traditionally serve any structural purpose. Hybrid masonry, a relatively new concept in seismic load resisting building systems that has already been implemented in several projects in areas of low seismicity [Abrams et al., 2010], serves to incorporate the masonry infill walls, designed and reinforced as shear walls, as a specifically-designed structural component of the building’s seismic force resisting system. A Type I hybrid masonry system uses steel connector plates to transfer only lateral forces from the steel frame to the top of the masonry wall. Type II hybrid masonry extends the purpose of the masonry wall to resisting both lateral forces and vertical compressive forces by eliminating the gap between the steel frame and the top of the wall and connecting the two components with a system of headed studs across the top of the masonry wall. Finally, Type III hybrid masonry contributes an additional load transfer mechanism by providing a similar headed stud connection along both sides of the masonry wall to resist vertical shear forces. The intent of this project is to investigate the incorporation of masonry infill walls into the lateral load resisting system, and more specifically, to develop a steel-masonry connector suitable for this purpose. Investigation will concentrate primarily on the development of a structural “fuse” to serve as a ductile link to dissipate seismic energy and enhance seismic performance of the lateral system. This project is part of a larger on-going NEESR-SG project under Principal Investigator Daniel Abrams of the University of Illinois (UIUC). Ductile fuses have been studied and tested by others for various applications, such as [Bruneau & El-Bahey, 2010] and [Aliaari & Memari, 2007].
1.2 Objective
The scope of this particular study is limited solely to the investigation of the Type I hybrid masonry system and will focus on the development of a steel-masonry connection using steel connector plates through-bolted to masonry walls to transfer horizontal shear forces from the steel frame into the masonry walls.
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2 Literature Review
2.1 Hybrid Masonry Lateral Load Resisting System
The hybrid masonry system, first proposed in 2006 [Biggs, 2006], incorporates the traditional, non-structural masonry infill wall into the lateral load resisting system of a structure through design of the infill wall as a masonry shear wall and application of a series of shear connector plates between the infill wall and the steel frame of the building. Hybrid masonry is classified into three categories, shown in Figure 1, based on the degree of confinement of load transfer between the steel frame and masonry wall [Biggs, 2011].
Figure 1. Three proposed types (Type I, Type II, and Type III, shown clockwise from top left) of
hybrid masonry systems, classified according to types of loads transferred to masonry walls [Biggs, 2011].
For the Type I system, the masonry wall is designed only for in-plane and out-of-plane loads, with a gap between the top of the masonry panel and the bottom of the steel beam above to prevent transfer of gravity loads to the wall. In this case, connector plates have also been detailed with vertical slotted holes to eliminate any vertical load transfer through these connections. In the Type II system, the masonry wall is designed for the in-plane and out-of-plane loads of a Type I system, but the gap at the top of the
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panel is not present, requiring the wall to serve additionally as a vertical loadbearing wall. In this case, hybrid masonry can serve to prevent progressive collapse as a means to increase redundancy in the structure. Finally, the Type III system, although still in the early stages of development, will provide for additional connectivity between the ends of the masonry panel and the wide flange steel columns, as a means to provide a greater degree of masonry confinement and more direct transfer of overturning moments.
The International Building Code [IBC, 2006] and The Building Code Requirements and Specifications for Masonry Structures [MSJC, 2008] currently provide for three different classifications of masonry shear walls: ordinary reinforced, intermediate reinforced, and special reinforced. Each of these classifications is assigned different seismic factors that are used for the overall design of the lateral force resisting system. These factors, R (response modification coefficient), Ω0 (system over-strength factor), and Cd (deflection amplification factor), are given in Table 1 for each of the classifications of masonry shear walls. Current applications of hybrid masonry systems have been designed using the seismic forces developed according to the IBC and the design values for reinforced masonry shear walls.
Table 1. Seismic Design Values for Reinforced Masonry Shear Walls.
To date, hybrid masonry systems have been implemented in Seismic Design Categories A, B, and C. Extension of hybrid masonry load resisting systems into higher SDCs is currently under investigation as the requirements of structural steel construction fall under the AISC Seismic Design Manual for SDC D and E [IMI, 2009]. Computer models have been developed for Type I hybrid masonry systems using Bentley RAM Elements and also with linear shell models to confirm the assumed behavior of the steel frame and masonry shear wall assembly [Abrams et al., 2010]. Based on the results from a comparison of these models with equations derived from statics, it was suggested that an appropriate design rule was to proportion steel columns to resist the axial forces resulting from total overturning moments and to proportion masonry walls to resist the total story shears. The total shear in the masonry wall and the column axial force can be represented by Equation [1] and Equation [2], respectively.
[1]
Wall Classification R Ω0 Cd
Ordinary 3 2.5 2.25
Intermediate 4 2.5 2.5
Special 5.5 2.5 4
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[2]
In these equations Fj is the lateral force at level j, n is the total number of stories in the building under consideration, Vk is the story shear in story k, hk is the individual story height, and L is the bay width.
Results from the same study also demonstrated that current building codes would be appropriate for the structural design of both steel frames and masonry walls, but it was determined that there exists a lack of structural details for the connection between the two systems. 2.2 Connector Plates
Design of the connector plates for hybrid masonry systems has been addressed in its simplest form but this design has focused solely on design for in-plane and out-of-plane masonry wall reactions [Biggs, 2011]. The preliminary design procedure was developed on the sole premise that “the connector’s only purpose is to transfer in-plane shear to the wall and also to brace the frame for out-of-plane loadings” [Biggs, 2011]. Using this design approach the masonry wall is assumed to crack under high seismic loading, with the connectors resisting the load through flexure but providing little to no additional ductility to the overall system. Refer to Figure 2 for the current detailing requirements of the connector plates. There have been numerous investigations into the development of structural “fuses”, a building component that absorbs energy through yielding behavior, and even through failure, in order to prevent damage to more critical elements. In one study such structural fuses were proposed and tested as a component that “initially engages the infill walls in seismic resistance of the frame, but ultimately isolates them” [Aliaari & Memari, 2007].
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Figure 2. Current state of connector plate detailing requirements as provided by the
International Masonry Institute.
For the purposes of determining the required flexural strength of the connector
plates, the plates have been idealized as a cantilever with a maximum moment equal to the total shear to each connector plate determined by computer analysis multiplied by the distance from the bottom of the beam flange to the center of the connector plate slotted hole. Alternatively, Biggs suggests that a second option is to detail the connector plate geometry in a manner that the plates may serve as a seismic energy-dissipating fuse in the system, minimizing damage to the masonry wall. 2.3 CMU Thru-Bolt Push-Out Tests
The Building Code Requirements and Specifications for Masonry Structures (ACI 530-08) [MSJC, 2008] provides guidelines and code requirements for the design of embedded anchor bolts in masonry, but through-bolted connections are never clearly addressed. In the commentary for ACI 530-08 it states that “the design equations provided in the Code stem from research conducted on cast-in-place headed anchor bolts and bent-bar anchors (J- and L-bolts) in grout. Therefore, the application of these provisions to post-installed anchors may be in question.” ACI 530-08 provides guidelines for determining the nominal bearing strength of masonry as well as the nominal shear strength of headed and bent-bar anchor bolts. For the purposes of determining the nominal shear strength, the Code addresses the following potential failure mechanisms: nominal shear strength governed by masonry breakout, nominal shear strength governed by masonry crushing, nominal shear strength governed by anchor bolt pryout, and nominal shear strength governed by steel yielding. Although, these provisions are not directly applicable to the through-bolted connection proposed for hybrid masonry systems, they can be used as a basis for an evaluation of the
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connection performance. The applicability of the individual failure mechanisms will be addressed in more detail in the analysis portion of this report. ASTM E488 [ASTM, 2003] provides guidelines for determining the strength of anchors in concrete and masonry elements, loaded in single shear. Based on the procedure outlined in this guideline, a similar methodology has been developed for the testing of the post-installed through-bolted connections to be employed in this project.
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3 Approach
3.1 Connector Plate Tests
3.1.1 Experimental Setup
The experimental test setup for this portion of the testing was developed to simulate the loading and restraints provided by the actual overall hybrid masonry assembly, while allowing for compliance with ATC-24 guidelines. Provision was also made for quick replacement of connector plates to facilitate the testing of multiple specimens with the only damage incurred by the test specimens themselves. The lateral load was applied to the connector plate by bearing of a ¾” diameter ASTM A325 bolt in a ¾” by 1 ¾” vertical slotted hole located 0.625” from the top edge of the plate. The headed end of the bolt was welded to a 6” channel section, which provided adequate stiffness to simulate the condition of the connection to a CMU wall. The web of the channel was reinforced with a ½” cover plate at the location of the bolt for additional web stiffness. Lateral displacement of the channel section and attached bolt was provided by the extension and retraction of a hydraulic actuator connected to the 3” square tube section as shown in Figure 3. Each test specimen was individually welded to a 6” by 9” by ½” hot rolled steel base plate with four bolt holes symmetrically placed. During testing, the base plate was connected to a stiffened W12 wide flange section with four ¾” ASTM A325 bolts to provide a fixed reaction—this setup isolates the behavior of the connector/fuse from the potential flexibility of the beam above the masonry wall in the hybrid system.
Figure 3. Connector plate test setup with specimen highlighted. Lateral load is applied to the
steel tube section at the top of the image.
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The tube section was restrained from vertical translation using two load rods, with pinned connections at each end, that were bolted through the W12 base and into the laboratory’s reinforced concrete strong floor. Additional braces were attached between three locations on the testing assembly and the laboratory wall located approximately 12 feet away to provide restraint against lateral twisting, rotation, and out-of-plane translation of the tube section. A load cell was used to measure the load applied by the hydraulic actuator while an internal LVDT measured displacement. Lateral deflection of the connector plates was measured at the centerline of the bolted connection relative to the specimen base using a string potentiometer as shown in the figure. 3.1.2 Test Specimens
The connector plate specimens tested in this project were manufactured from ASTM A36 bar steel hot rolled to 4-inch widths and varying thicknesses, as identified by the actual thicknesses reported for the individual test specimens. For the purposes of this report, “link” connector plates will reference plates that have a uniform width throughout the length of the specimen, while “fuse” connector plates will refer to specimens manufactured with regions of constant or tapered width significantly less than the width of the rest of the plate. Bent plates refer to plates manufactured with a 90-degree bend near the bottom of the plate, simulating the bend as would be required for attachment to the beam flange above in the actual hybrid masonry assembly. All plates were cut to shape with a milling machine, except for specimen P4_T4_FA2-P-01, which was cut using a hand-held plasma torch. 17 connector plates were tested in this phase of the project; six link connectors and 11 fuse connectors. In order to identify the ideal fuse geometry for the purposes of this project, three unique fuse types were tested, identified as either type A, S, or T. Figure 4 shows the typical dimensions of a link connector plate as well as the three types of fuse connectors utilized. Fuse type A, for aspect ratio, was the first fuse to be designed and tested. This fuse type is characterized by a narrowed region of constant width machined into the bar with the intent to develop a plastic hinge and increase the deformation capacity of the connector plate. The narrowed region was designed to force full utilization of the non-linearity of the stress-strain curve within the fuse. Type A fuses were machined with an aspect ratio (length of narrowed region, Lf, to reduced section width, wf) of either one, two, or three and using plates of various thicknesses.
The behavior of a type S fuse, for slotted, is governed by a 2-inch (50.8 mm) wide by 4-inch (101.6 mm) long slot removed from the 4-inch (101.6 mm) wide bar. The intent of this fuse was to develop plastic hinges within both one-inch wide legs adjacent to the slot to increase the deformation capacity of the connection, while providing greater flexural stiffness than fuse type A. Type S should allow an increase in the deformation capacity and ensure full utilization of the non-linearity of the stress-strain curve within the fuse. After observing the yielding behavior of the type A fuses, a modification was made to the design of the reduced width section to account for the linearly varying moment distribution in this region.
Fuse type T, for tapered, was the result of this modification and was designed with a linearly varying reduced section width to facilitate first yielding simultaneously along the entire length of the narrowed region of the plate. For a detailed explanation of the methodology used to develop the tapered fuse section, refer to the Appendix. Fuse
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type T was also machined with an aspect ratio of either two or three, similar to type A, and as identified in the specimen nomenclature.
Figure 4. Typical link and fuse connector details (in = 25.4 mm).
Table 2. Connector plate specimen naming convention (in = 25.4mm).
In order to differentiate between the various connector plate properties utilized in
each specimen test, a specimen naming convention was developed and is shown in
PX _ TX _ FX - X - 0XPlate
Width (in.)Thickness x 1/8 inch
Fuse Type
Misc.Specimen
No.[1] [2] [3] [4] [5]4 2 1 P4_T2-014 2 A2 1 P4_T2_FA2-014 3 1 P4_T3-014 3 A2 1 P4_T3_FA2-014 4 1 P4_T4-014 4 2 P4_T4-024 4 B 1 P4_T4-B-014 4 B 2 P4_T4-B-024 4 A1 1 P4_T4_FA1-014 4 A2 1 P4_T4_FA2-014 4 A3 1 P4_T4_FA3-014 4 A2 P 1 P4_T4_FA2-P-014 4 S 1 P4_T4_FS-014 4 T2 1 P4_T4_FT2-014 4 T3 1 P4_T4_FT3-014 4 T3 B 1 P4_T4_FT3-B-014 6 A2 1 P4_T6_FA2-01
Specimen Designation
[6]
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Table 2. Column 1 gives the width in inches of the bar, or plate, from which the connector was constructed, while Column 2 gives the thickness of the same bar in eighths of an inch. Column 3 identifies the fuse type, as well as the aspect ratio when applicable, utilized for the particular specimen. Column 4 identifies any miscellaneous information about the specimen such as whether the specimen was cut with a handheld plasma torch (P) or constructed as the bent plate to be used for attachment to the beam flange (B). Column 5 simply identifies the individual specimen number for such cases when a particular plate geometry required more than one test. Finally, the complete specimen name is assembled and listed in Column 6. All specimens are referenced using the names as shown in Column 6. 3.1.3 Procedure
Of primary concern for the development of an effective steel-masonry connection is the detailing of the connector plates in order to facilitate a ductile response during seismic events. In order to investigate the response of various shaped and sized steel connector plates, a testing assembly was constructed to simulate the loading and deflection that will be applied to the plates in the full-scale tests, shown in Figure 5. This test assembly allowed for studying the behavior of a wide variety of connector plate designs under cyclic loading with minimum required labor to change test specimens.
Figure 5. Graphic representation of the proposed full-scale test structure to be constructed at
UIUC.
Each plate was installed in the test assembly as shown in Figure 3 and loaded according to the displacement-controlled sequence shown in Figure 6, at a programmed strain rate of approximately 0.2 in./in. per minute according to specified loading rates in
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ASTM A370 and E8 [ASTM, 2008]. Flexibility of the test setup resulted in an actual loading rate approximately 5% slower than the programmed value.
Using the test assembly shown, a modification of the recommended ATC-24 single-specimen testing program was developed to investigate the hysteretic responses of each connector plate. According to this adaptation of the ATC-24 procedure, each specimen was cyclically loaded using a hydraulic actuator under displacement-based control. Each cycle began with the connector plate in the vertical position (origin) and proceeded with an outward extension of the actuator corresponding to the specified lateral displacement, followed by a reversal of direction back to the origin. The second half of the cycle consisted of the retraction of the actuator a distance corresponding to the specified lateral displacement and again followed by a reversal of direction back to the origin. The first half of each cycle, the extension and return to the origin, and the second half of each cycle, the retraction and return to the origin, have been termed as a positive excursion and a negative excursion, respectively. Each displacement level (±0.010”, ±0.015”,…, ±2.75”, ±3.00”) consisted of three cycles, i, and each cycle (i = 1, 2, 3,…, n-1, n) consisted of both a positive excursion, i+, and a negative excursion, i-. The deformation history used for testing of the connector plate specimens is shown in Table 3 and Figure 6. In determining the specific displacement levels for the testing routine, a value of the yield strain for a typical connector plate was approximated and used to calculate the corresponding lateral deflection. Taking this value, ±0.015 in., the first level of displacement was chosen to provide for the completion of at least three cycles prior to yielding. After assumed first yield had occurred, lateral deflections were increased step-wise according to the deformation history. Due the flexibility of the test setup, the specified displacement levels apply only to the programmed motion of the actuator. Actual plate displacements were determined to be approximately 10% less than these values. All data were plotted using the actual measured displacements of the specimens. The following general procedure was followed for each connector plate specimen:
1) Initiate hydraulics and data collection devices.
2) Attach connector plate specimen to the testing apparatus by bolting down the specimen base plate, tightening with an impact wrench for a “rigid” connection.
Table 3. Actuator displacement summary.
Deformation Step Number
Displacement (inch = 25.4mm)
Cumulative Cycles
Completed1 0.01 32 0.015 63 0.03 94 0.06 125 0.1 156 0.25 187 0.5 218 0.75 249 1 27
10 1.25 3011 1.5 3312 1.75 3613 2 3914 2.25 4215 2.5 4516 2.75 4817 3 51
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3) Attach nut and washers at location of the slotted hole, providing a finger-tight connection that will allow for plate rotation and vertical translation with respect to the bolt.
4) Adjust lateral displacement of the plate connection until the load cell registers an initial load near zero.
5) Begin cyclic displacement-controlled loading of the specimen according to the specified deformation history. Continue cyclic testing until plate failure or significant buckling failure occurs, indicated by a significant decrease in the load on the specimen.
Figure 6. Modified ATC-24 cyclic loading routine for steel connector plates, with three cycles at
each level of lateral displacement.
3.2 CMU Thru-Bolt Push-Out Tests
3.2.1 Experimental Setup
In order to determine the lateral capacity of a through-bolted connection to the bond beam of the hybrid masonry CMU wall, the experimental test setup was developed to load a single ¾” ASTM A307 threaded rod in double shear, bearing along the cross-section of the cell of a grouted concrete masonry block. In the hybrid masonry assembly, pairs of connector plates will be placed at specified locations along the top of the wall, one connector plate on each face of the wall. Figure 7 provides an illustration of the concept applied to the testing of these connections. In order to provide symmetric loading to the threaded rod extending from each side of the test specimen, the loading device shown in Figure 8 was constructed. This loading device, composed of a top piece and two legs, was constructed from a 4” channel section. The top piece was stiffened with a 0.5” steel plate welded to its flanges, providing for a contact surface with
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the actuator platen. Each leg was connected to the top piece using two ½” diameter bolts, allowing for insertion of shims to level the device. The ¾” bolt holes were provided in each leg of the device for attachment to the test specimens. Displacement of the loading device relative to the test specimen was measured by attaching a linear potentiometer (LPOT) near the location of the bolted connection, on each face of the test specimen. A steel angle was welded to opposite flanges on each leg at the height of the bolt holes to support the LPOT pins. The base of each LPOT was connected to an aluminum block and epoxied to the face of the CMU assembly. A load cell on the actuator assembly measured the applied load on the specimens. During testing, the ends of the specimens were supported by thin steel plates and ¼” thick plywood pieces to provide a uniform bearing surface in contact with the reinforced concrete laboratory floor. The bearing plates extended 2” inward from the ends of the specimen. A threaded rod was installed into a pre-drilled ¾” hole at approximately 10.5” from each end of the wall as shown in Figure 9.
Figure 7. CMU bolt push-out test concept.
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Figure 8. CMU bolt push-out test setup.
Figure 9. Schematic drawing showing the dimensions and layout of the test assembly for
loading of the CMU wall specimens.
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3.2.2 Test Specimens
In order to determine the capacity of the through-bolted connection at the top of the CMU wall, small wall specimens were constructed to model the behavior of this connection. Since the critical bolt location of the small wall tests was determined to be at the second row of blocks from the top of the wall and at least one cell from the edge of the wall, the bolted connection capacity was tested using a specimen 3 blocks high by 2 blocks wide as shown in Figure 10. Each specimen was of running bond construction using standard 8x8x16 inch units obtained from Tileco on the island of Oahu. For a detailed test report of the CMU blocks, refer to the Appendix section of this report. The middle row of blocks was detailed as the bond beam with a centered #3 steel reinforcing bar (As = 0.11 in.2) with a 180 degree hook at each end. A single, centered vertical #5 steel reinforcing bar (As = 0.31 in.2) was provided at each end of the specimens. Identical reinforcement was provided for all the specimens.
A ¾” hole was pre-drilled into each specimen at two mirrored locations, approximately 10.5” from the edge of the specimen and at the centerline of the second row. To minimize spalling of the test specimens, drilling of the holes for the threaded rod connection was performed by pre-drilling from both sides of the specimen with masonry drill bits of increasing diameter until the required ¾” hole size was achieved. Four specimens were constructed for testing, and each specimen was drilled to allow for two tests denoted as test A or test B in order of completion, for a total of eight push-out tests. Of the four specimens, two were partially grouted (PG) and two were fully grouted (FG). In order to address all possible block arrangements, the specimens were constructed with either two half blocks and a single whole block (Configuration 1) at the bond beam level or two whole blocks (Configuration 2) as shown in Figure 10(a) and Figure 10(b), respectively. For example, specimen FG-1B denotes the 2nd test of a fully grouted specimen constructed according to block configuration 1.
Figure 10. (a) CMU block configuration "1", (b) CMU block configuration "2".
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3.2.3 Procedure
Prior to testing, each CMU specimen was drilled at two locations as previously identified, to accommodate for test A and B loading. The specimen to be tested was rotated 90 degrees and placed under the four-post hydraulic frame and rested on the steel and plywood bearing supports, extending two inches from the edge of the specimen on each side. The loading device was placed over the specimen and attached to the threaded rod using nuts and washers. In order to insure a symmetrically applied load, shims were used as leveling devices at the connections of the legs to the top channel section. Once assembled, the loading device was centered and clamped to the raised actuator platen, to hold the apparatus in place. The actuator was then loaded downward at a displacement-controlled rate of 0.002 in./sec. in 0.25-inch intervals. Loading was stopped periodically to make observations and record behavior of the test specimen. After completion of test A, the specimen was removed from the test assembly and rotated 180 degrees so that the damaged end of the specimen faced vertically upward. A new piece of threaded rod was inserted into the 2nd hole and test B was completed just as the first test. It should be noted that there was often residual damage and cracking in the specimens from test A, which often propagated well into the region of the specimen that bolt B was bearing upon, potentially affecting the results of the second test on each specimen.
3.3 Full-Scale Steel-Masonry Subassemblies
Three 8” thick CMU walls have been constructed for the purpose of testing the hybrid masonry assembly in its completed state, incorporating the results and designs from both the connector plate tests and CMU bolt push-out tests. These subassemblies will be tested under similar loading conditions to those proposed for the full scale tests at UIUC, including monotonic shear tests and cyclic shear tests. The masonry walls for these tests were previously constructed and consist of 4 rows of 8” blocks, each row 5 blocks long, in running bond assembly. Analysis of the results from both the steel connector plate tests and the masonry wall bolted connection tests will be used to determine the final plate connection details for the connection of the CMU walls to a steel frame, consisting of two wide flange columns and beam. Testing of these subassemblies will continue until an adequate system is finalized for implementation in the UIUC full-scale hybrid masonry panel tests.
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4 Results
4.1 Shear Connector Tests
4.1.1 Test Results Summary
A plot of the hysteretic behavior (load vs. displacement) for each connector plate specimen was used to visually identify critical defining points (yield point, point of maximum load, and point of maximum displacement) in the behavior of the specimen. The displacement and corresponding load for each of these critical points was manually extracted from the plots for both the positive and negative excursions. Table 4 provides a summary of the identified points for each of the 17 specimens tested in this study. The load and corresponding displacement identified at the point of specimen yielding is listed in Column 2. The maximum load and corresponding displacement are listed in Column 3. Finally, Column 4 gives the maximum displacement and corresponding load for each specimen.
Table 4. Critical Points of Hysteretic Response (in = 25.4 mm, kip = 4.45 kN, kip-in = 113 J).
Specimen Designation
[1]Displ.(in) Load (kip) Displ.(in) Load (kip) Displ.(in) Load (kip)
(+/-) (+/-) (+/-) (+/-) (+/-) (+/-)P4_T2-01 0.10/-0.15 3.0/-3.3 0.40/-0.40 4.4/-4.3 1.80/-1.90 2.3/-2.3P4_T2_FA2-01 0.12/-0.13 0.8/-0.8 1.00/-0.94 1.4/-1.4 1.95/-1.96 1.0/-1.0P4_T3-01 0.17/-0.17 5.3/-5.0 0.90/-0.84 7.3/-7.4 1.92/-1.88 4.8/-5.1P4_T3_FA2-01 0.10/-0.14 1.2/-1.3 2.65/-1.90 2.5/-2.5 2.70/-2.64 1.5/-1.8P4_T4-01 0.14/-0.14 4.9/-5.2 0.38/-0.43 7.6/-8.0 0.38/-0.43 7.6/-8.0P4_T4-02 0.17/-0.19 5.5/-6.3 0.83/-0.79 9.5/-9.3 1.12/-1.12 0.3/-0.5P4_T4-B-01 0.12/-0.15 1.7/-2.0 1.20/-0.91 3.3/-3.3 2.20/-2.17 2.9/-1.9P4_T4-B-02 0.13/-0.18 1.6/-1.8 0.67/-0.91 3.2/-3.2 1.95/-1.91 1.7/-2.3P4_T4_FA1-01 0.13/-0.12 1.7/-1.7 2.20/-2.17 3.5/-3.5 2.20/-2.17 3.5/-3.5P4_T4_FA2-01 0.12/-0.17 1.5/-2.0 2.40/-2.38 3.3/-3.5 2.40/-2.38 3.3/-3.5P4_T4_FA3-01 0.19/-0.25 2.0/1.9 2.43/-2.40 3.4/-3.4 2.43/-2.40 3.4/-3.4P4_T4_FA2-P-01 0.14/-0.17 1.5/-1.4 1.19/-1.43 2.5/-3.5 1.45/-1.43 1.8/-2.5P4_T4_FS-01 0.11/-0.07 3.2/-3.0 0.92/-0.88 5.9/-5.8 2.10/-2.14 4.5/-4.3P4_T4_FT2-01 0.17/-0.15 1.5/-1.5 3.06/-2.86 3.4/-3.2 3.06/-2.86 3.4/-3.2P4_T4_FT3-01 0.17/-0.25 1.6/-1.5 2.83/-2.90 3.2/-3.0 2.83/-2.90 3.2/-3.0P4_T4_FT3-B-01 0.22/-0.28 1.2/-1.4 2.90/-2.87 2.6/-2.6 2.90/-2.87 2.6/-2.6P4_T6_FA2-01 0.12/-0.15 1.7/-2.0 3.06/-2.86 3.3/-3.3 2.20/-2.17 2.9/-1.9
[2] [3] [4]Yield Point Maximum Load Max.Displacement
In order to determine the actual material strengths of the ASTM A36 steel bar used for the connector plate specimens, tension tests were conducted according to ASTM E8 [ASTM, 2008c]. Table 5 provides a summary of the test results for the three thicknesses of steel bar tested.
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Table 5. ASTM E8 tension test results for connector plate steel bar (inch = 25.4mm, ksi = 6.895 MPa).
Figure 11. (a) Non-Ductile Weld Rupture: P4_T4-01; (b) Rupture of Base Metal: P4_T4-02.
For all the specimens tested, five characteristic hysteretic behaviors were identified. Each of the 17 specimens displayed one of these five typical responses, with varying degrees of magnitude. The first of these, non-ductile weld rupture, is typified by the behavior of specimen P4_T4-01 in Figure 11(a). This hysteretic response represents the behavior of a specimen failure due to rupture of the weld at the base of the specimen. This weld joins the plate to the base plate that is bolted to, and considered part of, the testing frame. Figure 11(b) shows the hysteretic response of specimen P4_T4-02. This hysteretic response characterizes a specimen which first yields but then initiates rupture in the base plate and gradually separates itself from the testing frame with the base weld still intact.
Stress (ksi) Strain Stress (ksi) Strain
0.25 49.629 0.00164 79.467 0.171 31000
0.375 50.016 0.00162 79.993 0.16571 31000
0.50 53.447 0.00168 83.293 0.16771 32000
0.75 52.953 * * * *
Yield Ultimate Young's Modulus, E (ksi)
Bar Thickness (in.)
*Data not available due to testing error and/or insufficient load capacity of test frame.
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Figure 12. (a) Yielding then buckling: P4_T2-01; (b) Yielding/Fatigue rupture: P4_T4_FT3-01.
The hysteretic response of specimen P4_T2-01 shown in Figure 12(a) is representative of a specimen that begins to yield but then undergoes local buckling in the link, causing the load on the specimen to decrease at increasing displacement levels. Once buckling initiated in specimens displaying this behavior, the specimen would only begin to rupture due to fatigue from repeated cycling at large displacements. Figure 12(b) shows the hysteretic response of specimen P4_T4_FT3-01. This hysteretic response characterizes a specimen that first yields in the fuse region and then undergoes eventual fatigue rupture in this region due to increasing, or repeating, displacement cycles. The hysteretic response of specimen P4_T4_FT3-B-01, shown in Figure 13, shows the effect of a bent plate connection in combination with the yielding behavior displayed by a type T fuse. This response is characteristic of a specimen that first yields in the fuse region and continues to respond as a ductile connection until significant stress concentration at either leg of the C-shaped weld begins to initiate weld rupture.
Figure 13. Bent plate behavior, ductile response followed by C-shaped weld rupture:
P4_T4_FT3-B-01.
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4.1.2 Specimen Behavior Narratives
P4_T2-01 Plate test specimen P4_T2-01 was tested up to and through the ± 2.00”
excursions, for a total of 39 complete loading cycles. Testing of this specimen did not continue beyond the ±2.00” level as significant buckling had already occurred and the load on the specimen had diminished significantly (Figure 14). Lateral-torsional buckling was determined to be the governing limit state for this specimen, with initiation of buckling observed around the ±0.5” cycles, at a load of 4.3 kips. This specimen reached a maximum load of 4.4 kips in the positive excursion direction and 4.3 kips in the negative excursion direction.
P4_T2_FA2-01
Plate test specimen P4_T2_FA2-01 was tested up to and through the ± 2.00” excursions, for a total of 39 complete loading cycles. Testing of this specimen did not continue beyond the ±2.00” level as significant buckling had already occurred and the load on the specimen had diminished significantly. Lateral-torsional buckling was determined to be the governing limit state for this specimen, with initiation of buckling observed around the ±1.00” cycles, at a load of 1.3 kips. This specimen reached a maximum load of 1.4 kips in the positive excursion direction and -1.4 kips in the negative excursion direction.
P4_T3-01
Plate test specimen P4_T3-01 was tested up to and through two complete cycles of the ± 2.00” excursions, for a total of 38 complete loading cycles. Prior to this test, an additional restraining brace was added to the test setup in order to minimize rotation of the loading hardware due to plate buckling. Due to the significant buckling induced in this plate, the additional restraint caused significant tensile forces to develop in the bolted connection. Tension in the bolted connection contributed to failure of the bolted connection prior to completion of testing. Testing of this specimen stopped due to the bolt failure. The test setup was then modified to eliminate flexural bending of the bolt that failed. The bolt is inserted through a tight fit clearance hole—the hole was previously 1/16” oversized per typical AISC details. Lateral-torsional buckling was determined to be the governing limit state for this specimen, with initiation of buckling observed around the ±1.25” cycles, at a load of 7.1 kips. This specimen reached a maximum load of 7.25 kips in the positive excursion direction and -7.4 kips in the negative excursion direction, at a displacement of 0.9 inches. P4_T3_FA2-01
Plate test specimen P4_T3_FA2-01 was tested up to and through the ± 2.75” excursions, for a total of 48 complete loading cycles. The first signs of significant plate
Figure 14. P4_T2-01, showing
typical buckling failure.
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rupture were detected at the base of the fuse at the beginning of the ±2.75” cycles, at a load of 2.5 kips. This rupture led to complete separation of the plate across the fuse base at the end of the ±2.75” cycles. Initial signs of lateral-torsional buckling were observed at the 3rd cycle at the ±2.00” displacement level, but plate rupture was identified as the governing limit state. This specimen reached a maximum load of 2.5 kips in the positive excursion direction and -2.5 kips in the negative excursion direction. P4_T4-01 / P4_T4-02
Plate test specimen P4_T4-01 was tested up to and through the ±0.75” excursions, for a total of 24 complete cycles. Initial signs of rupture of the fillet weld joining the connector plate to the base plate were observed during the 0.50” cycles, at a load of 7.5 kips. Complete separation of the base fillet weld occurred by the end of the 0.75” cycles (Figure 15). This specimen reached a maximum load of 7.6 kips in the positive excursion direction and -8.0 kips in the negative excursion direction. In order to determine if insufficient weld thickness was the cause of premature failure, a second specimen (-02) was tested using a heavier weld. Plate test specimen P4_T4-02 was tested up to and through six complete cycles at the 1.25” displacement level. Initial signs of weld rupture were observed at the ±1.00” displacement level, at a load of 9.5 kips, followed by base plate rupture around the weld connection. This specimen reached a maximum load of 9.5 kips in the positive excursion direction and -9.25 kips in the negative excursion direction. A combination of weld failure and base plate rupture was determined to be the governing limit state for this specimen.
P4_T4-B-01 / P4_T4-B-02
Plate test specimen P4_T4-B-01 was tested up to and through the ±2.00” excursions followed by 12 cycles at ±2.25” and three cycles at ±2.50”, for a total of 54 complete cycles. Initial signs of fillet weld rupture were identified during the ±1.00” cycles in the portion of the weld perpendicular to the wide face of the plate, at a load of 3.25 kips. After the specimen reached the ±2.25” displacement level one corner of the welded bent portion of the plate had separated completely, propagating the weld failure to the long weld parallel to the wide face of the plate. At the end of the 54th cycle the load capacity dropped to 1.1 kips. The specimen was completely ruptured by manually controlled displacement with four additional cycles at undefined displacement levels until the weld had completely separated at all three sides. This specimen reached a maximum load of 3.25 kips in the positive excursion direction and -3.3 kips in the negative excursion direction. In order to determine if defects in the weld were the causes of premature failure, a second specimen (-02) was tested, with increased attention paid to weld application and thickness. Plate test specimen P4_T4-B-02 was
Figure 15. P4_T4-02 showing weld failure and base plate
rupture.
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tested up to and including one cycle at the ±2.25” displacement level. During this testing, it was noted that initial yielding was occurring on the bent leg of the plate as early as the ±0.25” cycles. Weld failure began similarly to the previous bent plate, with one corner of the weld showing complete separation by the ±2.00” displacement level. This specimen reached a maximum load of 3.15 kips in the positive excursion direction and -3.15 kips in the negative excursion direction. Weld failure was determined to be the governing limit state for these specimens. P4_T4_FA1-01
Plate test specimen P4_T4_FA1-01 was tested up to and through the ±2.25” excursions, for a total of 42 cycles. Yielding of the fuse section was first noticed through signs of flaking at the base of the fuse section during the ±0.25” displacement cycles, at a load of 2.1 kips. Initial rupture of the specimen was identified at the base of the fuse section at the ±2.00” displacement cycles, leading to complete separation at the fuse base at the end of the ±2.25” displacement cycles (Figure 16). This specimen reached a maximum load of 3.5 kips in the positive excursion direction and -3.5 kips in the negative excursion direction. Fuse base rupture was determined to be the governing limit state for this specimen. No significant buckling behavior was observed.
P4_T4_FA2-01
Plate test specimen P4_T4_FA2-01 was tested up to and through the ±2.50” excursions with an additional cycle at ±2.50”, for a total of 46 cycles. Yielding of the fuse section was first noticed through signs of flaking at the base of the fuse section at approximately the ±0.25” displacement cycles, at a load of 2.0 kips. Initial rupture of the specimen was identified at the base of the fuse section at approximately the ±2.00” displacement cycles, leading to complete separation at the fuse base at the end of four ±2.50” displacement cycles (Figure 16). This specimen reached a maximum load of 3.25 kips in the positive excursion direction and -3.5 kips in the negative excursion direction. Fuse base rupture was determined to be the governing limit state for this specimen. No significant buckling behavior was observed. P4_T4_FA3-01
Plate test specimen P4_T4_FA3-01 was tested up to and through the ±2.50” excursions with an additional cycle at ±2.50”, for a total of 46 cycles. Yielding of the fuse section was first noticed through signs of flaking at the base of the fuse section at the ±0.25” displacement cycles, at load of 2.0 kips. Initial rupture of the specimen was identified at the base of the fuse section at the ±2.25” displacement cycles, leading to
Figure 16. Specimens P4_T4_FA1-01, _FA2-01, and _FA3-01 showing
typical fuse base rupture (shown from left to right).
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complete separation at the fuse base at the end of four ±2.50” displacement cycles (Figure 16). This specimen reached a maximum load of 3.4 kips in the positive excursion direction and -3.4 kips in the negative excursion direction. Fuse base rupture was determined to be the governing limit state for this specimen. No significant buckling behavior was observed. A machining error was noted to have caused a slightly over-sized slotted hole in this specimen, leading to approximately 0.10” of slip in the connection at the top of the plate. Surface flaking was not observed over the entire fuse region but stopped at approximately 4/5th of the distance from the base to the top of the fuse. P4_T4_FA2-P-01
Plate test specimen P4_T4_FA2-P-01 was tested up to and through the ±1.50” excursions with an additional cycle at ±1.50”, for a total of 34 cycles. Yielding of the fuse section was first noticed through signs of flaking at the base of the fuse section at approximately the ±0.25” displacement cycles, at a load of 1.7 kips. Initial rupture of the specimen was identified at the base of the fuse section at the ±1.25” displacement cycles, at a load of 2.4 kips. This led to complete separation at the fuse base at the end of four ±1.50” displacement cycles (Figure 17). This specimen reached a maximum load of 2.45 kips in the positive excursion direction and -3.5 kips in the negative excursion direction. Fuse base rupture was determined to be the governing limit state for this specimen. No significant buckling behavior was observed. P4_T4_FS-01
Plate test specimen P4_T4_FS-01 was tested up to and through the ± 2.25” excursions, for a total of 42 complete loading cycles. Testing of this specimen did not continue beyond the ±2.25” level as significant buckling had already occurred and the load on the specimen had diminished significantly to 2.9 kips (Figure 18). Out-of-plane buckling was determined to be the governing limit state for this specimen, with initiation of significant buckling observed around the ±1.00” cycles, at a load of 5.9 kips. This specimen reached a maximum load of 5.85 kips in the positive excursion direction and -5.8 kips in the negative excursion direction.
Figure 17. Fuse roughness and early rupture of
P4_T4_FA2-P-01 due to cutting with a hand-held
plasma torch.
Figure 18. P4_T4-FS-01 showing significant buckling at
±2.00” displacement.
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P4_T4_FT2-01 Plate test specimen P4_T4_FT2-01 was tested
up to and through the ±3.00” excursions with an additional 13 cycles at this displacement level, for a total of 64 complete loading cycles. Uniform flaking of the tapered fuse region was observed as early as the ±0.25” displacement level, at a load of 1.76 kips, and initial signs of lateral torsional buckling behavior were identified at the ±2.25” displacement level, at a load of 2.8 kips. Although buckling behavior became significant, no decrease in load was observed so testing was continued until rupture of the plate at the 64th cycle (Figure 19). A combination of lateral-torsional buckling and plate fatigue were identified as the governing limit state for the specimen. This specimen reached a maximum load of 3.4 kips in the positive excursion direction and -3.15 kips in the negative excursion direction. P4_T4_FT3-01
Plate test specimen P4_T4_FT3-01 was tested up to and through the ±3.00” excursions with an additional 28 cycles at this displacement level, for a total of 79 complete loading cycles. Loading of this specimen was stopped after the 57th cycle. The specimen was uninstalled from the test setup and reinstalled when testing was later resumed until failure occurred at the end of the 79th cycle. Uniform flaking of the tapered fuse region was observed as early as the ±0.25” displacement level, at a load of 1.6 kips, and initial signs of lateral torsional buckling behavior were identified at the ±2.25” displacement level, at a load of 2.6 kips (Figure 20). Although buckling behavior became significant, no decrease in load was observed so testing was continued until failure. A combination of lateral-torsional buckling and eventual plate fatigue were identified as the governing limit state for the specimen. This specimen reached a maximum load of 3.2 kips in the positive excursion direction and -3.3 kips in the negative excursion direction. Figure 21 shows the final failed state of both tapered fuse specimens.
Figure 19. P4_T4-FT2-01 displaying a combination of
lateral-torsional buckling and rupture due to fatigue.
Figure 20. P4_T4-FT3-01 displaying uniform distribution of yielding along entire fuse
section.
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Figure 21. Specimens P4_T4_FT2-01 and _FT3-01, showing buckling
and tapered fuse base rupture (shown from left to right).
P4_T4_FT3-B-01
Plate test specimen P4_T4_FT3-B-01 was tested up to and through the ±3.00” excursions with an additional 15 cycles at this displacement level, for a total of 66 complete cycles. Uniform flaking of the tapered fuse region was observed as early as the ±0.25” displacement level, at a load of 1.2 kips. Yielding of the plate specimen in the region welded to the base plate was also observed at approximately the 0.50” displacement level, at a load of 1.8 kips. Initial signs of fillet weld rupture were identified during the ±2.00” cycles in the portion of the weld perpendicular to the wide face of the plate, at a load of 2.4 kips. After the specimen had completed several cycles at the ±3.00” displacement level one corner of the welded bent portion of the plate had separated completely, propagating the weld failure to the long weld parallel to the wide face of the plate. Loading was continued until the weld had completely separated at all three sides (Figure 22). This specimen reached a maximum load of 2.55 kips in the positive excursion direction and -2.6 kips in the negative excursion direction. Weld failure was determined to be the governing limit state for this specimen.
Figure 22. P4_T4_FT3-B-01 showing rupture of C-shaped
fillet weld on bent leg.
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P4_T6_FA2-01 Plate test specimen P4_T6_FA2-01 was tested
up to and through the ±2.25” excursions with an additional two cycles at this displacement level, for a total of 44 complete loading cycles. Yielding of the fuse section was first noticed through signs of flaking at the base of the fuse section at approximately the ±0.25” displacement cycles, at a load of 2.9 kips. Initial rupture of the specimen was identified at the base of the fuse section at approximately the ±2.00” displacement cycles, at a load of 5.0 kips, leading to complete separation at the fuse base at the end of five ±2.25” displacement cycles (Figure 23). This specimen reached a maximum load of 5.1 kips in the positive excursion direction and -5.1 kips in the negative excursion direction. Fuse base rupture was determined to be the governing limit state for this specimen. No significant buckling behavior was observed.
Figure 23. P4_T6_FA2-01 showing typical fuse base
rupture.
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4.2 CMU Thru-Bolt Push-Out Tests
4.2.1 Specimen Behavior and Failure Mechanisms
FG-1A/1B CMU test specimen FG-1A showed first signs of minor cracking due to bearing at
one of the specimen supports at approximately 20 kips. Cracking was then observed around the threaded rod bearing area along with mortar joint splitting on both sides of the middle half block at approximately 21 kips. Diagonal shear cracks began to form at approximately 24 kips, along with further cracking around the bolt bearing area. At 26 kips, diagonal shear cracks had propagated to within a couple inches of the supports and some horizontal cracks were observed above the location of the threaded rod. The maximum load reached by the specimen was 26.48 kips. At the end of testing, shearing type deformation was observed in the threaded rod as well as some cracking in the grout at the ends of the specimen (Figure 24). No significant cracking was observed to have propagated into the area appropriated for test FG-1B.
CMU test specimen FG-1B showed first signs of mortar joint separation along the horizontal surface of the half block immediately below the location of the threaded rod at approximately 21 kips. At this time slight vertical mortar joint separation was also observed along one side the full block immediately above the location of the threaded rod. Vertical mortar joint separation was observed along one side of the middle half block at a load of 27 kips. Testing of the specimen continued to a load of 32.13 kips, at which point one end of the threaded rod sheared off, with the other end of the threaded rod showing significant shear deformation.
Figure 24. (a) Final cracking patterns for specimen FG-1A; (b) Typical cracks around rod bearing area as seen in specimen FG-1A.
FG-2A/2B
CMU test specimen FG-2A showed first signs of minor cracking around the threaded rod bearing area along with vertical mortar joint separation along the sides of the bottom cell of the middle full block at approximately 21 kips. Diagonal shear cracks began to form at approximately 26 kips and propagated to the bottom of the specimen shortly after. Testing of this specimen continued until significant cracking had taken
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place, with the half block resting on the support beginning to rotate under the load. The maximum load reached by the specimen was 25.75 kips (Figure 25a). At the end of testing, shearing type deformation was observed in the threaded rod along with significant cracking around the rod bearing surface (Figure 25b). Some cracking was observed to have propagated into the area appropriated for test FG-2B.
CMU test specimen FG-2B showed first signs of horizontal shear cracking along with mortar joint separation along the sides of the bottom cell of the middle full block as well as the tops of the end half blocks at approximately 19 kips. The load on the specimen did not increase much beyond this point and testing continued until significant crack opening and mortar joint separation had occurred. The maximum load reached by the specimen was 19.61 kips. Some random cracking was also observed in the grout at the ends of specimen.
Figure 25. (a) Final cracking patterns for specimen FG-2A; (b) Typical threaded rod shear deformation as seen in specimen FG-2A. Load values shown on specimens were rounded to
the nearest whole number. PG-1A/1B
CMU test specimen PG-1A showed first signs of cracking around the threaded rod bearing area along with vertical mortar joint separation along the sides of the bottom cell of the middle half block at approximately 20 kips. Diagonal shear cracks began to form at approximately 24 kips, at which point the specimen reached its maximum load of 24.19 kips. Testing was continued until further displacement led to propagation of mortar joint separation up along the entire height of the full block, i.e. three cells from the bottom of the specimen. At the end of testing, slight shearing type deformation was observed in the threaded rod as well as splitting cracks through the ungrouted cells at the ends of the specimen. Although no cracking was observed to have propagated into the area appropriated for test PG-1B, mortar joint separation had propagated into that area.
CMU test specimen PG-1B showed first signs of mortar joint separation along the sides of the middle half block at approximately 22 kips. Diagonal shear cracks began to form at approximately 25 kips. Loading continued until the specimen reached its maximum load of 25.26 kips. Continued deformation of the specimen beyond its
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maximum load led to significant vertical mortar joint separation, propagating almost through the entire height of the specimen (Figure 26). At the end of testing, significant opening of shear cracks had occurred and shearing type deformation of the threaded rod was also observed.
Figure 26. Specimen PG-1B showing mortar joint separation along the entire height of the right
side of the middle blocks, typical of partially grouted specimens.
PG-2A/2B
CMU test specimen PG-2A showed first signs of vertical mortar joint separation along the sides of the middle full block at approximately 20 kips. Cracks around the threaded rod bearing area and some horizontal mortar joint separation along the tops of the end half blocks were observed at approximately 22.5 kips. Diagonal shear cracks began to form at approximately 24 kips, with the specimen maximum load of 24.48 kips occurring shortly after. Testing was stopped after mortar joint separation began propagating through almost the entire height of the specimen. At the end of testing, significant opening of shear cracks had occurred and shearing type deformation of the threaded rod was also observed (Figure 27). Although no cracking was observed to have propagated into the area appropriated for test PG-2B, mortar joint separation had propagated into that area.
CMU test specimen PG-2B showed first signs of mortar joint separation along both surfaces of one of the end half blocks at approximately 21 kips, followed by cracking around the threaded rod bearing area and vertical mortar joint separation of the other end half block. Diagonal shear cracks began to form at approximately 27 kips, with the specimen reaching its maximum load of 27.30 kips shortly after. Continued deformation of the specimen beyond its maximum load led to significant vertical mortar
Page 30
joint separation, propagating almost through the entire height of the specimen. At the end of testing, significant opening of shear cracks had occurred and shearing type deformation of the threaded rod was observed.
Figure 27. Final cracking patterns of specimen PG-2A, with vertical mortar joint separation
visible above the location of the threaded rod.
4.2.2 Test Results
A summary of the maximum loads reached by all the test specimens, as well as the deflections at which these loads occurred, is provided in Table 6. In addition to the CMU bolt push-out test specimens, three CMU prism specimens were constructed and tested to determine the compressive strength of the masonry, fmt, according to ASTM C1314 [ASTM, 2008b]. The results from these specimen tests are listed in Table 7.
Page 31
Table 6. CMU bolt push-out maximum load values and corresponding downward deflection (in = 25.4 mm, kip = 4.45 kN).
Specimen Test Max Load (kips) Deflection (in.)A 26.48 0.710
B 32.13 0.600
A 25.75 0.792
B 19.61 0.578
A 24.19 0.442
B 25.26 0.458
A 24.48 0.442
B 27.30 0.466
FG1
FG2
PG1
PG2
Table 7. Summary of results from CMU prism tests.
The results for the individual test specimens are grouped as “A” tests and “B” tests and plotted in Figure 28 and Figure 29, respectively, with the displacement values on the x-axis and the corresponding load value on the y-axis. These plots represent the data collected for each specimen and show the difference in behavior between the fully grouted and partially grouted specimens as well as the effect of damage incurred during the “A” tests on the behavior of the specimens during the “B” tests.
SpecimenMaximum
Compressive LoadCross-Sectional
AreaObserved Mode of
FailureCompressive Strength, fmt
(lbs = 4.45 N) (in2 = 645.16 mm2) (psi = 0.00689 N/mm2)
[1] [2] [3] [4] [5]
P1 137500 58.14Semi-conical break, face shell separation, no grout cracking
2365
P2 138000 58.14 Face shell separation 2374
P3 155500 58.14 Face shell separation, some grout cracking
2675
Page 32
Figure 28. Load vs. Displacement for all CMU specimen "A" tests.
Figure 29. Load vs. Displacement for all CMU specimen "B" tests.
Page 33
5 Analysis
5.1 Comparison of Connector Plates (Link and Fuse Types A, S, and T)
Of all the connector plates tested, the full-width specimens without the fuse detail provided the highest maximum load values, as the greater specimen width provided a larger cross-section for higher flexural loads. Either plate buckling or failure of the connection of the specimen to the test assembly ultimately governed behavior of these specimens. Such a plate design would be preferable only for situations requiring a non-ductile connection with higher lateral load requirements.
Fuse type A provided a much more ductile response than the full width plates, although a significant decrease in the load capacity was observed due to the reduced cross-section. Failure of these specimens was also much more predictable and they maintained their maximum load for multiple cycles before rupture occurred due to high strain levels at the base of the fuse region.
Fuse type S performed equally as well as the type A specimens up until larger displacement levels at which point buckling behavior began to cause a reduction in the maximum load achieved in subsequent excursions.
Fuse type T provided the most ductile behavior of all the specimens. The tapered fuse width allowed for a nearly uniform distribution of initial yielding, identified through the uniform flaking of mill scale along the fuse length.
Implementation of the bent plate configuration introduced another design consideration into the detailing of specimens. Specimens bearing this connection type typically failed by failure of the C-shaped weld, without reaching the energy dissipation potential of a similar specimen without the bent plate configuration. Further testing is necessary in developing this connection detail to compensate for the introduction of earlier weld rupture. Table 8 summarizes the cumulative displacement and total energy
Table 8. Specimen Cumulative Displacement and Total Energy Dissipation Values (in = 25.4 mm, kip-in = 113 J).
Specimen CumulativeDisplacement (in)
Total Energy Dissipated (kip-in)
[1] [2] [3]
P4_T4-01 13.364 16.207
P4_T2_FA2-01 107.561 67.413
P4_T4_FA2-P-01 73.718 74.276
P4_T4-02 54.077 90.023
P4_T4_B-02 113.705 119.025
P4_T2-01 101.492 119.666
P4_T4_B-01 271.477 205.452
P4_T4_FA1-01 131.433 260.585
P4_T3_FA2-01 193.302 270.003
P4_T3-01 94.909 278.346
P4_T4_FA3-01 165.132 316.197
P4-T4-FS-01 126.714 325.956
P4_T4_FA2-01 169.192 328.713
P4_T6_FA2-01 143.132 378.051
P4_T4_FT3-B-01 394.675 487.936
P4_T4_FT2-01 383.307 698.707
P4_T4_FT3-01 544.697 978.726
Page 34
dissipated for each specimen, in columns 2 and 3, respectively. The rows in the table were sorted numerically according to the values in column 3 from least to greatest in order to provide a rank of the energy dissipating characteristics of the specimens. The value of energy dissipation shown for each specimen was determined by integrating the area under the hysteretic loop for each excursion and then summing these values for the entire testing routine. Similarly, the cumulative displacement were obtained by summing the maximum measured displacement value reached for each excursion completed by the specimen up until the point of failure or significant buckling that validated ending the test.
5.2 ATC-24 Connector Plate Analysis
In order to facilitate performance evaluation of the connector plate specimens tested, an analysis of the hysteretic test results was conducted to generate all of the parameters recommended in ATC-24 [ATC, 1992]. The intention of the ATC-24 recommendations is to provide all the data necessary for a deduction of a complete set of performance parameters, addressing measures of specimen strength characteristics, stiffness characteristics, deformation capacity, and energy dissipation for each excursion of the specimen loading history. With these results, future analytic models can be developed for further exploration and/or verification of earlier predictions. Table 9 through Error! Reference source not found. show the ATC-24 recommended results for three representative connector plate specimens. All of the experimentally determined values listed in the tables have been normalized with respect to analytically determined values of yield force, Qy, and yield displacement, δy, based on an assumed material yield strength, Fy, of 42 ksi. Column 1 lists the entry number for each row of data, starting with 1 at the first excursion and increasing with each subsequent excursion. Column 2 lists the absolute value of the programmed actuator displacement level corresponding to each excursion. Column 3 lists the excursion number from which each row of data was extracted, followed by a positive sign (+) or a negative sign (-) noting either a compressive or a tensile force in the actuator load cell, respectively. Although previously an excursion was identified by displacement direction, ATC-24 documentation requires an excursion to be identified by the sign of the load on the specimen since residual deformation can make displacement levels deviate from the origin. The maximum deflection, δ, reached by a specimen during a given excursion is listed in Column 4. The deflection at the start of each excursion, δ0, is listed in Column 5. The difference between this value and the deflection at the start of the subsequent excursion (i.e. the point at which the load on the specimen changed signs) represents the plastic deformation range, Δδpm, for each excursion and is listed in Column 6. It should be noted that the plastic deflection range for the final entry of each specimen test has been omitted from the tables as this value would often prove to be a misrepresentation of the physical test in specimens that either failed prior to completion of the excursion or when loading was stopped before completion of the excursion.
The load, Q, corresponding to the maximum deflection in each excursion is given The load, Q, corresponding to the maximum deflection in each excursion is given in Column 7, while Column 8 lists the maximum overall load, Qmax, attained by the specimen during each excursion. The initial loading stiffness, K0, and initial unloading stiffness, K, are given in Columns 9 and 10, respectively. These values correspond to the slope (change in load vs. change in displacement) of the first 20 (loading) and last
Page 35
20 (unloading) data values in each excursion. Finally, the area under the hysteretic loop of each excursion was used to quantity the specimen energy dissipation, as listed in Column 11.
The ATC-24 based hysteretic analysis of specimens P4_T4-01, P4_T4_FA3-01, and P4_T4_FT3-01 are shown in Table 9, Table 10, and Error! Reference source not found. respectively, providing a representation of three of the general failure modes observed: base connection failure, fuse base rupture, and fuse buckling/fatigue rupture, respectively. All other specimen tables can be found in the Appendix.
Page 36
Sp
ecim
enP
4_T
4-0
1F
y =
42.0
0ks
i(a
ssu
me
d)Q
y =
5.6
0ki
ps!
y =
0.0
593
in.
En
try
Dis
pla
cem
ent
Le
vel
Pea
k D
efle
ctio
n
De
flect
ion
at
Sta
rt o
f E
xcu
rsio
n
Pla
stic
D
efle
ctio
n R
ang
e
Loa
d a
t P
eak
De
flect
ion
Ma
x. L
oad
in
Exc
urs
ion
Initi
al L
oadi
ng
Stif
fnes
sIn
itia
l Unl
oadi
ng
Stif
fne
ssH
yste
resi
s A
rea
(in.)
(+
/-)
i+i-
!/!
y!
0/!
y"
!pm
/!y
Q/Q
yQ
ma
x/Qy
K0/(
Qy/
! y)
K/(
Qy/
! y)
A/(
Qy/
! y)
[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
[11]
10
.01
1+
0.11
20.
112
-0.0
28-0
.004
-0.0
09-0
.065
-0.0
65
0.00
0007
21
-0.
182
0.09
80
.03
50
.03
10
.051
0.2
890.
115
0.0
0001
23
2+
0.11
20.
112
-0.0
21-0
.002
-0.0
08-0
.047
-0.0
47
0.00
0013
42
-0.
175
0.10
50
.02
10
.04
70
.051
0.2
930.
202
0.0
0001
85
3+
0.12
60.
126
-0.0
28-0
.002
-0.0
07-0
.040
-0.0
40
0.00
0018
63
-0.
203
0.11
20.
007
0.0
150.
051
0.35
6-0
.00
60.
0000
227
0.0
24
+0.
091
0.09
1-0
.091
-0.0
02-0
.011
-0.0
41-0
.04
10.
0000
248
4-
0.19
60.
007
0.1
19
0.0
66
0.0
730.
366
0.25
00
.000
035
95
+0.
098
0.09
8-0
.105
-0.0
02-0
.011
-0.0
14-0
.01
40.
0000
371
05
-0.
189
-0.0
07
0.1
050.
072
0.0
720.
359
0.30
40.
0000
4811
6+
0.09
80.
098
-0.0
98-0
.002
-0.0
11-0
.001
-0.0
01
0.00
0050
12
6-
0.20
30.
000
0.14
70
.015
0.07
20.
343
-0.0
02
0.00
0064
130.
037
+0.
126
0.12
6-0
.168
-0.0
01-0
.053
0.25
60.
256
0.00
0075
14
7-
0.34
9-0
.04
20
.217
0.14
10
.141
0.33
10.
728
0.00
0106
15
8+
0.15
40.
154
-0.1
89
0.0
00
-0.0
530.
261
0.26
10
.000
114
16
8-
0.36
3-0
.02
10
.182
0.12
40
.142
0.35
10.
361
0.00
0144
17
9+
0.14
70.
147
-0.1
75
0.0
00
-0.0
520.
247
0.24
70
.000
153
18
9-
0.35
6-0
.02
80
.203
0.13
80
.142
0.34
30.
660
0.00
0185
190.
0610
+-0
.29
40.
147
-0.2
10
-0.1
44-0
.144
0.3
020.
617
0.0
0022
82
01
0-
0.62
9-0
.04
90
.224
0.27
30
.285
0.31
30.
605
0.00
0308
21
11+
-0.2
86
0.14
0-0
.21
0-0
.129
-0.1
440.
288
0.55
50
.000
346
22
11-
0.62
2-0
.04
90
.210
0.27
90
.285
0.30
80.
742
0.00
0421
23
12
+-0
.28
60.
119
-0.1
96
-0.1
44-0
.144
0.3
120.
713
0.0
0045
72
41
2-
0.62
9-0
.06
30
.238
0.26
60
.285
0.30
20.
580
0.00
0530
250.
1013
+-0
.59
40.
126
-0.2
38
-0.2
63-0
.280
0.3
010.
562
0.0
0062
92
61
3-
1.03
4-0
.07
70
.266
0.45
80
.463
0.24
40.
582
0.00
0816
27
14
+-0
.59
40.
161
-0.2
59
-0.2
79-0
.285
0.2
850.
721
0.0
0088
92
81
4-
1.03
4-0
.08
40
.259
0.45
20
.462
0.19
90.
569
0.00
1031
29
15
+-0
.59
40.
154
-0.2
38
-0.2
55-0
.285
0.2
470.
520
0.0
0111
53
01
5-
1.04
1-0
.06
30
.259
0.46
10
.462
0.23
30.
573
0.00
1273
Exc
urs
ion
[3]
Table 9. ATC-24 based table for specimen P4_T4-01.
Page 37
S
pec
imen
P4_
T4
-01
Fy =
42.0
0ks
i(a
ssu
me
d)Q
y =
5.6
0ki
ps!
y =
0.0
593
in.
En
try
Dis
pla
cem
ent
Le
vel
Pea
k D
efle
ctio
n
De
flect
ion
at
Sta
rt o
f E
xcu
rsio
n
Pla
stic
D
efle
ctio
n R
ang
e
Loa
d a
t P
eak
De
flect
ion
Ma
x. L
oad
in
Exc
urs
ion
Initi
al L
oadi
ng
Stif
fnes
sIn
itia
l Unl
oadi
ng
Stif
fne
ssH
yste
resi
s A
rea
(in.)
(+
/-)
i+i-
!/!
y!
0/!
y"
!pm
/!y
Q/Q
yQ
ma
x/Qy
K0/(
Qy/
! y)
K/(
Qy/
! y)
A/(
Qy/
! y)
[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
[11]
Exc
urs
ion
[3]
310.
2516
+-1
.97
80.
168
-0.4
33
-0.7
62-0
.773
0.2
351.
018
0.0
0205
23
21
6-
2.65
6-0
.25
90
.720
0.89
90
.918
0.15
40.
437
0.00
3704
33
17
+-2
.06
90.
440
-0.5
80
-0.8
35-0
.835
0.11
50.
512
0.0
0455
73
41
7-
2.67
0-0
.12
60
.545
0.91
80
.918
0.10
40.
496
0.00
5508
35
18
+-2
.07
60.
426
-0.5
59
-0.8
34-0
.838
0.11
80.
521
0.0
0624
43
61
8-
2.67
7-0
.11
90
.524
0.91
80
.918
0.08
80.
488
0.00
7123
370.
5019
+-4
.45
90.
419
-1.5
80
-1.2
00-1
.233
0.0
900.
404
0.0
1274
23
81
9-
5.91
3-1
.13
93
.348
1.34
81
.348
0.06
10.
632
0.02
5817
39
20
+-4
.16
52.
195
-2.9
07
-1.2
73-1
.283
0.0
490.
444
0.0
3468
54
02
0-
5.94
1-0
.69
92
.817
1.34
41
.363
0.03
50.
492
0.04
4478
41
21
+-4
.20
02.
125
-2.8
17
-1.2
90-1
.300
0.0
430.
445
0.0
5294
64
22
1-
6.01
1-0
.69
92
.865
1.32
91
.341
0.03
00.
478
0.06
2655
430.
7522
+-7
.24
02.
167
-5.3
26
-1.3
92-1
.414
0.0
370.
425
0.0
8403
14
42
2-
11.5
25
-3.1
45
10.8
540.
786
1.3
720.
054
0.40
50.
1290
214
52
3+
-11.
224
7.67
4-1
8.70
2-0
.013
-1.1
810.
116
0.00
10
.170
018
46
23
-13
.705
-11.
007
14.6
700
.007
0.03
10.
022
0.01
10
.171
063
47
24
+-1
1.26
63.
620
-14.
830
-0.0
03-0
.006
0.00
50.
000
0.1
7128
44
82
4-
13.7
40-1
1.17
524
.629
0.0
030.
006
0.01
10.
000
0.1
7161
149
0.75
25+
13.4
26
13.4
26
---
0.0
00
-0.0
020.
001
0.00
1--
-
Page 38
Sp
ec
imen
P4
_T4_
FA3-
01
Fy =
42.
00
ksi
(ass
um
ed
)Q
y =
1.4
0ki
ps
!y =
0.08
26in
.
En
try
Dis
pla
cem
en
t L
eve
lP
eak
D
efle
ctio
n
Def
lect
ion
at
Sta
rt o
f E
xcu
rsio
n
Pla
stic
D
efle
ctio
n R
ang
e
Load
at P
eak
De
flect
ion
Max
. L
oad
in
Exc
urs
ion
Initi
al L
oadi
ng
Stif
fnes
sIn
itial
Unl
oadi
ng
Stif
fnes
sH
yste
resi
s A
rea
(in.)
(+
/-)
i+i-
!/! y
! 0/!
y"
!pm
/!y
Q/Q
yQ
ma
x/Qy
K0/(
Qy/
! y)
K/(
Qy/!
y)A
/(Q
y/! y
)
[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
[11]
10
.01
1+
-0.0
75-0
.03
0-0
.02
0-0
.032
-0.0
34
0.3
68
0.3
840
.00
0002
21
--0
.040
-0.0
350.
005
0.0
190.
086
0.9
280.
970
0.0
0001
73
2+
-0.0
90-0
.05
0-0
.01
5-0
.027
-0.0
35
0.3
13
0.2
300
.00
0019
42
--0
.100
-0.0
50
0.10
00
.077
0.0
81-0
.29
7-0
.29
00
.00
0040
53
+-0
.136
-0.0
05
-0.0
60
-0.0
39-0
.03
90
.26
20
.254
0.0
000
526
3-
-0.0
60-0
.03
50.
025
0.0
140.
081
0.8
10
-0.0
34
0.0
000
557
0.0
24
+-0
.151
-0.0
45
-0.0
80
-0.0
37-0
.03
90
.19
90
.152
0.0
000
728
4-
-0.1
25-0
.125
0.09
00
.010
0.11
70
.441
0.44
10
.000
120
95
+-0
.156
-0.0
50
-0.0
95
-0.0
33-0
.03
60
.17
90
.129
0.0
001
401
05
--0
.130
-0.1
300.
105
0.0
040.
117
0.4
500.
450
0.0
0017
911
6+
-0.2
61-0
.03
5-0
.22
6-0
.014
-0.0
35
0.0
96
0.0
340
.00
0219
12
6-
-0.2
61-0
.261
0.23
60
.007
0.11
50
.224
0.22
40
.000
230
13
0.0
37
+-0
.346
-0.0
45
-0.2
91
-0.0
31-0
.03
50
.03
40
.056
0.0
002
911
47
--0
.336
-0.3
360.
291
0.0
050.
224
0.0
940.
094
0.0
0040
11
58
+-0
.351
-0.0
55
-0.2
86
-0.0
29-0
.03
20
.02
50
.053
0.0
004
571
68
--0
.341
-0.3
410.
306
0.0
050.
222
0.0
940.
094
0.0
0055
31
79
+-0
.346
-0.0
50
-0.2
86
-0.0
23-0
.03
10
.01
80
.040
0.0
006
081
89
--0
.316
-0.3
160.
281
0.0
070.
222
0.0
930.
093
0.0
0068
01
90
.06
10
+-0
.717
-0.0
50
-0.6
62
-0.0
16-0
.03
20
.02
90
.015
0.0
008
052
01
0-
-0.7
02-0
.702
0.67
20
.003
0.44
60
.064
0.06
40
.001
069
21
11+
-0.7
17-0
.04
0-0
.67
2-0
.022
-0.0
29
0.0
21
0.0
190
.00
1183
22
11-
-0.7
02-0
.702
0.68
20
.007
0.44
60
.029
0.02
90
.001
337
23
12
+-0
.712
-0.0
30
-0.6
82
-0.0
10-0
.02
40
.01
30
.009
0.0
014
342
41
2-
-0.6
97-0
.697
0.68
20
.006
0.44
40
.028
0.02
80
.001
594
25
0.1
01
3+
-1.1
79-0
.02
5-1
.05
3-0
.089
-0.0
95
0.0
21
0.0
410
.00
1754
26
13
--1
.059
-1.0
591.
043
0.0
070.
734
0.0
080.
008
0.0
0215
32
71
4+
-1.0
89-0
.025
-1.0
08-0
.018
-0.1
040.
006
-0.0
900.
0022
712
81
4-
-1.0
19-1
.019
1.00
80
.007
0.73
20
.004
0.00
40
.002
588
29
15
+-1
.164
-0.0
25
-1.0
18
-0.0
90-0
.10
40
.00
60
.042
0.0
027
083
01
5-
-1.0
29-1
.029
1.01
30
.005
0.73
10
.008
0.00
80
.002
972
Exc
urs
ion
[3]
Table 10. ATC-24 based table for specimen P4_T4_FA3-01.
Page 39
Sp
eci
me
nP
4_T
4_
FA
3-0
1F
y =
42.0
0ks
i(a
ssum
ed)
Qy =
1.4
0ki
ps
!y =
0.0
826
in.
Ent
ryD
isp
lace
me
nt
Leve
lP
eak
D
efle
ctio
n
Def
lect
ion
at
Sta
rt o
f E
xcu
rsio
n
Pla
stic
D
efle
ctio
n
Ran
ge
Loa
d a
t Pea
k D
efle
ctio
nM
ax. L
oad
in
Exc
urs
ion
Initi
al L
oadi
ng
Stif
fnes
sIn
itia
l Unl
oad
ing
Stif
fne
ssH
yste
resi
s A
rea
(in.)
(+
/-)
i+i-
!/! y
! 0/!
y"
! pm/!
yQ
/Qy
Qm
ax/Q
yK
0/(
Qy/
! y)
K/(
Qy/
! y)
A/(
Qy/
! y)
[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
[11
]
Exc
urs
ion
[3]
310
.25
16+
-2.5
94-0
.02
0-1
.114
-1.0
58-1
.06
50.
003
0.84
40
.00
36
9132
16-
2.49
8-1
.104
1.73
61.
432
1.43
20.
012
1.02
70.
0107
6933
17+
-2.5
740
.607
-1.5
85-1
.158
-1.1
58
0.0
030.
786
0.0
13
872
3417
-2.
523
-0.9
581.
585
1.4
221.
433
0.01
00.
825
0.01
8281
3518
+-2
.559
0.5
97-1
.565
-1.1
71-1
.17
10.
001
0.77
90
.02
12
0836
18-
2.51
8-0
.948
0.94
31.
436
1.43
60.
008
0.89
20.
0251
5337
0.5
019
+-5
.448
0.0
05-3
.396
-1.4
83-1
.48
30.
007
1.04
80
.05
21
2838
19-
5.45
8-3
.371
6.59
21.
597
1.59
70.
011
0.98
30.
1041
3339
20+
-5.4
583
.206
-6.3
26-1
.596
-1.5
96
0.0
030.
776
0.1
51
509
4020
-5.
458
-3.1
006.
246
1.6
071.
647
0.00
10.
815
0.19
9872
4121
+-5
.408
3.1
20-6
.166
-1.6
34-1
.63
40.
004
0.85
90
.24
59
7142
21-
5.45
3-3
.025
6.09
01.
654
1.67
00.
001
0.86
60.
2931
7243
0.7
522
+-8
.358
3.0
40-8
.895
-1.7
43-1
.74
30.
003
0.87
70
.37
18
8044
22-
8.43
3-5
.845
11.7
141.
750
1.7
980.
006
0.82
80.
4819
8245
23+
-8.3
535
.854
-11.
569
-1.8
06-1
.80
60.
006
0.86
80
.58
83
8346
23-
8.42
3-5
.684
11.4
681.
815
1.8
34-0
.002
0.96
20
.697
320
4724
+-8
.358
5.7
59-1
1.46
8-1
.811
-1.8
29
0.0
030.
783
0.8
03
707
4824
-8.
423
-5.6
7411
.428
1.8
391
.846
-0.0
020.
922
0.9
1235
649
1.0
025
+-1
1.29
35
.739
-14.
253
-1.9
04-1
.90
50.
003
0.85
81
.05
55
7850
25-
11.4
13
-8.4
88
17.
062
1.9
171
.94
20.
008
0.90
71
.23
45
7451
26+
-11.
293
8.5
48-1
6.97
7-1
.938
-1.9
49
0.0
030.
846
1.4
10
592
5226
-11
.40
8-8
.41
31
6.9
271.
891
1.9
68
0.0
090.
818
1.5
89
471
5327
+-1
1.25
38
.488
-16.
841
-1.9
64-1
.96
90.
005
0.89
21
.76
55
7754
27-
11.4
18
-8.3
43
16.
851
1.9
301
.98
30.
000
0.85
61
.94
48
2955
1.2
528
+-1
4.23
38
.498
-19.
676
-2.0
30-2
.03
10.
005
0.87
32
.16
09
1956
28-
14.
393
-11.
167
22.
505
2.0
512
.06
70.
011
0.96
32
.41
69
7257
29+
-14.
213
11.3
23-2
2.43
5-2
.034
-2.0
70
0.0
060.
772
2.6
69
340
5829
-1
4.3
83-1
1.0
922
2.4
052.
061
2.0
90
0.0
060.
914
2.9
26
071
5930
+-1
4.22
311
.288
-22.
355
-2.0
38-2
.08
70.
011
0.75
73
.17
97
0160
30-
14.
388
-11.
067
22.
335
2.1
002
.10
30.
005
0.94
33
.43
68
10
Page 40
Sp
ecim
enP
4_T
4_F
A3
-01
Fy =
42.0
0ks
i(a
ssu
me
d)Q
y =
1.4
0ki
ps!
y =
0.0
826
in.
En
try
Dis
pla
cem
ent
Le
vel
Pea
k D
efle
ctio
n
De
flect
ion
at
Sta
rt o
f E
xcu
rsio
n
Pla
stic
D
efle
ctio
n R
ang
e
Loa
d a
t P
eak
De
flect
ion
Ma
x. L
oad
in
Exc
urs
ion
Initi
al L
oadi
ng
Stif
fnes
sIn
itia
l Unl
oadi
ng
Stif
fne
ssH
yste
resi
s A
rea
(in.)
(+
/-)
i+i-
!/!
y!
0/!
y"
!pm
/!y
Q/Q
yQ
ma
x/Qy
K0/(
Qy/
! y)
K/(
Qy/
! y)
A/(
Qy/
! y)
[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
[11]
Exc
urs
ion
[3]
611.
5031
+-1
7.13
711
.25
3-2
5.15
9-2
.133
-2.1
450.
007
1.00
33
.732
624
62
31
-17
.383
-13
.886
28.0
341.
983
2.1
770.
009
0.45
14.
0719
296
33
2+
-17.
147
14.1
22
-27.
943
-2.1
77-2
.177
0.02
30.
900
4.4
0730
36
43
2-
17.4
08-1
3.8
1127
.903
2.1
692.
196
0.00
00.
883
4.7
4791
26
53
3+
-17.
157
14.0
62
-27.
863
-2.1
93-2
.195
0.01
50.
876
5.0
8446
96
63
3-
17.4
13-1
3.7
7627
.878
2.17
22
.208
0.00
70.
855
5.42
5560
671.
7534
+-2
0.07
214
.11
2-3
0.76
8-2
.248
-2.2
480.
039
0.94
15
.806
624
68
34
-20
.433
-16
.651
33.6
122.
271
2.2
740.
006
0.93
76.
2350
706
93
5+
-20.
067
16.9
57
-33.
547
-2.2
70-2
.278
0.08
40.
880
6.6
5727
67
03
5-
20.4
23-1
6.5
7033
.532
2.26
32
.295
0.01
30.
881
7.08
7159
71
36
+-2
0.07
716
.94
2-3
3.52
2-2
.273
-2.2
870.
082
0.91
07
.510
711
72
36
-20
.423
-16
.555
33.4
772.
305
2.3
050.
011
1.11
17.
9415
8173
2.00
37+
-23.
037
16.9
01
-36.
331
-2.3
31-2
.344
0.11
80.
917
8.4
1225
67
43
7-
23.3
43-1
9.4
1539
.221
2.36
42
.365
0.01
61.
126
8.93
3146
75
38
+-2
3.05
719
.79
1-3
9.17
1-2
.351
-2.3
530.
193
0.90
09
.446
622
76
38
-23
.348
-19
.370
39.1
512.
346
2.3
770.
030
0.88
99.
9675
767
73
9+
-23.
062
19.7
61
-39.
106
-2.3
39-2
.354
0.24
70.
870
10.
4815
537
83
9-
23.3
58
-19
.33
039
.096
2.3
78
2.3
780.
042
1.10
211
.002
794
792.
2540
+-2
6.0
27
19.7
56
-42
.021
-2.3
73-2
.402
0.2
880.
802
11.5
6482
38
04
0-
26.2
93
-22
.24
444
.915
2.3
89
2.4
270.
199
0.91
41
2.1
7867
08
14
1+
-26.
022
22.6
66
-44.
880
-2.4
19-2
.419
0.42
90.
905
12.
7851
558
24
1-
26.3
03
-22
.20
444
.855
2.3
50
2.4
340.
406
0.63
21
3.3
9962
98
34
2+
-25.
982
22.6
36
-44.
835
-2.3
82-2
.425
0.42
20.
811
14.
0057
158
44
2-
26.3
28
-22
.18
444
.820
2.4
25
2.4
300.
419
1.05
21
4.6
1986
385
2.50
43+
-28.
957
22.6
11-4
7.68
9-2
.464
-2.4
700.
432
0.87
81
5.27
3903
86
43
-29
.23
8-2
5.0
74
50.5
442
.39
22
.476
0.4
190.
864
15.
981
140
87
44
+-2
8.94
725
.45
0-5
0.49
9-2
.465
-2.4
780.
417
0.87
91
6.67
7462
88
44
-29
.26
8-2
5.0
29
50.5
792
.39
52
.401
0.4
451.
084
17.
378
948
89
45
+-2
8.98
225
.53
0-5
0.58
9-2
.400
-2.4
610.
411
0.75
81
8.06
6176
90
45
-29
.79
0-2
5.0
39
52.6
510
.33
21
.860
0.4
190.
290
18.
468
365
912.
5046
+27
.597
27.5
92-4
4.86
5-0
.001
-1.3
470.
090
0.09
01
8.65
9227
92
46
--2
9.78
5-1
7.2
83--
-0
.025
0.04
4-0
.015
0.01
0--
-
Page 41
S
pe
cim
en
P4
_T4_
FT
3-0
1F
y =
42.0
0ks
i(a
ssu
me
d)Q
y =
1.02
kip
s!
y =
0.0
871
in.
Ent
ryD
isp
lace
me
nt
Leve
lP
eak
D
efle
ctio
n
Def
lect
ion
at
Sta
rt o
f E
xcu
rsio
n
Pla
stic
D
efle
ctio
n
Ra
nge
Loa
d at
Pe
ak
Def
lect
ion
Ma
x. L
oad
in
Exc
urs
ion
Initi
al L
oad
ing
Stif
fnes
sIn
itial
Unl
oad
ing
Stif
fnes
sH
yste
resi
s A
rea
(in.)
(+
/-)
i+i-
!/! y
! 0/!
y"
!p
m/!
yQ
/Qy
Qm
ax/Q
yK
0/(
Qy/!
y)K
/(Q
y/!y)
A/(
Qy/
! y)
[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
[11]
10.
2516
+
216
-2.
023
0.0
29-0
.19
01.
596
1.6
25
0.7
020.
725
0.00
3857
317
+-2
.889
-0.1
90
-0.5
71
-1.5
26
-1.5
300
.419
0.7
620.
0077
034
17-
2.03
7-0
.738
0.56
21.
622
1.62
20
.499
0.76
80.
0120
325
18+
-2.8
84-0
.18
6-0
.56
6-1
.53
2-1
.532
0.4
120.
781
0.01
5654
618
-2.
037
-0.7
330.
547
1.62
31.
623
0.4
940.
808
0.01
9685
70.
5019
+-5
.654
-0.2
00
-2.9
08
-1.7
88
-1.7
990
.372
0.8
070.
0555
328
19-
4.85
5-3
.098
5.40
71.
773
1.77
50
.394
0.72
50
.115
048
920
+-5
.635
2.2
89-5
.37
3-1
.76
6-1
.780
0.3
920.
771
0.16
9968
1020
-4.
888
-3.0
705.
349
1.75
71.
787
0.3
830.
680
0.22
5324
1121
+-5
.640
2.2
56-5
.27
3-1
.78
9-1
.789
0.3
830.
765
0.27
8852
1221
-4.
859
-3.0
085.
245
1.79
71.
797
0.4
380.
907
0.33
2947
130.
7522
+-8
.457
2.2
32-7
.96
2-1
.85
5-1
.863
0.3
780.
805
0.42
3885
1422
-7.
686
-5.7
0610
.656
1.88
31
.887
0.34
20.
873
0.54
5740
1523
+-8
.410
4.91
6-1
0.56
1-1
.91
0-1
.911
0.3
310.
811
0.66
4781
1623
-7.
672
-5.6
3010
.452
1.92
11
.931
0.31
30.
855
0.78
5732
1724
+-8
.424
4.81
2-1
0.41
3-1
.93
7-1
.937
0.4
420.
833
0.90
5046
1824
-7.
667
-5.5
7310
.394
1.94
81
.949
0.33
20.
912
1.02
5938
191.
0025
+-1
1.19
44.
802
-13
.04
5-2
.01
2-2
.015
0.3
590.
858
1.18
6145
2025
-10
.518
-8.2
2915
.734
2.03
62
.036
0.32
70.
918
1.38
3865
2126
+-1
1.19
47.
477
-15
.67
3-2
.04
8-2
.052
0.3
740.
832
1.57
9152
2226
-10
.528
-8.1
7715
.620
2.05
72
.062
0.31
80.
938
1.77
7459
2327
+-1
1.18
97.
434
-15
.58
7-2
.05
2-2
.074
0.3
760.
810
1.97
3671
2427
-10
.537
-8.1
4315
.587
2.07
42
.077
0.27
20.
877
2.17
2596
251.
2528
+-1
4.00
27.
434
-18
.22
8-2
.12
6-2
.136
0.3
460.
845
2.41
2996
2628
-13
.345
-10.
780
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Page 45
5.3 CMU Thru-Bolt Push-Out Tests
5.3.1 Connection Failure Mechanisms
Design values for shear applied to the through-bolted masonry connection implemented in the hybrid masonry system are not clearly addressed in ACI 530, although several equations for embedded anchor bolts and headed studs are provided. These may be extended to provide an approximation of the nominal strengths and associated limit states that could be expected in a through-bolted connection.
ACI 530-08 gives an upper bound for nominal shear strength in equation 3-20. This value assumes that the factored load in shear is much greater than the factored moment (i.e. / 0.25), which can be reasonably assumed to be the case for the type of connection under investigation. For this condition, ACI 530-08 limits the maximum nominal masonry shear strength to the following:
6 (ACI 530-08 EQN 3-20) [3]
Evaluation of this equation gives the nominal upper limit shear strength for the specimens in this test as 2 6 10.5 . 7.625 . 2471 1 1000 ⁄ 47.78 , multiplying by two to account for the two potential shear failure planes in the specimen. This upper limit accounts for both masonry shear strength and the nominal shear strength provided for any reinforcement, or (ACI 530-08 EQN 3-19). Since only one #3 bar is provided in the specimen to resist shear along the two potential shear failure planes, it is more reasonable to disregard the contribution from the reinforcement and simply treat the single bar as a means to prevent brittle breakout. With that assumption, the nominal masonry shear strength is reduced only to the capacity from Vnm. This value is determined through the following:
4.0 1.75
0.25 (ACI 530-08 EQN 3-22) [4]
Where, in this case Pu = 0. Three potential assumptions can be made in order to apply and evaluate this equation for the conditions of this configuration. A conservative approach would be to assume , which is true at midspan, since 1⁄ . This equation then reduces to 2.25 . Evaluation of this equation for this experimental configuration gives the following:
2 2.25 10.5 . 7.625 . 2471 1 1000 ⁄ 17.91
Another reasonable approach is to determine the moment at a distance /2 from the support, assuming the crack initiates at mid-depth of the shear depth, dv. This results in 0.5⁄ and 3.125 . Evaluation of this equation based on the stated assumption gives a nominal masonry shear capacity equal to 24.87 kips.
One final approach considered for the evaluation of this term is in consideration of the original intent of the ⁄ ratio, a representation of the ratio of the tensile
Page 46
force induced by the bending moment to the actual shear force. As the tension force causes flexural cracking, the aggregate interlock and/or mechanical state of stress will be less effective to resist shear. For these test specimens, the bending moment is actually applied to the full 32 inch depth and therefore it could be assumed that
0⁄ , giving 4 . Evaluation of this equation based on the stated assumption gives a nominal masonry shear capacity equal to 31.84 kips. By inspection, the assumption that 0.5⁄ provides the closest approximation to the actual mean maximum load, while the first approach also serves to provide a reasonable, yet slightly conservative, value as well.
In addition to shear strength, ACI 530-08 addresses bearing strength and the strength of embedded anchor bolts in shear and/or tension. The ACI 530 specified nominal bearing strength of masonry is given as 0.6 , where the bearing area, Abr, is defined in Section 1.9.5. Since the threaded rod is bearing across the entire width of the wall specimens, the provision for an increased bearing area does not apply. The average load at which initial cracks were observed in the test specimens was between 20 and 21 kips, much greater than expected according to the code-prescribed bearing capacity of 0.6 2471 10.5 . 0.75 . 1 1000 ⁄ 8.48 . Therefore, this strength value provides an extremely conservative estimate of the through bolted connection capacity.
In addition to masonry shear capacity, ACI 530-08 provides four potential limiting equations for the nominal shear strength of headed and bent-bar anchor bolts.
4 (ACI 530-08 EQN 3-6) [5]
1050 (ACI 530-08 EQN 3-7) [6]
2.0 8 (ACI 530-08 EQN 3-8) [7]
0.6 (ACI 530-08 EQN 3-9) [8]
These equations provide an assessment of the nominal shear strength of the connection as governed by masonry breakout, masonry crushing, anchor bolt pryout, and bolt shear yielding, in the order shown. Equation (7) does not apply to the through bolted connection implemented in this test, as it is dependent upon a tensile cone breakout of an embedded anchor and would be highly unlikely in this configuration. Equation (5) depends upon the failure of the projected shear area of one-half of a right circular cone. For the purposes of a through bolted connection, it will be assumed that this projected shear area is the rectangular projection at 45 degree angles from the location of the anchor bolt to the surface of the masonry (Figure 30). This gives the values as 4 21 . 7.625 . 2471 31.84 .
Page 47
Figure 30. Assumed projected shear area for masonry breakout failure (in = 25.4 mm).
A complete, detailed evaluation of these equations is provided in the Appendix of this report. Table 11 summarizes the values obtained for the CMU bolt pushout test specimens along with critical points identified in the actual behavior of the test specimens for the purposes of identifying applicability of current ACI 530-08 design equations. Table 11. Summary of Applicable ACI 530-08 Design Values and Measured Test Results (kip =
4.45 kN).
By inspection, both masonry shear strength and bolt shear yielding provide conservative estimates of the capacity of the thru-bolted connection. The other two applicable limit states, masonry breakout and masonry crushing, are most likely affected by the substitution of the typical embedded connection with a thru-bolted connection, which introduces a set of completely different restraints into the connection. Because of the various additional restraints provided in this system, it appears bolt shear yielding will typically be the governing, and most predictable, limit state in determining the capacity of the masonry assembly. Masonry shear failure may have also governed the behavior of the specimens, but was limited by the presence of the vertical #5 steel located in the cells directly below the threaded rod connection, thereby confirming the importance of the end cell reinforcement in the detailing of the masonry wall.
Masonry Breakout
Masonry Crushing
Anchor Bolt Pryout
Bolt Shear Yielding
Typical Initial
Cracking Load
Mean Maximum Load prior to Failure
24.87 kips 31.84 kips 6.03 kips N/A 19.01 kips 20-21 kips 25.7 kips
Nominal Shear Strength of Anchor BoltTest Results
ACI 530-08 Prescribed Limit States
Masonry Shear
Strength, Vnm
Page 48
5.3.2 Comparison of Wall Specimens (Partially vs. Fully Grouted)
Although the actual maximum loads and displacement of the various test specimens remained relatively consistent regardless of the individual designs, the actual observed behaviors varied slightly. The partially grouted specimens displayed significant propagation of mortar bed separation through almost the entire height of the specimen. In the fully grouted specimens, mortar bed separation was typically limited to only one cell and was restrained by the grout in the adjacent cell. When limiting residual damages is a design objective, fully grouted specimens provided a more desirable outcome.
Page 49
6 Conclusions / Recommendations
Connection detailing often proves to be a critical aspect in achieving desirable energy dissipation levels and controlling structural response during seismic events, and each and every substantial earthquake continues to demonstrate the importance of these connections. The results from this study will serve as the basis for the further development of the connecting components to be used in the Type I Hybrid Masonry Seismic Structural System, as well as facilitate the design of the full-scale models to be tested at a later phase in the project. Based on the scope of this study and the results obtained, the following conclusions, and recommendations for further study, can be made: Connector Plate Tests:
Weld strength and detailing are critical in preventing premature failure of link connector plates.
Buckling behavior controls the maximum load for most connector plates (link and fuse) with thicknesses less than 0.5 in. (12.7 mm). The load capacity generally decreased during post-buckling displacements.
Fuse type A total energy dissipation potential increases with an in increase in
fuse aspect ratio for those aspect ratios addressed in this study.
Fuse type T is recommended in applications where all non-linear behavior is to be achieved in the connector plates due to its superior energy dissipating characteristics.
Link connectors are recommended when the non-linear behavior is to be
achieved entirely within the CMU walls.
Further testing is recommended to determine the appropriate connection geometry for attaching connector plates to the underside of the steel beam due to the significant reduction in connector capacity when a welded bent plate connection was used. Bolted connections and puddle welds may be effective in strengthening this connection. There is a need to explore other potential connection details.
CMU Thru-Bolt Push-Out Tests:
Detailing of boundary steel for thru-bolts located near the edge cell of the CMU wall may play an important role in providing adequate capacity for the connection.
Fully grouting CMU walls may play a significant role in minimizing propagation of mortar joint separation due to lateral loads on thru-bolted connections.
Page 50
Bolt shear yielding and masonry shear failure may both provide a reasonable design value for the strength of the thru-bolted connection nearest to the edge of the CMU wall.
Bolt shear yielding can be a potential limit state used to determine a design value
for the strength of a thru-bolted connection away from the edge of the wall.
Further testing is recommended to investigate the effects of the interaction of a line of multiple thru-bolted connections in a single CMU wall panel, in addition to determining the appropriate restraint conditions for testing the thru-bolted connections—specifically, in the Type I hybrid system there will not be any shear reaction at the top of the wall.
Page 51
7 References
Abrams (2011): Abrams, Daniel P. (2011). “NSF NEESR Research on Hybrid Masonry Structural Systems.” Proceedings to the 11th North American Masonry Conference, Minneapolis, MN, 2011.
Abrams et al (2010): Abrams, D. P., Fahnestock, L.A., Eidini, M., "Basic Mechanisms for
Hybrid Masonry Structures" Proceedings to 2010 ASCE Structures Congress, Orlando, FL, USA, 2010.
Aliaari & Memari (2007): Aliaari, M. and Memari, A. (2007). “Experimental Evaluation of
a Sacrificial Seismic Fuse Device for Masonry Infill Walls.” ASCE Journal of Architectural Engineering, June 2007, 111-125.
ASCE (2005): ASCE 7-05: Minimum Design Loads for Buildings and Other Structures.
American Society of Civil Engineers, Reston, VA, 2005. ASTM (2003): ASTM Standard E488 (2003). “Standard Test Methods for Strength of
Anchors in Concrete and Masonry Elements,” ASTM International, West Conshohocken, PA, 2003.
ASTM (2008a): ASTM Standard A370 (2008). “Standard Test Methods and Definitions
for Mechanical Testing of Steel Products,” ASTM International, West Conshohocken, PA, 2008.
ASTM (2008b): ASTM Standard C1314 (2008). “Standard Test Method for Compressive
Strength of Masonry Prisms,” ASTM International, West Conshohocken, PA, 2008.
ASTM (2008c): ASTM Standard E8 (2008). “Standard Test Methods for Tension Testing
of Metallic Materials,” ASTM International, West Conshohocken, PA, 2008. ATC (1992): Guidelines for Cyclic Seismic Testing of Components of Steel Structures.
ATC-24, Applied Technologies Council, Redwood City, California, 1992. Biggs (2006): Biggs, David T. (2006). “Hybrid Masonry Structures.” Proceedings to 10th
North American Masonry Conference, St. Louis, MO, 2006. Biggs (2011): Biggs, David T. (2011). “Using Hybrid Masonry Bracing For Steel
Frames.” Proceedings to the 11th North American Masonry Conference, Minneapolis, MN, 2011.
Bruneau & El-Bahey (2010): Bruneau, M. and El-Bahey, S. (2010). “Structural Fuses
and Concrete-Filled Steel Shapes for Seismic- and Multi-Hazard Resistant Design.” 2010 NZSEE Conference.
Page 52
FEMA. (2007). “FEMA 461 – Interim Testing Protocols for Determining the Seismic Performance Characteristics of Structural and Nonstructural Components.” Applied Technology Council, Redwood City, California, 2007.
IBC (2006): International Building Code (IBC), International Code Council, Falls Church,
VA 2006. IMI (2009): IMI Technology Brief 02.13.01, Hybrid Masonry Design, International
Masonry Institute, Annapolis, MD, 2009. IMI (2010): IMI Technology Brief 02.13.02, Hybrid Masonry Construction, International
Masonry Institute, Annapolis, MD, 2010. IMI (2011a): International Masonry Institute. “Hybrid Masonry/Steel Details.” Retrieved
from http://www.imiweb.org/design_tools/structural_masonry/details_hybrid.php. IMI (2011b): Masonry Detailing Series. International Masonry Institute. Retrieved from
http://www.imiweb.org/design_tools/masonry_details/index.php MSJC (2008): MSJC 2008: The Masonry Standards Joint Committee, Building Code
Requirements and Specification for Masonry Structures. The Masonry Society, Boulder, Colorado, 2008.
NCMA (2009): NCMA TEK 14-9A 2009: NCMA TEK 14-9A, “Hybrid Concrete Masonry
Design,” National Concrete Masonry Association, Herndon, VA 2009.
A-1
Appendices
A-2
Appendix A: Supplier Test Report for all CMU Blocks used in this Test Series
A-3
Appendix B: Connector Plate Hysteretic Responses
A-4
A-5
A-6
Appendix C: Connector Plate Specimens ATC-24 Analysis Tables
The following ATC-24 based analysis tables were produced using a spreadsheet
to process the corrected data files for the individual connector plate test specimens. The following formula was used to distinguish the individual excursions from each other by identifying the regions in the data where load values changed from positive to zero to negative, and vice versa, thus indicating the end of one excursion and the beginning of the next.
=IF(OR(AND(AND(“n-2”>“n-1”, ABS(“n-2”-“n-1”)>J2), “n-1”>0, n<=0, OR(“n+1”<n, ABS(“n+1”-n)<J2)), AND(OR(“n-2”<“n-1”, ABS(“n-2”-“n-1”)<J2), “n-1”<0, n>=0, AND(“n+1”>n, ABS(“n+1”-n)>J2))), H16+1, H16)
[9]
The formula shown was developed to identify specific locations in the lists of data
points where the two previous values were increasing and the two subsequent values were decreasing and vice versa, as well as distinguishing between actually changes in value and noise due simply to vibrations in the testing equipment.
Table 13. Excerpt of test data and analysis used to identify critical values.
Table, an excerpt from the spreadsheet used to develop the ATC-24 based
summary for specimen P4_T2_FA2-01, is provided for the purpose of illustrating this methodology. For a specific entry in the data set, n, identified by the column labeled “Entry”, formula [10] would first determine if the load value in Column [F] of entry (n-2) was greater than the value of entry (n-1) in the same column. If this was true, then the formula would check if the absolute value of the difference of the two values was greater than some predetermined value of noise in the signal, in order to differentiate between noise and an intentional change in load. In the case of this specimen, any
Entry Time LR4 LR5Dpot‐161‐1915
MTS1Displ MTS2LoadExcursion(delta)
Excursion(force)
SelectDisplSelectedLoad
[A] [B] [C] [D] [E] [F] [G] [H] [I] [J]
1 28.3 0.002965 ‐0.0027379 0.00041 ‐0.001282 0.00486866 0 0 0 0
2 28.35 ‐0.002224 ‐0.0065711 0.00124 ‐0.001099 0.00257903 0 0 0 0
3 28.4 ‐0.003706 ‐0.0005476 ‐0.00041 ‐0.000825 0.00028941 0 0 0 0
4 28.45 ‐0.009636 0.0021904 ‐0.00207 0.000641 ‐0.0065795 1 1 0 0
5 28.5 ‐0.002965 0.0060235 ‐0.00497 0.002472 ‐0.0134483 1 1 0 0
6 28.55 ‐0.001483 0.0071187 ‐0.00539 0.004304 ‐0.015738 1 1 0 0
7 28.6 0.001482 0.0071187 ‐0.00497 0.005403 ‐0.0172643 1 1 0 0
8 28.65 0.003706 0.0038332 ‐0.00539 0.006044 ‐0.0165011 1 1 0 0
9 28.7 0.003706 0.0060235 ‐0.0058 0.005861 ‐0.0134483 1 1 0 0
10 28.75 0.013342 0.0010952 ‐0.00539 0.00522 ‐0.0088691 1 1 0 0
11 28.8 0.013342 ‐0.0021903 ‐0.00497 0.003938 ‐0.0020002 1 1 0 0
12 28.85 0.017789 ‐0.0104042 ‐0.0029 0.002381 0.0102111 1 2 0 0
13 28.9 0.017789 ‐0.0147849 ‐0.00083 0.000183 0.0193696 1 2 0 0
14 28.95 0.022236 ‐0.0169753 0.00124 ‐0.002107 0.0262384 2 2 0 0
A-7
change less than 0.001 was regarded as noise. This noise value was stored in cell J2 of the spreadsheet. If this criteria was met, the formula would then check that load for entry (n-1) was a positive value and that entry n was less than or equal to zero, indicating a negative slope and that the load on the specimen was approaching and either passing or arriving at zero load. If this initial set of criteria was met, the formula would then confirm that entry (n+1) was also greater than entry n, and if it was not, it would determine if this condition was not met due of noise in the data, that is, if the absolute value of the difference between the two entries was less than the noise value. If all of these conditions were met, the formula would iterate the value in cell H, signifying that a zero intercept had occurred and the test was in a new excursion.
Alternatively, for the case of a positive slope in the data, formula [10] would first determine if the load value in Column [F] of entry (n-2) was less than the value of entry (n-1) in the same column, and if it was not, it would determine if this condition was not met due of noise in the data, that is, if the absolute value of the difference between the two entries was less than the noise value. If this criteria was met, the formula would then check that load for entry (n-1) was a negative value and that entry n was greater than or equal to zero, indicating a positive slope and that the load on the specimen was approaching and either passing or arriving at zero load. If this initial set of criteria was met, the formula would then confirm that entry (n+1) was greater than entry n, and that the absolute value of the difference of the two values was greater than the noise value. If all of these conditions were met, the formula would iterate the value in cell H, signifying that a zero intercept had occurred and the test was in a new excursion. Although this formula served to identify a majority of the zero intercepts of the data, there were typically a few excursion start and end values that were not detected. In these cases, the data was manually inspected to identify their location and the excursion count was manually iterated. Once the individualized excursions were identified and separated from each other using this methodology, simple formulas were used to identify such values as the local extrema for each excursion, as well as the values at the beginning and end of each excursion. Using all of these values, simple manipulation and calculations produced the values as shown in the following summary tables for each specimen tested.
A-8
Sp
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594
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[7]
[8]
[9]
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10.
01
1+
-0.1
26
-0.0
14
-0.0
70-0
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77
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47
0.0
00
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21
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3-0
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30.
035
0.0
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0.1
40
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40
0.0
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50
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44
2-
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77
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77
0.0
350
.000
0.0
51-0
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5-0
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0.0
00
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53
+-0
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9-0
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9-0
.014
-0.0
22-0
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60
.17
40
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80
.00
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16
3-
-0.0
70
-0.0
49
0.0
210
.001
0.0
500
.84
6-0
.007
0.0
00
023
70.
02
4+
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81
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28
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70-0
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-0.0
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0.1
04
0.0
46
0.0
00
034
84
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4-0
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40.
070
0.0
050
.081
0.4
94
0.4
94
0.0
00
047
95
+-0
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5-0
.03
5-0
.056
-0.0
25-0
.02
90
.08
80
.06
90
.00
005
810
5-
-0.0
70
-0.0
70
0.0
490
.003
0.0
820
.58
40
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40
.00
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811
6+
-0.1
95
-0.0
28
-0.0
70-0
.028
-0.0
29
0.0
95
0.0
95
0.0
00
078
126
--0
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1-0
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40.
077
0.0
010
.081
0.4
99
0.0
01
0.0
00
091
130.
03
7+
-0.4
47
-0.0
14
-0.0
98-0
.032
-0.0
32
0.0
45
0.0
46
0.0
00
128
147
-0.
202
-0.0
84
0.0
490
.174
0.1
760
.61
40
.74
90
.00
016
115
8+
-0.4
40
-0.0
21
-0.0
70-0
.008
-0.0
31
0.0
29
0.0
13
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195
168
-0.
181
-0.0
84
0.0
630
.164
0.1
760
.63
20
.81
90
.00
022
417
9+
-0.4
61
-0.0
42
-0.0
70-0
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-0.0
31
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25
0.0
17
0.0
00
258
189
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181
-0.1
120
.112
0.1
760
.176
0.6
17
0.8
15
0.0
00
309
190.
06
10
+-0
.70
5-0
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1-0
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48-0
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20
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40
.67
60
.00
038
020
10
-0.
495
-0.0
91
0.0
630
.369
0.3
730
.62
50
.78
70
.00
045
321
11+
-0.7
05
-0.0
35
-0.0
63-0
.147
-0.1
53
0.0
36
0.7
52
0.0
00
520
2211
-0.
495
-0.0
98
0.0
700
.368
0.3
710
.62
00
.75
90
.00
058
723
12
+-0
.70
5-0
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5-0
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41-0
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20
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50
.91
40
.00
065
324
12
-0.
502
-0.0
91
0.0
840
.357
0.3
710
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90
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60
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025
0.1
01
3+
-1.0
96
-0.0
28
-0.1
05-0
.377
-0.3
77
0.0
37
0.7
04
0.0
00
856
261
3-
0.92
1-0
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0.1
260
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0.6
310
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0.7
68
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01
083
271
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96
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14
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98-0
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0.7
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208
281
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220
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0.5
79
0.6
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340
291
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77-0
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85
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26
0.6
55
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481
301
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0.92
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70.
098
0.6
280
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0.6
61
0.7
73
0.0
01
676
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A-9
Sp
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nP
4_T
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594
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310.
25
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A-11
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Exc
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[4]
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[6]
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[8]
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[11]
Exc
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[3]
611.
5031
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2.9
37
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9.8
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Exc
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A-27
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5
Exc
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on
[3]
A-31
Sp
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nP
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Exc
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[3]
310.
25
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2
A-32
Sp
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nP
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Exc
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---
A-33
Sp
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nP
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Exc
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[3]
A-34
Sp
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nP
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Exc
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310.
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Exc
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611.
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A-36
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Exc
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A-37
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Exc
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310.
25
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N 1
6- WA
S N
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PR
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LY C
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LE
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ED
Beg
in R
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. 1
A-54
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15
Beg
in R
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. 2
A-56
Sp
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me
nP
4_T
4_
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A-65
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[1]
[2]
[4]
[5]
[6]
[7]
[8]
[9]
[10
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1]
Exc
ursi
on
[3]
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---
A-66
Appendix D: Tapered Fuse Design Procedure
The procedure outlined here shows the process used to design the fuse width at the top and bottom of the fuse section such that the extreme fibers yield at an equal horizontal force. For these fuses, the procedure relies on a triangular moment diagram increasing linearly from zero at the location of the center of the slotted hole (bolt through masonry) to a maximum at the welded connection to the test setup.
Calculate the stress at the top of the fuse, , and the stress at the bottom of the fuse, , in terms of, and , the width at the top and bottom of the fuse, respectively, by Equation 8. Refer to Figure 31 for the fuse dimensions referenced in the provided equations.
6
6
[10a]
6
6
[10b]
Setting equal to and solving for the ratio of the fuse widths gives Equation 9.
[11]
Equation 9 is valid for any orientation of the fuse, with “top” and “bottom” in reference to the fuse’s orientation in the actual hybrid masonry assembly. Setting a top fuse width,
, equal to 2 inches, for a 4 inch fuse length and the bolt through the masonry at 12
inches from the bottom of the fuse gives 2" 8"/12" 1.633 in. (41.5 mm). This is the width of the bottom of the fuse detail for specimen P4_T4_FT2-01: a tapered fuse with an aspect ratio, ⁄ , equal to 2. The length of the fuse detail is 4 inches (101.6 mm) and the largest width of the tapered fuse is 2 inches (50.8 mm). Similarly, specimen P4_T4_FT3-01 has a 6 inch (152.4 mm) fuse length ( ⁄ = 3) which gives
Figure 31. Tapered fuse design dimensions.
A-67
2" 6"/12" 1.414 in. (35.9 mm) - the width of the bottom of the fuse detail.
