DESIGN OF SELF-CENTERING MOMENT RESISTING FRAME AND EXPERIMENTAL LOADING SYSTEM

Transcription

1 DESIGN OF SELF-CENTERING MOMENT RESISTING FRAME AND EXPERIMENTAL LOADING SYSTEM 1. Abstract Scott Swensen The University of Utah REU Institution: Lehigh University REU Advisors: Dr. Richard Sause and Dr. James Ricles Graduate Student Mentor: Ying-Cheng Lin Recently, much attention has been focused on moment frame designs that sustain little to no residual damage after the design basis earthquake is applied. One method that shows promise is the use of post-tensioning (PT) steel strands that run parallel to the beams to create a self centering effect upon unloading. Additionally, the implementation of friction energy dissipation devices at the beam-column interfaces provides energy dissipation. During lateral loading, the self-centering moment resisting frame (SC-MRF) expands as the beams rotate in relation to the columns creating a phenomenon called gap opening between each beam and column. The floor system must be designed in a way as to not interfere with this behavior. A proposed method of dealing with this expansion involves the transfer of story shear to the SC-MRF at a single rigid bay at each face of the building. This paper is focused on the design of an experimental SC-MRF frame that utilizes a web friction device (WFD). Included is the design and explanation of an experimental force transfer system, floor diaphragm system, and bracing system that were created to mimic the systems that would be implemented in a full-scale building. 2. Introduction 2.1 Post-tensioned Self-Centering Moment Resisting Frames Conventional steel moment frames subjected to earthquake loading dissipate energy and mitigate structure collapse, but many such structures perform these tasks through the development of plastic hinges in the moment frame beams, causing residual structural drift and damage. Recently, researchers have been focusing efforts on developing steel frame designs that will undergo little to no residual damage after the design basis earthquake (DBE) is applied. One such design that is being developed at Lehigh University utilizes high-strength steel post tensioning cables to create moment resistance at the beam-column interface. Experiments have shown that such post tensioning can provide initial joint stiffness similar to that of a conventional welded moment connection (Wolski 2006). When moment is initially applied to the connection, the post tensioning strands hold the beam firmly to the column and both beam flanges are in compression. Once applied moment surpasses moment capacity of the connection, force is completely released at either the

2 top or the bottom flange of the beam at the beam-column interface. This occurrence of released compression is called decompression. As more moment is applied after decompression, gap opening at the beam-column interface is further resisted by energy dissipation devices. When additional moment is applied and forces from the energy dissipation device are overcome, either the top or bottom flange of the beam separates from the column and gap opening occurs. The post-tensioning force in the strands increases at gap opening due to strand elongation (Garlock et al. 2007). This increased tension serves to self-center the frame after loading. 2.2 Energy Dissipation Devices After gap opening of the connection, energy is dissipated through the implementation of an energy dissipation device. In the past, this has been achieved through the installation of steel angles at the beam-column connections that deform plastically at gap opening and closing. This design method, however, requires replacement of the yielded steel angles after the design earthquake is applied. Other energy dissipation devices include viscoelastic copolymer devices, elastomeric devices, and viscous fluid devices. These kinds of devices, however, have limitations such as inconsistent behavior with varying temperatures and loading frequencies. Some of these methods are also not easily constructible (Wolski 2006). Another approach to energy dissipation is to utilize friction devices. These devices utilize two or more pieces of metal that slide past each other, creating frictional forces that act opposite to the direction of sliding. Through Petty s research (1999), it was observed that when UNS 260 half hard cartridge plates are used at the connections, good energy dissipation is provided and the device is relatively durable. The friction coefficient determined to be most appropriate for these brass cartridge plates is 0.45 (Wolski 2006). Friction devices that are located only below the lower beam flange have previously been tested (Ricles et al. 2006). This bottom flange configuration was designed as to not interfere with the floor slab that would be placed over the beams in a real structure. Because of the un-symmetric location of this device, energy is dissipated differently depending on whether the beam is rotating relative to the column from the top flange or the bottom flange. In order to create similar energy dissipation regardless of the point of rotation, a new design will be implemented where the friction device is located at the web of the beam; this device was named the web friction device (WFD). In this design, channel sections with bolt holes are welded to the column on either side of the beam web. The beam web has slotted holes that align with the sets of channel bold holes to allow for rotation of the column relative to the beam. Between the channel section and the beam web are brass friction plates with hole configurations similar to that of the channel sections. These plates create a friction surface between the brass plates and the beam web. As the beam moves relative to the column, the friction devices dissipate energy, regardless of the

3 direction or point of rotation. Figures 1 and 2 show the configuration of the connection and the forces present at gap opening. Figure 1. Connection before decompression Figure 2. Connection after gap opening Based on previous research, the reaction of the connection to applied moment was estimated. Figure 3 illustrates the rotation of the joint based on applied moment. It is observed that when moment is first applied, the connection behaves similarly to a conventional moment frame connection and resists gap opening. At decompression, one of the beam flanges no longer carries compressive force. Gap opening, however, is still inhibited by static friction in the WFD. As the moment is increased, this static friction is overcome and gap opening occurs. Upon unloading, energy is dissipated by the WFD through dynamic friction until gap opening closes. Because this frictional force remains when the gap has closed, complete compression of both beam flanges does not always occur at unloading, though residual connection rotation should be extremely minimal. Because the device is located at the center of the beam web, the connection behaves similarly when the opposite moment is applied. Figure 3 displays the hysteretic behavior of the WFD. Figure 3. Hysteretic behavior of WFD connection

4 2.3 Floor Diaphragm Behavior One inherent problem that occurs in post-tensioned steel frames is posed by gap opening. When gap opening occurs at all bays of the moment frame, the frame width expands. This is illustrated on an exaggerated scale in Figure 4 where the distance between columns is seen to increase at gap opening. If the floor diaphragm were to be rigidly fastened to the moment frame, performance of the frame would be inhibited or damage would occur to the floor diaphragm during lateral loading. In order to achieve desired performance of the frame, the floor diaphragm must be able to transfer lateral shear forces to the frame system while also allowing uninhibited gap opening Figure 4. Illustration of moment frame expansion at gap opening Since the interior framing system is to be composed of rigid connections, frame expansion will only occur on the outside frames. This causes the interior gravity frames to flex in order to reach the contact points in the SC-MRF (Garlock et al. 2006). This flexure of the gravity beams is illustrated in Figure 5. Figure 5. Gravity beam flexure caused by SC-MRF gap opening

5 King (2007) has found that this gap opening can be achieved without damage to the floor diaphragm by using one rigid bay at each floor on each of the outside SC-MRFs of the building. At the rigid bays, a horizontal composite steel plate shear wall (C-SPSW) is installed at the floor-moment frame interface in order to transfer floor diaphragm shear to the moment frame. The C-SPSW is formed by placing a shear plate over half the width of the rigid bay that reaches to the first interior set of beams. This steel plate is set with steel shear studs at a spacing of 12 inches on center and a 6 inch concrete slab is poured over the shear studs. The concrete serves to add to the shear strength of the system and to brace the steel plate against shear buckling. The C-SPSW is connected to the SC-MRF by shear studs that penetrate the plate and the top flange of the SC-MRF beam. Stiffness of the C-SPSW is supplemented by adding boundary elements such as column sections to the edges of the shear plate. The rest of the bays are designed to slide over the moment frames, minimizing resistance. At all other locations except for the four rigid bays at each floor, concrete is poured over ribbed steel decking; this floor system is not connected to any part of the interior framing, ensuring that all story shear will be transferred to the SC-MRFs through the rigid bays. Figure 6 and Figure 7 show the location of rigid bays on a typical floor plan and the design of the C-SPSW. Figure 6. Floor diaphragm design showing rigid and sliding bays

6 C-SPSW Boundary Elements Gravity Beam SC-MRF Column SC-MRF Beam Figure 7. Composite steel plate shear wall at rigid bay 2.4 Bracing In order to prevent weak-axis buckling of the SC-MRF columns and beams, bracing must be provided. In the prototype, gravity beams that run perpendicular to the SC-MRF provide this bracing at the columns. Special consideration must be made for the installation of these gravity beams because the post tensioning cables and web stiffeners run through the flanges of the column parallel to the web at these locations. A special design was formulated that provides bracing but doesn t interfere with the post tensioning or the stiffeners (King 2007). Figure 8 shows the design of the column bracing system. A steel seat angle section is installed under the post tensioning holes and the column stiffeners. This seat extends several inches past the flanges of the SC-MRF column. The gravity beam ends just outside the flange of the SC-MRF column and sits on the seat. Bracing is provided by bolts that run through slotted holes in the seat and fixed holes in the flange of the gravity beam. The slot runs parallel to the web of the SC-MRF column, allowing lateral movement and expansion of the SC-MRF while providing rigid bracing to the SC-MRF column against weak-axis buckling.

7 SC-MRF Column Gravity Beam PT Strands SC-MRF Beams Seat Angle Figure 8. Prototype building column bracing In addition to column bracing, the SC-MRF beams are also braced against buckling. As with the columns, bracing at the SC-MRF beams is complicated by the presence of the post-tensioning cables. In order to brace the beams at these locations, a channel is placed that stretches from the SC-MRF beam to the first interior beam parallel to the SC-MRF in the building (King 2007). This channel fits between the post tensioning and sits on an angled seat to which it is bolted. As with the column bracing, the bolts are fitted through slotted holes that allow lateral translation between the SC-MRF beams and the channel but provide rigid beam bracing. The configuration of the prototype SC-MRF beam bracing is shown in Figure 9. This design braces the SC-MRF beam only if the channel section does not buckle under the bracing force. The bracing channel must be designed to be compact enough to avoid this buckling. SC-MRF Beam PT Strands Bracing Channel Seat Angle Figure 9. Prototype building beam bracing

8 3. System Design 3.1 Frame Design Based on the prototype for the self-centering moment frame that utilizes the beam web friction device, a test specimen was designed for laboratory testing. Because of laboratory size and configuration and equipment restrictions, a scaling factor of 0.6 was utilized to design the test frame based on plans for the full-scale prototype frame. All other frame pieces were subsequently adjusted and scaled to meet seismic code and ensure desirable performance. Because many structural members are only available in fixed sizes, most pieces of the frame did not scale exactly and members had to be chosen that best approximate the desired geometry. The test frame is a four-floor, two-bay design. Figure 10 shows the resulting design frame along with critical dimensions. 3.2 Floor Diaphragm Design Figure 10. Experimental frame The floor diaphragm to be placed on each story of the test frame was designed to simulate a floor system that utilizes the rigid bay concept. In the experimental frame, one bay was designed as the rigid bay. Because of constraints caused by the size of the loading system and the configuration of the testing area, the experimental floor system design is much smaller than that of a conventional building. This means that though the testing of this frame will not simulate the actual floor diaphragm weight, the behavior at frame expansion should be similar. Because each beam is fitted with cover plates and shear plates that lie on top of the SC-MRF beam flanges, the floor system was designed in a

9 way as to not be inhibited from sliding by these pieces. For ease and consistency of design, these plates were all designed to be the same thickness, ½ inch. The experimental floor system was designed to be installed on the laboratory floor before assembly of the frame. The floor system will then be attached to the SC-MRF beam and transported to position on the frame. First, a 1.5C steel decking sheet will be placed on the beam cover plates, the ribbing running parallel to the length of the beam. This decking will be 15 inches wide and 197 inches long in each bay. Decking must terminate 4 inches from the face of each column to ensure that the floor diaphragm will not impact the column upon frame expansion. The plate must be cut so that a trough in the ribbing runs along its centerline. This trough in the steel decking will be placed along the centerline of the beam. In places where the steel decking is not supported either by the cover plates or the shear plate, ½ inch-thick shims will be placed under the trough of the decking at the centerline of the beam to ensure that the decking does not sag. These shims will be placed at 12 inches on center at these locations and welded to the top of the SC-MRF beams. In the rigid bay, shear studs will be placed through the steel decking and the shear plates only at the locations of the shear plates. These shear studs will be placed 6 inches from the center line of the beam and spaced at 12 inches on center. A four-inch concrete slab will then be poured over each stretch of the steel decking, leaving spaces at the columns. Except for the shear plate connections to the beam at the rigid bay, the floor system must have no other connection to the frame system. Figure 11 illustrates the design of the floor system on the rigid bay beam. Figure 11. Design of floor diaphragm at rigid bay In order to accurately simulate the floor sliding effect, the slabs on each of the two bays must be connected. This is somewhat difficult because of the presence of the interior SC- MRF column that occupies most of the space between the two slabs. A special design was formulated in which the two concrete slabs on each floor will be connected by steel rebar that extends around each side of the central column. Two bars must be embedded into each slab so they fit between the column and the loading beams. Both the embedded length and the overlap length of each bar must be at least 12 inches in order to provide sufficient development length in the rebar. After assembly of the frame, the two sets of rebar will be lashed together and a narrow 1½ inch wide channel of concrete will be

10 poured between the SC-MRF column and the loading beams. Figure 12 shows the configuration of the connection between the two slabs. Figure 12. Connection of floor slabs In addition to this connecting element between the two concrete slabs on each floor, elements that represent the geometry of the floor system at the outside SC-MRF columns will also be implemented. First, a connection to the slab that is furthest away from the actuator was designed. Because of size and weight considerations, it was determined that these simulated pieces of the floor diaphragm shouldn t be made of concrete. Instead, wood pieces were designed. Though the weight and material properties of the real floor system will not be simulated by these pieces, the gap opening s effect on the floor diaphragm will be clearly observed during testing. Figure 13 shows the design of the proposed wood slab extension at the end of the frame that is furthest from the actuator. During the construction of the concrete slabs, threaded bolts will be embedded into the outside surface of the concrete slab. The U-shaped wood slab extension will be fitted with nuts that can be used to secure the piece to the concrete slab. Figure 13. Simulated floor system away from actuator At the end of the frame that is near the actuator, the size of the simulated floor diaphragm is limited by the proximity of the actuator attachment to the loading system. Because of this, a smaller simulated floor diaphragm was designed. Figure 14 shows the configuration of this system at the end of the frame closest to the actuator. Note that

11 instead of a U-shaped wood section, a rectangular section is attached to the slab at each side of the exterior SC-MRF column. Other design features such as the threaded bolt attachment are similar to the simulated floor system implemented at the far end of the frame. 3.2 Story Shear Transfer System Figure 14. Simulated floor system at actuator An actuator attached to the end of a force transfer system will provide simulated story shear for each floor. For ease of construction and implementation, a force transfer system used previously in another experiment will be altered to meet the needs of testing this frame. The force transfer system is composed of two sets of loading beams that sit 15.5 inches apart. These loading beams are composed of HSS12x12x¼ steel sections filled with concrete. At the midpoint of each SC-MRF beam, these loading beams are replaced with loading links. The loading links are composed of two vertical steel plates between two steel end plates. At the end of the series of loading beams, the two sets of beams are held together with steel spreader beams. These beams serve to keep the distance between the two loading beams relatively constant. The entire system will sit between two sets of rigid framing at the test site. The rigid framing will serve to brace the loading system and the test frame. The rigid frame will be simulated at the bay closest to the actuator. At this bay, four shear plates will be installed to represent the C-SPSW system. Each of these plates will be ½-inch thick, 6 inches wide, and 30 inches long. These plates will be welded to the side of each loading beam, ensuring that the bottom of each shear plate lies at the midheight of the beam. The other end of each plate will then be welded to the top of the top flange of the SC-MRF with a ¼ inch fillet. When the actuator at each floor pushes the force transfer system, the entire story shear force will be translated to the shear plates, and then into the top flange of the beam at the rigid bay. The configuration of the loading beams and the shear plates is shown in Figure 15.

12 Loading Beams SC-MRF Columns Loading Links SC-MRF Beams Shear Plates Spreader Beam Figure 15. Force transfer system Actuator Force 3.3 Gravity Force Transfer System In the prototype building, the SC-MRF beams receive a small gravity load due to the member s observed live and dead loads distributed over its tributary area. In order to find if it is appropriate to distribute the weight of the force transfer system to the beams in the SC-MRF to simulate this gravity force, an analysis was made of the expected gravity loads distributed to the SC-MRF beams on the prototype. The desired gravity loads on each floor were determined based on scaled live and dead loads distributed over the tributary areas of each beam. The weight of the force transfer system at each floor was then estimated. Table 1 shows the resulting desired floor loads and the weight of the force transfer system. Table 1. Comparison of desired floor load and weight of force transfer system Floor Scaled Load (plf) Length (ft) Weight of Force Floor Load Transfer System (kips) (kips) RF F F F It is observed that the weight of the force transfer system is considerably larger than the desired weight on the beam to simulate an actual distributed floor load. An option that was considered was to transfer some of the weight of the force transfer system to the beams, and the remainder to the SC-MRF columns. This design, however, uses more members and would cost more money and take more time to construct. Also, the precise distribution of load in the columns and beams would still be unknown. Because of this, it was decided to place the entire weight of the loading beams onto the SC-MRF columns. This was achieved through the implementation of a column force transfer device. The device is shown in Figure 16.

13 Figure 16. Force transfer device In this device, a 10 inch long 6x6x¾ angle is welded to the tips of the flanges of each SC- MRF column using a ¼ inch fillet weld on the outside of the column flanges. This angle is reinforced with ½ inch thick stiffeners at each end of the seat angle. Four inch long 3x2x½ angles are then welded to the bottom of the loading beams and accept one end of each seat angle. Enough tolerance between the seat angles and the friction angles must be provided in order to ensure that the loading system will fit into place. Also, because these loading beams may spread apart during testing due to the actuator force, this maximum deflection distance must be calculated to ensure that the friction angles will not slide off the seat angles. Through this system, gravity load is transferred by shear through the seat angle and into the column flanges. Relative movement of the loading system in relation to the frame is allowed in the plane of the frame, but prohibited in all other directions. Figure 17 is a three-dimensional rendering of the device as seen from below. Loading Beams SC-MRF Beams SC-MRF Column Loading Beam Angle Seat Angle Figure 17. Gravity force transfer device at SC-MRF column

14 3.4 Bracing System The steel angle welded to the bottom of the loading beams at the columns provides bracing when the angles come into contact with the seat angles of the gravity force transfer device. Because the rigid frame in the laboratory setup braces the loading beams, the loading beams can brace the columns through the steel plates on the gravity force transfer device. In order to simulate SC-MRF beam bracing, it was necessary to design a new bracing device. Because of restrictions imposed by the small area available for bracing placement, a channel section bracing configuration similar to that devised for the prototype structure could not be achieved. Instead, the bracing device shown in Figure 18 was devised. This illustration shows the beam braces as seen from the underside of the SC-MRF beam. Loading Beams Beam Braces SC-MRF Beam Figure 18. SC-MRF beam bracing detail In this system, two triangular plates that are welded to the bottom of each loading beam project toward the bottom of the SC-MRF beam. At the surface between these plates and the SC-MRF beam, another plate is welded to the first plate to form a T-shape. This device is installed at both sides of the beam at the same location. This design provides ample bracing while allowing uninhibited lateral displacement of the SC-MRF frame. Two of these braces will be placed on each of the SC-MRF beams at their third-points. 4. Conclusions Upon comparing the experimental force transfer system to the actual force transfer that would be seen in the prototype building, it was concluded that the experimental system simulates the actual system quite accurately. Figure 17 shows an overview of the experimental loading system and Figure 18 shows the force distribution in the prototype frame.

15 Figure 19. Experimental loading system Figure 20. Prototype loading system It is noted that in the experimental frame the entire gravity force from the loading system is placed on the three columns. In the prototype building, the gravity force is distributed over the SC-MRF beams. The gravity load on the beams, therefore, is not comparable to the load in the experimental frame. This, however, is not expected to alter experimental data considerably because the tributary area and floor loads that should be placed on the beam are fairly small in comparison to the lateral loads that will be applied to simulate story shear. The story shear in the experimental frame is distributed by two shear plates on each side of the beam in two locations along the rigid bay. In the prototype building, there is only one shear plate on one side of the rigid bay that is the width of half the bay. The out-ofplane length of the prototype shear plate stretches to the interior frame of the structure, which is the length of one of the SC-MRF beams. But the shear plates in the experimental system are only six inches long each. Because of this difference in shear plate length, the potential of the prototype plate to buckle in shear will not be observed. The plates will, however, provide a means to transfer the lateral story shear to the frame in much the same way as the C-SPSW in the prototype. The floor system in the experimental frame is much smaller than the actual tributary area of the floor diaphragm of the prototype structure because of testing restrictions. Also, the slab will be poured in two sections before placement and connected in place, as opposed to pouring the whole slab at once. In the test structure, the spaces around the columns are simulated with wood slab extension pieces instead of concrete. These differences, however, are not expected to cause a problem in the observation of gap opening on the floor slab. This gap opening behavior during lateral loading is the primary structural behavior that is being investigated in the floor system.

16 The experimental bracing system designed is quite similar to the prototype bracing system. While comparable members could not be used to brace the SC-MRF, similar bracing reactions were achieved at each column through the implementation of the gravity force transfer device. While the prototype building calls for three bracing points per beam, the experimental system will consist of two bracing points per beam. This difference is not expected to affect experimental data significantly since out of plane beam buckling is not expected to be significant. The effectiveness of this frame and loading system design will be observed during laboratory testing. If the loading system is found to behave satisfactorily and to effectively imitate the prototype loading scenario, this designed loading system could be used in future earthquake testing scenarios. 5. Acknowledgements This research was made possible through funding from the George E. Brown, Jr. Network for Earthquake Engineering Simulation (NEES) Program of the National Science Foundation (NSF). Dr. Richard Sause and Dr. James Ricles of the Advanced Technology for Large Structural Systems (ATLSS) Center provided extensive support, guidance, and feedback throughout the research process. Ying-Cheng Lin, a graduate student at Lehigh University, also supplied essential assistance. 6. References AISC (2005). Steel Construction Manual, 3rd Edition. American Institute of Steel Construction, Chicago, IL. Garlock, M. (2002). Full-Scale Testing, Seismic Analysis, and Design of Post- Tensioned Seismic Resistant Connections for Steel Frames, Ph.D. Dissertation, Civil and Environmental Engineering Department, Lehigh University, Bethlehem, PA. Garlock, M., Jiu, J., and King, A. (2006). Construction Details for Self-Centering Moment Resisting Frame Floor Diaphragms. Proceedings, U.S.-Taiwan Workshop on Self-Centering Structural Systems. Taipei, Taiwan. Garlock, M., Sause, R., and Ricles, J. (2007). Behavior and Design of Posttensioned Steel Frame Systems. Journal of Structural Engineering, 133(3), Herrera, R. (2005). Seismic Behavior of Concrete Filled Tube Column-Wide Flange Beam Frames. Ph.D. Dissertation, Civil and Environmental Engineering Department, Lehigh University, Bethlehem, PA. King, A. (2007). Design of Collector Elements for Steel Self-Centering Moment Resisting Frames. M.S. Thesis, School of Civil Engineering, Purdue University, West Lafayette, IN.

17 Petty, G. (1999). Evaluation of a Friction Component for a Post-Tensioned Steel Connection. M.S. Thesis, Department of Civil and Environmental Engineering. Lehigh University, Bethlehem, PA. Ricles, J., Sause, R., Wolski, M., Seo, C-Y., and Iyama, J. (2006). Post-Tensioned Moment Connections with a Bottom Flange Friction Device for Seismic Resistant Self-Centering Steel MRFs. Proceedings, 4th International Conference on Earthquake Engineering. Taipei, Taiwan. Sause, R., Ricles, J., Liu, J., Garlock, M., and VanMarcke, E. (2006). Overview of Self- Centering EQ Resistant Steel Frames Research. Proceedings, 2 nd U.S.-Taiwan Workshop on Self-Centering Structural Systems. Taipei, Taiwan.