EXPERIMENTAL TESTING OF SELF-CENTERING CONCRETE WALLS AND WALL-TOFLOOR CONNECTIONS
K M TWIGDEN, J WATKINS, Q MA, R S HENRY Department of Civil and Environmental Engineering, The University of Auckland, New Zealand
S SRITHARAN, Z CHENG Civil, Construction and Environmental Engineering, Iowa State University, USA
SUMMARY This paper presents an overview of two current experimental programs investigating the seismic response of self-centering walls and behaviour of a connector that facilitates the connection between the floor and a rocking wall. The first series of experiments investigated unbonded post-tensioned (PT) precast concrete wall systems, with and without additional energy dissipating elements. The walls were subjected to pseudo-static cyclic loading, snapback testing and forced vibration testing. The analyses of the experimental results focused on the moment contribution ratio, the effect of this on residual drift, and the equivalent viscous damping ratio for each wall configuration. The second series of experiments investigated a wall-to-floor connector that is suitable for use in conjunction with self-centering, rocking walls. The experimental tests demonstrated the connection has suitable tension and shear strength. The connection required minimal reinforcement around the insert embedded in the wall panel to ensure sufficient transfer of shear forces with minimal spalling of concrete. A sub-assemblage test comprised of a selfcentering wall and a floor unit will verify the connection’s ability to isolate the floor system from the uplift and rotation of the wall. INTRODUCTION There has been extensive research in recent years investigating low-damage seismic resisting systems. With approximately 80% of the multi-storey buildings in the Christchurch CBD demolished following the earthquakes, the potential benefit of low-damage seismic resisting systems is obvious and they will likely be desired during rebuilding of the city. Unbonded PT precast concrete walls have previously been shown to demonstrate excellent seismic performance (Priestley et al. 1999; Perez et al. 2007; Restrepo & Rahman 2007; Sritharan et al. 2008). However, only a limited number of experimental tests have been conducted to investigate the dynamic response of PT concrete walls (Marriott 2009). The behaviour of an unbonded PT wall under lateral loading is characterised by a rocking mechanism. The post tensioning elements are typically designed to remain elastic during design level earthquake loading to ensure a sufficient self-centering behaviour. The rocking mechanism results in a single crack opening at the wall base, ensuring damage is limited to the wall toes. Due to the lack of inelastic response, PT concrete members provide limited energy dissipation and can be designed with hysteretic energy dissipating elements. The controlled rocking mechanism and uplift at the wall base causes a relative vertical
displacement between the wall panel and floor diaphragms. Unless allowance is made to accommodate this relative vertical displacement, this displacement incompatibility will significantly influence the building’s seismic response, adversely affecting the wall's ability to self-center and damage the floor diaphragms. Two experimental test programs are currently being undertaken collaboratively at the University of Auckland in New Zealand and Iowa State University in the US to investigate the aforementioned issues associated with self-centering concrete walls. The first set of experiments involves the cyclic and dynamic testing of self-centering PT concrete walls, to investigate several design parameters and their influence on the seismic response. The second set of experiments investigates a novel wall-to-floor connector, which allows the floor diaphragm to be isolated from the uplift and rotation of a rocking wall while allowing inertia forces to be transferred from the floor to the wall. TESTING OF SELF-CENTERING CONCRETE WALLS An extensive experimental test program is currently underway investigating the cyclic and dynamic behaviour of unbonded PT concrete walls with and without additional hysteretic energy dissipating elements. The walls are subjected to a range of loading conditions including pseudo-static reverse cyclic, snap-back free vibration, and forced vibration. Future testing will incorporate shake table tests and field tests. Test specimen This series of tests consisted of two different wall systems in a total of four different wall configurations. Details of the wall configurations are summarised in Table 1 and the crosssections of the two different wall systems are shown in Figure 1. The two wall systems investigated were Single Rocking Wall (SRW) and Precast Wall with End Columns (PreWEC). The SRW consisted of a single precast concrete panel that was 0.8 m long, 3 m high and 0.125 m thick. The PreWEC system consisted of the same precast concrete wall panel with the addition of end columns attached by O-shaped energy dissipating steel connectors. An in-depth explanation of the PreWEC concept can be found in Sritharan et al. (2008). All of the wall configurations were post-tensioned with either two or three, 15.2 mm diameter prestressing steel strands. The end columns were constructed from concrete filled square hollow sections and were post-tensioned to the foundation with a single high-strength steel bar. A steel angle base frame constructed from 25x25x5 equal angle was cast into each precast wall end for additional confinement and protection of the cover concrete. Since only minor damage was expected in the wall toes, the specimen had a symmetrical design that allowed the wall to be inverted and tested on both ends. A total of five wall panels, or ten wall ends, were constructed for this series of testing. The combinations of loading type and wall configuration details for each wall end are summarised in Table 1. The level of post-tensioning and the number of energy dissipating O-connectors were varied in the PreWEC designs. The moment contribution ratio between self-centering and energy dissipating components is represented by λ, which is calculated using Equation (1) (New Zealand Standard 2006). A lower λ indicates a higher strength contribution from the energy dissipating devices in comparison to the strength contribution of the post tensioning elements and axial load. 𝜆=
𝑀𝑝𝑡 + 𝑀𝑁 𝑀𝑆
where 𝑀𝑝𝑡 , 𝑀𝑁 , and 𝑀𝑆 are the flexural strength contributions from the post-tensioned tendons, the axial load, and the energy dissipating devices respectively.
(1)
Table 1: Test Matrix for the Self-centering Concrete Walls Tests
Wall
PT
SRW
[email protected]
PreWEC-1
[email protected]
PreWEC-2
[email protected]
PreWEC-3
[email protected]
Wall End 1 2 3 4 5 6 7 8 9 10
Loading Type
No. of connectors per joint
λ
-
-
Pseudostatic cyclic
Free vibration
Forced vibration
√
√
Shake table
√ √ √ 4
√
2.3
√ √
√ 6
1.3
√
6
1.0
√
√ √
Figure 1: Cross-section of the test specimens Test setup A photo of the test setup used for the pseudo-static cyclic testing is shown in Figure 2(a). The wall panel in each test was seated in a pocket on top of the foundation that was posttensioned to a strong floor. Grout was flowed into a pocket in the foundation to provide an even bearing surface at the wall-to-foundation interface. Concrete mass blocks were attached to the top of the wall providing anchorage for the tendons, seismic mass for the dynamic testing, and a loading beam for the actuator and eccentric mass shaker (EMS) to be attached. A lateral support frame provided out-of-plane restraint to the wall. For the pseudostatic reverse cyclic testing, a ± 300 kN capacity hydraulic actuator applied the lateral displacements to the top of the wall. Free vibration testing was undertaken using a quick release mechanism that could be remotely triggered to release the wall from a specified lateral displacement and an EMS was placed on top of the mass block to conduct the forced vibration testing. Loading protocol Only the pseudo-static cyclic testing has been completed at the time of writing, and the preliminary results are discussed herein. The loading history applied to the wall for the pseudo-static reverse cyclic tests was based on ACI ITG-5.1 recommendations (2008). Three force based cycles were first applied to each wall followed by displacement controlled
cycles with the maximum reaching 3% drift. The SRW was tested first to provide a baseline measure for comparison with the PreWEC system. Following the SRW, PreWEC-1 and PreWEC-2 were tested using the same lateral displacement protocol.
b) Wall condition at 3% drift (PreWEC-2)
a) Test setup
c) Connector deformation at 2% drift (PreWEC-2)
Figure 2: Test setup and PreWEC 2 observations Experimental Observations and Results The three wall configurations subjected to reverse cyclic loading performed exceptionally well when compared to the expected performance of traditional reinforced concrete walls. In each test, the wall panel experienced only minor damage, limited to a small amount of spalling in the wall toes at large lateral drifts. The essentially undamaged PreWEC-2 wall is shown at 3% lateral drift in Figure 2(b). The lateral force-displacement response of SRW is shown in Figure 3(a), whose response envelope approximately followed a bilinear elastic response. Two cyclic tests were performed on the same PreWEC-1 wall. The first test (PreWEC-1a) relied solely on the connectors to transfer the displacement demand into the post-tensioned columns as lateral displacement was applied to the wall. Following the first test, the fractured O-connectors were removed and new connectors were welded onto the PreWEC-1a wall. A restraint was added during the second test (PreWEC-1b) to enable the columns to displace laterally with the wall without relying on the O-connectors. The two lateral force-displacement responses of PreWEC-1a and b are presented in Figure 3(b). A similar global forcedisplacement response was observed between the two PreWEC-1 tests. The main difference is the presence of small stiffness degradation in the PreWEC-1a test which is barely noticeable in the PreWEC-1b test. The lack of stiffness degradation during PreWEC-1b directly demonstrates the influence of concrete damage and pre-stress loss on the force-
125
125
100
100
Lateral Force Resistance(kN)
Lateral Force Resistance(kN)
displacement response, despite the quantity of damage being very small. It is also important to note that a more desirable force-displacement response was achieved for PreWEC-1b as the O-connectors fractured at 3% drift instead of 2.5% drift.
75 50 25 0 -25 -50 -75 -100
SRW
-125 -3
-2
-1 0 1 Lateral Drift (%)
2
75 50 25 0 -25 -50 -75 PreWEC-1a PreWEC-1b
-100 -125 -3
3
-2
-1 0 1 Lateral Drift (%)
2
3
b) PreWEC – 1a and b
a) SRW
Lateral Force Resistance(kN)
125 100 75 50 25 0 -25 -50 -75 -100
PreWEC-2
-125 -3
-2
-1 0 1 Lateral Drift (%)
2
3
c) PreWEC-2 Figure 3: Lateral force-displacement results for cyclic testing The force-displacement response of PreWEC-2 is presented in Figure 3(c). The PreWEC-2 showed an increase in energy dissipation over both the SRW and PreWEC-1 due to the presence of 1.5 times the number of connectors. Fracture of the O-connectors started to occur at 3% drift, similar to that observed in PreWEC-1b cyclic test. Figure 2(c) depicts an Oconnector at the start of the test (Left) and then at 2% drift (Right). The relative vertical displacement is evident from the change in location of the horizontal lines marked on the Oconnector. The calculated equivalent viscous damping (EVD) for each cycle of the four wall tests is shown in Figure 4(a). A large increase in damping for the PreWEC is observed in comparison to the SRW. Interestingly the SRW's EVD is approximately constant at 5%. The difference in EVD between the PreWEC1a and b is due to two factors, first, the damping provided by the permanent deformation in the wall toe would not have occurred in the
second test (PreWEC-1b), second, the increased horizontal load on the O-connectors would have caused further plastic strain and therefore energy dissipation. At 2% lateral drift, the increase in EVD between PreWEC-1b (14%) and PreWEC-2 (18%) is 4%. Considering that the 5% inherent damping is present regardless of additional energy dissipating devices, increasing the connector number by 50% has also approximately increased the EVD by 50%. The residual drift was defined as the displacement at zero lateral force after unloading from the peak displacement of a cycle, and these are plotted for each of the wall tests in Figure 4(b). It is shown that residual drift increases with an increase in energy dissipating elements, and residual drift increases with lateral drift. It is interesting to note that as O-connectors fracture the residual drift recovers and decreases. For the SRW, the residual drifts only exceeded 0.1% for cycles greater than 1% lateral drift. 1
20
SRW PreWEC-1a PreWEC-1b PreWEC-2
15 10
0.6 0.4 0.2
5 0 0
SRW PreWEC-1a PreWEC-1b PreWEC-2
0.8
Residual Drift (%)
Equivalanet Viscous Damping (%)
25
0.5
1 1.5 2 2.5 Lateral Drift Cycle (%)
3
3.5
a) Equivalent viscous damping
0 -3
-2
-1 0 1 Lateral Drift Cycle (%)
2
b) Residual drift
Figure 4: Equivalent viscous damping and residual drift from cyclic testing WALL-TO-FLOOR CONNECTION TESTING Experimental testing is currently being conducted to investigate a new connector suitable for wall-to-floor connections for buildings designed with self-centering walls. As previously noted the connector is intended to isolate the floor system from the uplift and rotation of the wall system while allowing inertia forces to be transferred from floor to wall. The connector achieves this isolation through the use of a vertically slotted insert embedded into the wall panel and a horizontal V shaped plate affixed to the floor diaphragm. Three different series of tests have been planned to characterise both the connector behaviour and system performance. A combination of monotonic and pseudo-static reverse cyclic direct shear or tension loading is being applied during each of the tests, as summarised in Table 2. First, individual connectors were tested with a steel block replicating the embedded insert to characterise the connector behaviour, strength and ductility. Second, small concrete sections with the embedded inserts were tested to determine the influence of the insert’s anchorage in the concrete. Lastly, a sub-assemblage test representing a single wall-to-floor section will be tested to investigate the global and local responses of the wall-tofloor assembly and to validate the connections' ability to isolate the floor diaphragm from the uplift and rotation of the wall.
3
Table 2: Wall-to-floor testing program Specimen Phase 1
Steel Block & Connector
Phase 2
Embedded Insert & Connector
Phase 3
Wall Panel & Floor System
Loading Protocol Monotonic Tension Cyclic Monotonic Shear Reverse Cyclic Monotonic Shear Unidirectional Cyclic Reverse Cyclic
Test setup The test setups in phase one and two of the experimental program are shown in Figure 5. During phase one, the supporting steel assemblage shown in Figures 5(a) and 5(b) allowed the connectors to be tested in tension and shear. To isolate the performance of the connectors, the slotted insert was emulated by a solid steel block that was designed to have a higher strength than the connector being tested. Stiffeners were added to the supporting steel plate assemblage to ensure that the large moments generated as a result of the eccentricity between the connector and testing machine’s actuator were resisted elastically with minimum deformation to the plate. The second phase of testing incorporated the influence of the embedded inserts anchorage on the connector performance. As shown in Figures 5(c), slotted inserts were cast into concrete blocks 915 mm long, 533 mm wide and 127 mm thick. Two configurations of the concrete blocks were cast, one without any reinforcement and the second with two 19 mm U shaped reinforcing bars around the insert. The concrete blocks were held rigidly in place to the testing machines masthead through clamping. A steel tube fixed to the testing machines base was displaced vertically to apply the shear force to the connector. At the time of testing, the concrete unconfined compressive strength was measured to be 65 MPa. Steel slot
Steel slot
Embedded insert
Steel tube Stub Concrete block
Connector Connector a) Phase 1 shear
b) Phase 1 tension
Connector c) Phase 2 shear
Figure 5: Wall-to-floor connection phase 1 and 2 test setups The loading protocol used for the tests was a combination of pseudo-static monotonic, unidirectional cyclic, and reverse cyclic loading, with loading rates of either 80 kN/min or 4.5 mm/min. Unidirectional cyclic loading was used during the phase two tests due to limitations of the laboratory equipment. Four to six displacement transducers were used
during each test to measure the relative shear/tension displacements, in and out-of-plane rotations and column uplifts. Experimental Observations and Results Selected results from the first two phases of connector testing are presented in Figures 6 and 7. Figure 6 shows the measured force-displacement response for the phase one connector in tension. This highlights a maximum tension strength of approximately 200 kN. The tension capacity of the connector was governed by the yield strength of the bolt that held the connection together. An examination of the connector V plate did not indicate any signs of inelastic behaviour during the test; therefore only the bolt required replacement before retesting. Furthermore, the magnitude of tensile force expected for a wall-to-floor connection is typically a small percentage of the shear force, and the connector was found to have comparable tension and shear capacity. Therefore, the tension force on the bolt will not significantly influence the performance of the connector.
Figure 6: Wall-to-floor connection Phase 1 cyclic tension The measured force-displacement results of the reverse cyclic shear loading in phase one is shown in Figure 7(a). The maximum shear strength recorded for the connector was 183 kN with an elastic stiffness of approximately 65 kN/mm. A decrease in the stiffness only occurred after the initiation of cracking at the inner section between the V and stem at the first cycle of 4.85 mm, as shown in Figure 7(a). Failure in both phases of shear testing occurred when the connector fractured due to accumulated inelastic strains at the inner section between the V and stem. The accumulation of strain was due to the stem flexing back and forth from the reverse cyclic loading. The horizontal line in Figure 7(a) shows that after unloading, the connector underwent a small transverse deformation before reloading in the opposing direction. This feature will allow the connector to accommodate small horizontal movements that are associated with wall rocking. Since multiple connectors will be used between a rocking wall and floor system, the zero resistance between unloading and reloading of an individual connector response will not be a concern. This will be verified in the system test. The force-displacement response from the unidirectional cyclic phase two shear test is shown in Figure 7(b). The maximum strength and stiffness of the connection in phase two was measured as 224 kN and 106 kN/mm respectively. Phase two testing confirmed that the anchorage of the embedded insert in the concrete block was sufficient, with only minor spalling of surface concrete observed around the embedded insert at loads greater than 200 kN. An increase in the strength between phase two and phase one was observed, however, this could be attributed to the unidirectional cyclic loading and lower cycles in phase two. The results plotted in Figure 7(b) also indicate that the connection can be reliably
designed to resist loads of up to 135 kN within the elastic range without the need for reinforcement around the embedded connector. The addition of U shaped reinforcement bars around the insert limited the minor spalling of concrete but had no effect on the connector strength.
a) Phase 1
b) Phase 2
Figure 7: Wall-to-floor connection cyclic shear testing Sub-assembly test The final phase of the test program will verify the ability of the connector to transfer a combination of shear and tensile forces while isolating the floor diaphragm from the wall uplift. The sub-assemblage test comprises of a 1/3 scale prototype unbonded PT concrete wall panel with steel beams representing the floor diaphragm. The wall panel is 4673 mm high, 2260 mm long and 178 mm thick and grouted into a pocket in the foundation. To simulate the load path in a real building, an actuator will be used to push and pull the floor beams horizontally. In turn, the wall-to-floor connections will transfer the shear and tensile forces generated into the wall system, while isolating the floor beams from the self-centering walls relative displacement with the slotted insert. The calculated maximum base shear for the wall is 480 kN, which is sufficient to either yield two wall-to-floor connections or be resisted by four connections within their elastic range. The wall will also subject the connections to equivalent wall uplift and rotation of the full-scale prototype wall. Two to four connectors per test will be utilised at different locations along the wall to capture the connections' ability to distribute forces throughout the wall-to-floor interface.
Figure 8: Phase 3 Test Setup
CONCLUSIONS Four pseudo-static cyclic tests on SRW and PreWEC systems were presented. The results showed approximate bi-linear and flag-shape hysteresis responses for both systems respectively. The equivalent viscous damping for the SRW was calculated to be 5% and the incorporation of O-connectors in the PreWEC system increased the EVD to between 14% and 18% at 2% lateral drift. The increase in EVD appears to be directly proportional to the increase in O-connectors. The residual drift was found to increase with decreasing moment contribution ratio and increase with increasing lateral drift. These experimental tests provided definitive examples of the cyclic behaviour of post-tensioned walls systems and validated their superior performance when subjected to large lateral drifts. Three phases of testing on a wall-to-floor connection detail suitable for self-centering walls were presented. The results from the first two phases characterised the connections behaviour when subjected to shear and tension loading. The measured tension and shear capacity of the connector were 200 kN and 224 kN respectively. The tensile strength was governed by the bolt yield strength and the V plate fracture during the shear tests. In shear loading, the connector resisted loads up to 135 kN elastically and had an elastic stiffness of 106 kN/mm. The final phase of testing will verify the connection’s ability to isolate the floor diaphragm from the uplift and rotation of the wall system while allowing inertia forces to be transferred from floor to wall. ACKNOWLEDGEMENTS The authors would like to acknowledge the assistance provided by Dan Ripley, Mark Byrami, Alex Shegay, James King, Doug Wood and Owen Steffens and are grateful for materials donated by BBR Contech, Mapei, Sika, Stresscrete Northern Ltd., Williams Form Engineering Corp. and B.S Italia. Financial support for these projects was provided by the University of Auckland, Fulbright New Zealand and the Earthquake Commission of New Zealand, National Science foundation in the US (CMMI Grant No: 1041650) and B.S Italia. REFERENCES ACI Innovation Task Group 5. (2008). Acceptance criteria for special unbonded posttensioned precast structural walls based on validation testing and commentary : an ACI standard. Farmington Hills, Mich.: American Concrete Institute. Marriott, D., (2009) The Development of High-Performance Post-Tensioned Rocking Systems for the Seismic Design of Structures. Civil Engineering, University of Canterbury Christchurch. New Zealand Standard 2006. Concrete Structures Standard NZS 3101. Perez, F. J., Sause, R. and Pessiki, S. (2007). Analytical and experimental lateral load Behavior of unbonded posttensioned precast concrete walls. Journal of Structural Engineering-Asce Vol. 133 No. 11: 1531-1540 Priestley, M. J. N., Sritharan, S. S., Conley, J. R. and Pampanin, S. (1999). Preliminary results and conclusions from the PRESSS five-story precast concrete test building. PCI Journal Vol. 44 No. 6: 42-67 Restrepo, J. I. and Rahman, A. (2007). Seismic performance of self-centering structural walls incorporating energy dissipators. Journal of Structural Engineering Vol. 133 No. 11: 1560-1570 Sritharan, S., Aaleti, S., Henry, R. S., Liu, K. Y. and Tsai, K. C. (2008). Introduction to PreWEC ad key results of a proof of concept test. M.J. Nigel Priestley Symposium, North Lake Tahoe, California, August 4-5, 2008 M. J. N. Priestley, M. J. Kowalsky and S. Sritharan. Pavia, Italy, IUSS Press: xiv, 267 p.
