EXPERIMENTAL TESTING OF THE NEXT GENERATION OF LOW DAMAGE ROCKING BRIDGE PIERS

LIU R. & PALERMO, A.

Department of Civil and Natural Resources Engineering, University of Canterbury, New Zealand

SUMMARY Single column bents also known as cantilever or hammerhead piers are a common structural configuration of substructure for highway bridges in New Zealand. In terms of seismic design, bridges using this type of substructure rely heavily on the robustness of the plastic hinge (typically designed to form at the column base) and redundancy in the number or longitudinal bars to prevent either local failure or collapse of the entire structure. A similar type argument can be made for low damage solutions such as the Hybrid PRESSS system also known as dissipative controlled rocking (DCR). Current research on DCR has focussed on performance with no work directed towards the issue of seismic structural redundancy and robustness. In this contribution, the issue of seismic structural redundancy and robustness for DCR is discussed in addition to a description given of experimental work being conducted at the University of Canterbury on a 2/3 scale low damage pier subassembly designed to test some ideas conjectured to improve the seismic structural redundancy and robustness of DCR. INTRODUCTION The design technology known as the Hybrid PRESSS system (Priestley, 1996) or dissipative controlled rocking (DCR), is the combination of member end rocking, unbonded post-tensioning, and dissipative devices (across the rocking interface). Focussing on its application to bridges, DCR has been shown to be superior in terms of seismic performance both in numerical simulation and laboratory testing (Guerrini, Asce, Restrepo, Massari, & Vervelidis, 2015; Marriott, 2009; Moustafa & ElGawady, 2016; Palermo, Pampanin, & Calvi, 2005). An aspect of this design strategy which up till now has not been explored is seismic structural redundancy. THE ISSUE OF SEISMIC STRUCTURAL REDUNDANCY AND ROBUSTNESS Structural redundancy in the general sense refers to how many extra parts (be it members, supports, or number of components in each connection) a structure has over the minimum number required to function in case of unexpected failure of a component to prevent collapse. The term “seismic structural redundancy” has a similar definition but where it refers to the amount of redundancy a structure has in case of unexpected/premature failure of a part(s) due to seismic loading. For bridges, the seismic structural redundancy depends largely on the choice of substructure (pier and foundation) and hinging mechanism (Figure 1). A multi-column

bent is more seismically structurally redundant than a cantilever pier as there are more plastic hinges which can be relied upon to resist seismic action. Another term which will be used in this paper is “structural robustness”. Where, in this context the use of this term refers to structural design details which reduce the likelihood of local and or global failure of structure under a given load case. As an example of this concept for seismic loading, the structural robustness of a plastic hinge in a traditional reinforced concrete structure is dependent on the adequacy of the reinforcement detailing in the plastic hinge zones. If the transverse reinforcement is spaced too far apart in the plastic hinge zone then the entire structure is not likely to be seismically structurally robust due to the high likelihood of premature buckling and fracture of the longitudinal bars.

Figure 1. Examples of hinging mechanisms with the highest allowed ductility limits (New Zealand Transport Agency, 2013). Leftmost and 3rd from the left show hinging mechanism under transverse seismic loading for a cantilever pier and two column bent respectively.

Now returning to the main issue: simply supported beam bridges with cantilever hammerhead piers (Figure 1 left and second from left), rely heavily on hinge detailing for robustness and on the large number of longitudinal bars for redundancy to prevent premature failure of the hinge zone and collapse of the structure. For low damage design strategies, such as dissipative controlled rocking (DCR) applied to cantilever hammerhead piers, a similar sort of reasoning regarding structural redundancy and robustness can be applied, but, with the source of redundancy against overturning being different to that of a traditional reinforced concrete solution. For DCR, structural redundancy under seismic loading depends largely on the restoring moment provided by the post-tensioning and multiplicity in the number of energy dissipating devices at the base of the pier. The DCR design strategy then derives its robustness from the fatigue life of the dissipative devices (important for long duration and sequential earthquake events) and the displacement capacity of the central post-tensioning prior to yielding. Due to DCR being highly dependent on these two systems for seismic structural redundancy and robustness, this had led the authors to investigate the issue of improving the structural robustness of DCR in what is termed the next generation of low damage rocking bridge piers. IMPROVING SEISMIC STRUCTURAL REDUNDANCY OF DISSIPATIVE CONTROLLED ROCKING The authors propose three different solutions for improving the seismic structural redundancy of DCR: 1) have two sets of dissipative devices across one rocking interface but where the second set is only activated once the rocking rotation exceeds the design limit (Liu & Palermo, 2016b); 2) spread the rotational gap opening demand over multiple joints and have dissipative devices across each of those joints (Figure 2a); 3) combine DCR with foundation rocking to make use of rocking isolation under seismic loading greater than the design seismic load (Figure 2b). All the solutions are more robust than conventional design since they consider the activation of multi-layer dissipators and/or rocking sections along the pier. In this contribution, the specimen being described makes use of the latter two solutions.

a) b)

Figure 2: a) Segmental column DCR & b) Combined DCR and Pile cap rocking. In a) the dissipative devices across the higher joint behave inelastically for loadings greater than the

design load, while, in b) pile cap rocking is only activated for loadings greater than the design load.

SPECIMEN DESCRIPTION The specimen is a fully precast, cantilever, low damage pier with “pile cap” (Figure 3a). It is 2/3 scale of a pier which would be used for a typical two lane New Zealand highway bridge (Liu & Palermo, 2016a), but, can be considered 1:1 scale for a single lane New Zealand highway bridge. Overall the specimen is 6300mm tall and has a footprint of 4030mm x 3970mm. The column is octagonal in section (Figure 3b), 1m in diameter, has an overall clear length of 3400mm, and a cantilever length of 4000mm (column base to point of lateral loading). The entire structure is made of eight parts (Figure 6): two ground beams, a pile cap, two concealing blocks, two column segments, and a loading beam.

a) b)

Figure 3: a) South elevation of the test specimen. b) Close up of the two column segments and dissipative devices.

Structural configurations As previously mentioned, this specimen has three structural configurations in which it can be tested: segmented column DCR, conventional DCR, and DCR combined with pile cap rocking. To be able switch between configurations, two joints can be locked or unlocked: the pile-cap to ground beam interface, by loosening or tightening the four hold down anchors either side of the pile cap; and the joint between the two column segments, by not grouting all the starter bars coming out of the bottom column segment.

Figure 4: Depiction of the three different structural configurations in which the specimen will be tested in.

Design The dimensions, design strength, and design displacement were determined by scaling off the same prototype structure used by (Liu & Palermo, 2016a). The seismic action design parameters are also the same as that used by (Liu & Palermo, 2016a). The strut and tie design method as according to NZS 3101: 2006 (Standards New Zealand, 2006) was used to design most of the reinforcing within the structure including the post-tensioning anchorage zones. The design of the unbonded reinforcement crossing the joint between column segments was determined based off numerical multi-section moment-rotation-axial load analysis developed by the authors based on work by (Ou, 2007). The sizing of the plan dimensions of the pile cap was determined also from multi-section moment-rotation analysis. Construction Construction of the reinforced concrete components of the specimen was undertaken by Bradford Precast Ltd, Ashburton. Due to the size and complexity (number of concrete components, reinforcing density, alignment of common ducts, etc.) of the specimen Building Information Modelling (BIM) was used to ensure the constructability of the specimen on top of: close collaboration between Bradford Precast and UC; the use of CNC fabricated steelwork; and the use of alignment fixtures to position reinforcing bars and ducts where possible. Assembly of the specimen in the Structural Engineering Laboratory at the University of Canterbury consisted of the following construction stages: Levelling and alignment of ground beams; placement of the pile cap on the ground beams and anchorage to the strong floor; placement and temporary anchorage of the first column segment to the pile cap by eight M36 grade 8.8 threaded rods; installation of the first half of the central post-tensioning bar; installation of the concealing blocks with column base shear keys on the pile cap; installation of the second column segment which was temporarily anchored to the pile cap by internal debonded Freyssibars; installation of the second half of the central post-tensioning bar; installation of the cap beam with the gravity loading system; attachment of the lateral loading ram to the cap beam; stressing the central post-tensioning bar; and grouting four of the 12 starter bars crossing the joint between the two column segments.

a) b) c)

Figure 5: a) BIM model of specimen, oblique view. b) Partially exploded BIM model of specimen showing modelling of reinforcement and ducts, oblique view. c) BIM model section

through column, concealing blocks and pile cap showing internal reinforcement.

TESTING ARRANGEMENT & VARIABLES EXPLORED Two types of loads are simulated in the testing arrangement: gravity loading from the scaled bridge superstructure and lateral seismic loading. The simulated gravity loading is provided by two vertical 1000kN hydraulic rams which are anchored to the strong floor either side of the specimen. Each gravity loading ram applies 500kN of compression force to the specimen via the loading beam. The gravity load is actively controlled and kept constant as a function of lateral displacement of the specimen. Due to the gravity loads not being provided by weight force from mass at the top of the specimen, P-Δ effects are not simulated in this experiment. The lateral seismic loading is unidirectional, cyclic, and applied quasi-statically. It is provided by a single horizontal 1000kN hydraulic ram connecting the specimen to the laboratory strong wall.

Figure 6: Front and side view of the test set up. Major parts of the specimen and set up annotated.

The loading protocol used was derived from ACI T1.1-01 (ACI Innovation Task Group 1, 2001) where: three fully reversed cycles are applied at each drift ratio (ratio of deck displacement to pier length); the first drift ratio is within the linear elastic response range; and subsequent drift ratios are between 1.25 and 1.5 times the previous value. The ULS drift ratio for the specimen is 1.7% (Δ = 68mm) whilst the MCE drift ratio is 2.7% (Δ = 108mm). The maximum displacement which can be applied symmetrically to the specimen is 4.37% drift (Δ = 175mm), where, this value dictated by the stroke available from the lateral loading ram.

Figure 7: Loading protocol displacement history.

Test Variables In the experimental program for this specimen, the variables which are controlled will be: number of column base dissipative devices, joints which will be allowed to open, and post-tensioning force. Throughout all tests the gravity load is kept constant at 1MN. Table 1 below gives a summary of all the tests which will be performed on the specimen.

Table 1: Summary of the tests which will be undertaken in the experimental program. Fpu is the characteristic failing load of the post-tensioning being referred to.

Test number and configuration description Pier post-tensioning kN

Pier base dissipators

no.

(T1-SegDCR): Segmented PRESSS/DCR pier 340 (0.17Fpu) 8

(T2-SegPT): Segmented post-tensioned rocking pier 290 (0.14Fpu) -

(T3-SegPT): Segmented post-tensioned rocking pier 690 (0.34Fpu) -

(T4-SegPT): Segmented post-tensioned rocking pier 526 (0.26Fpu) -

(T5-DCR): PRESSS/DCR pier 340 (0.17Fpu) 8

(T6-PT): Post-tensioned rocking pier 340 (0.17Fpu) -

(T7-PT): Post-tensioned rocking pier 690 (0.34Fpu) -

(T8-PT): Post-tensioned rocking pier 526 (0.26Fpu) -

(T9-PCDCR): Pile cap rocking + DCR pier 400 (0.20Fpu) 8

EXPECTED OUTCOMES FROM EXPERIMENT This experiment is being conducted not only as a means of obtaining experimental evidence to: validate the theoretical predictions used to model and design the structure; and validate numerical modelling for future parametric analysis. But, also to investigate the effect of scale on the practicalities of design and construction of DCR bridge piers. Other than the Wigram Magdala Place bridge in Christchurch (Routledge, Cowan, & Palermo, 2016), the specimen described in this paper is at the time of writing the largest example of a DCR bridge pier.

0 1000 2000 3000 4000 5000 6000-5

-4

-3

-2

-1

0

1

2

3

4

5D

rift

,

%

Displacement step

0 1000 2000 3000 4000 5000 6000

-180

-140

-100

-60

-20

20

60

100

140

180

Dis

pla

cem

ent,

mm

At the time of writing of this paper, Test 1 (T1-SegDCR) had been completed and presented in figures 8 and 9 are respectively the global force-displacement behaviour and measured gap opening at the column base and joint between column segments on the east face of the column. Overlaid on both figures are the predictions derived from modelling conducted prior to construction and the limit states of the structure. In terms of preliminary analysis of these results: the force-displacement prediction seems accurate for large drifts and grossly overestimates the initial stiffness of the system (the authors believe this to be due to imperfections at the rocking interfaces; the force-displacement is asymmetric due to two dissipators on one side of the column pulling out from there bottom attachment because they were not fully screwed to the bottom attachment leading to reduced strength during the positive phase of displacement; and the prediction made regarding the quantification of the size of gap opening which will happen at both joint seems relatively accurate however, the data suggests more a bilinear relationship in gap opening which will be investigated further.

Figure 8: Force displacement (in terms of drift) from test 1 compared against prediction.

Figure 9: Comparison of predicted gap opening at the column base and joint between segments with data from test 1.

CONCLUSION In conclusion, the issue of seismic structural redundancy and robustness was discussed for the design strategy of dissipative controlled rocking. Three solutions to improve the seismic structural redundancy of DCR were stated. A description of experimental work being undertaken at the University of Canterbury on a 2/3 scale low damage pier and pile cap utilising two of the methods to improve the seismic structural redundancy of DCR (segmental DCR and DCR combined with pile cap rocking) was given. ACKNOWLEDGEMENTS Research funders: UC Quake Centre and Natural Hazards Research Platform; Sponsorship and specimen manufacture: Bradford Precast Ltd; Contributions to experimental work: UC technical staff Gavin Keats, Mosese Fifita, and Peter Coursey. REFERENCES ACI Innovation Task Group 1. (2001). Acceptance criteria for moment frames based on

structural testing. Farmington Hills, Michigan, USA: American Concrete Institute. Guerrini, G., Asce, S. M., Restrepo, J. I., Massari, M., & Vervelidis, A. (2015). Seismic Behavior

of Posttensioned Self-Centering Precast Concrete Dual-Shell Steel Columns. Journal of Structural Engineering, 141(4), 1–11. https://doi.org/10.1061/(ASCE)ST.1943-541X.0001054.

Liu, R., & Palermo, A. (2016a). Large scale testing of a low damage substructure connection in a precast concrete bridge. In The New Zealand Concrete Industry Conference 2016. Auckland, N.Z.: The New Zealand Concrete Industry.

Liu, R., & Palermo, A. (2016b). Pier to deck interaction and robustness of PRESSS hybrid rocking : issues affecting hammerhead pier bridges. In 2016 NZSEE Conference (pp. 1–9). Christchurch, N.Z.

Marriott, D. (2009). The Development of High-Performance Post-Tensioned Rocking Systems for the Seismic Design of Structures. University of Canterbury.

Moustafa, A., & ElGawady, M. A. (2016). Damage-resistant segmental double-skin bridge column with replaceable energy dissipaters. In The sixth International Conference on Structural Engineering, Mechanics and Computation. Cape Town, South Africa.

New Zealand Transport Agency. (2013). Bridge manual 3rd Edition (3rd ed.). Ou, Y. (2007). Precast segmental post-tensioned concrete bridge columns for seismic regions.

University of New York at Buffalo. Palermo, A., Pampanin, S., & Calvi, G. M. (2005). Concept and development of hybrid

solutions for seismic resistant bridge systems. Journal of Earthquake Engineering, 9(6), 899–921.

Priestley, M. J. N. (1996). The PRESSS Program Current Status and Proposed Plans for Phase III. PCI Journal, 41(2), 22–40.

Routledge, P. J., Cowan, M. J., & Palermo, A. (2016). Low-damage detailing for bridges – A case study of Wigram-Magdala Bridge. In 2016 NZSEE Conference (pp. 1–8). Christchurch, N.Z.

Standards New Zealand. NZS3101 Concrete structures standard: Part 1 - the design of concrete structures (2006). New Zealand.