Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska 10NCEE

GROUTED SPLICE SLEEVE CONNECTION ALTERNATIVES FOR PRECAST

REINFORCED CONCRETE BRIDGE PIERS IN MODERATE-TO-HIGH SEISMIC

REGIONS

M.J. Ameli1, J.E. Parks1, D.N. Brown1, and C.P. Pantelides2

ABSTRACT

Connections between precast concrete bridge elements in the substructure are some of the most critical components in bridges constructed using accelerated bridge construction. Researchers are in the process of investigating the suitability of various connection configurations in moderate-to-high seismic regions. Load capacity, ductility level, and reparability are three significant acceptance criteria for any connection considered in earthquake-prone regions. The seismic performance of four half-scale bridge components connected by means of Grouted Splice Sleeves (GSS) was studied. Cyclic quasi-static loading was applied to two column-to-footing and two column-to-cap beam connections. The column-to-footing connections incorporated one type of GSS where the bars were grouted at both ends (GGSS); the column-to-cap beam connections used a different GSS type where one bar was threaded into one end and the other bar was grouted into the opposite end (FGSS). Splice sleeves were located in the column ends for each connection category in the first phase of the research. In the second phase, the splice sleeves were located inside the footing and inside the cap beam, in order to investigate the effect of the splice sleeve location in the overall response of the test specimens under seismic loading. Experimental results showed that the performance of all four test specimens was satisfactory and the connections were viable. Improved seismic response was observed when the sleeves were located inside the footing and the cap beam rather than the corresponding column end. 1Graduate Student Researcher, Dept. of Civil and Environmental Engineering, University of Utah, UT 84112 2Professor, Dept. of Civil and Environmental Engineering, University of Utah, UT 84112

Email: c.pantelides@utah.edu Ameli M.J., Parks J.E., Brown D.N., Pantelides C.P. Grouted Splice Sleeve Connection Alternatives for Precast Reinforced Concrete Bridge Piers in Moderate-to-High Seismic Regions. Proceedings of the 10th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska 10NCEE

Grouted Splice Sleeve Connection Alternatives for Precast Reinforced

Concrete Bridge Piers in Moderate-to-High Seismic Regions

M.J. Ameli1, J.E. Parks1, D.N. Brown1, and C.P. Pantelides2

ABSTRACT Connections between precast concrete bridge elements in the substructure are some of the most

critical components in bridges constructed using accelerated bridge construction. Researchers are in the process of investigating the suitability of various connection configurations in moderate-to-high seismic regions. Load capacity, ductility level, and reparability are three significant acceptance criteria for any connection considered in earthquake-prone regions. The seismic performance of four half-scale bridge components connected by means of Grouted Splice Sleeves (GSS) was studied. Cyclic quasi-static loading was applied to two column-to-footing and two column-to-cap beam connections. The column-to-footing connections incorporated one type of GSS where the bars were grouted at both ends (GGSS); the column-to-cap beam connections used a different GSS type where one bar was threaded into one end and the other bar was grouted into the opposite end (FGSS). Splice sleeves were located in the column ends for each connection category in the first phase of the research. In the second phase, the splice sleeves were located inside the footing and inside the cap beam, in order to investigate the effect of the splice sleeve location in the overall response of the test specimens under seismic loading. Experimental results showed that the performance of all four test specimens was satisfactory and the connections were viable. Improved seismic response was observed when the sleeves were located inside the footing and the cap beam rather than the corresponding column end.

Introduction Recent advancements in bridge construction include innovative methodologies that bring about ease of construction and acceleration of the overall project delivery time. Prefabrication of bridge elements contributes to this construction method and facilitates the whole construction process, whether the bridge is new or a replacement.

Connections between such elements in the bridge substructure are some of the most critical components in bridges constructed using accelerated bridge construction. Researchers are in the process of investigating the suitability of various connection configurations in moderate-to-high seismic regions. Such connections undergo high levels of earthquake-induced

1Graduate Student Researcher, Dept. of Civil and Environmental Engineering, University of Utah, UT 84112 2 Professor, Dept. of Civil and Environmental Engineering, University of Utah, UT 84112

Email: c.pantelides@utah.edu Ameli M.J., Parks J.E., Brown D.N., Pantelides C.P. Grouted Splice Sleeve Connection Alternatives for Precast Reinforced Concrete Bridge Piers in Moderate-to-High Seismic Regions. Proceedings of the 10th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

deformations and stresses, while concentrated damage accumulates in a limited and localized area. Load capacity, ductility level, and reparability are three significant acceptance criteria for any connection of this type, in earthquake-prone regions. Strength properties of individual Grouted Splice Sleeve (GSS) assemblies were studied in the past [1]. Tension test results provided in-depth information on the potential failure modes, ultimate load capacity, and flow of the tensile load. To date, only a few large-scale experimental studies have been conducted utilizing GSS connections. Aida et al. (2005) and Yoshino et al. (1996) reported on the cyclic performance of large-scale specimens representing a bridge and a building subassembly, respectively [2, 3]. One particular type of GSS was incorporated in these specimens to connect the precast components together. Test results indicated acceptable response to cyclic loads when compared to cast-in-place control specimens. Haber et al. (2013) discussed the results of three specimens, two of which had GSS connections while the third specimen was cast monolithically to serve as a control specimen [4]. All specimens exhibited similar performance in terms of ultimate load capacity and energy dissipation, but not ductility capacity. NCHRP report 698 evaluated several connection types applicable to Accelerated Bridge Construction (ABC), specifically for moderate-to-high seismic regions [5]. The GSS connection was classified as a practical and promising connection type in that report, requiring more research on both strength and displacement properties. In particular, the inelastic behavior of such connections under cyclic loads and the sensitivity of the response to the location of the sleeves were highlighted as issues in need of further research.

As part of an extensive research program, four half-scale specimens were tested under a

cyclic quasi-static displacement history. Two column-to-footing specimens incorporated one type of GSS where the bars were grouted at both ends (GGSS), and two column-to-cap beam specimens used a different GSS type where one bar was threaded into one end of the sleeve and the other bar was grouted into the opposite end (FGSS).

Experimental Program

Half-scale models were designed and detailed in accordance with the principles of capacity based design, to simulate typical prototype bridges built in the State of Utah. The octagonal columns were built using No. 8 longitudinal bars in a circular arrangement and confined with a No. 4 spiral. The longitudinal and transverse reinforcement ratio was 1.3% and 1.9%, respectively. The footings and cap beams were reinforced using No. 8 longitudinal bars and No. 4 double hoops; they were designed to remain elastic as capacity-protected members. Test Matrix The design process followed the AASHTO LRFD Bridge Design Specifications and the AASHTO Guide Specifications for LRFD Seismic Bridge Design [6, 7]. The latter prohibits splicing of the column longitudinal rebar inside the plastic hinge region for bridges in high-seismic zones. To address this issue, the response of the bridge subassemblies was studied by changing the location of the splice sleeves. Fig. 1 shows the specimen details. Splice sleeves were placed inside the column ends for specimens GGSS-1 and FGSS-1, whereas for specimens GGSS-2 and FGSS-2, they were placed outside the column plastic hinge region and inside the footing and cap beam, respectively. In Fig. 1, the width of the footing for specimens GGSS-1

Figure 1. Details of Test Specimens.

and GGSS-2 was 3 ft, whereas the width of the cap beam for specimens FGSS-1 and FGSS-2 was 2 ft. Table 1 lists the test matrix and the compressive strength of the concrete and grout used in construction of the test specimens, and tested in accordance with ASTM standards. Test Procedure A cyclic quasi-static lateral displacement history comprised of ascending amplitudes of the predicted column yield displacement was applied as illustrated in Fig. 2. Two cycles were applied for each displacement [8]. Specimens FGSS-1 and FGSS-2 were tested upside down. An axial load 6% of the column axial capacity was applied by means of a hydraulic actuator resting on the column using post tensioning rods. Linear variable differential transformers mounted on opposite sides of the columns measured vertical displacements near the column end. String potentiometers captured the lateral displacement of the column top. The specimens were instrumented with several strain gages especially in the plastic hinge region and in the joint area.

Test Results

The hysteresis loops of the four specimens demonstrated a typical hysteretic behavior of precast components in which the bond-slip effects are pronounced [9]. Fig. 3 shows the lateral force-displacement curves in addition to the major damage states, which were: end of crack formation and initiation of spalling, yield penetration, rebar fracture, and rebar pull out from sleeves.

Figure 2. Test setup for GGSS-1 and GGSS-2.

GGSS-1 and GGSS-2 exhibited full and stable loops during all cycles as opposed to FGSS-1 and FGSS-2, which experienced some pinching, associated with bond deterioration inside the sleeves of these specimens during the last few cycles. Observations and Damage States There were similarities and differences between the test specimens due to the sleeve type and sleeve location. Ductile failure was observed for GGSS-1 and GGSS-2, while FGSS-1 and FGSS-2 failed due to bond deterioration and consequent rebar pullout. Only flexural cracks developed in specimens with splice sleeves in the column end, GGSS-1 and FGSS-1, whereas both flexural and inclined cracks formed in specimens incorporating sleeves in the footing and the cap beam, i.e. GGSS-2 and FGSS-2. For GGSS-1 and FGSS-1, the first sign of cracking was observed at the column to footing interface or the column to cap beam interface, where a ¼ -in. bed grout was cast at the time of assembling the precast components; this was followed by another crack around the top of the sleeves. For GGSS-2 and FGSS-2, the first crack developed 1 ft above the column end, before the interface bed grout formation. All major flexural cracks formed by the end of the 3% drift ratio, in addition to initiation of considerable concrete spalling. Fig. 4(a) presents the state of damage at the end of the 3% drift ratio.

Cracks opened further and concrete spalling intensified after the 3% drift ratio up to the last drift ratio. Yield penetration occurred during the 6% drift ratio for GGSS-1 with a depth of 1 in. around the two extreme column bars. The test was terminated at the end of the 9% drift ratio when both of these bars fractured due to low cycle fatigue. The test for FGSS-1 was terminated at the end of the 6% drift ratio because the column bars pulled out of the sleeves, as a result of

Table 1. Test Matrix.

Specimen Connection Type Sleeve Type Sleeve

Location

Compressive Strength (ksi) Concrete Grout

28-day

Test day

28-day

Test day

GGSS-1 Column-Footing Grouted-Grouted Column end 5.3 5.9 14.4 14.4 FGSS-1 Column-Cap beam Fastened-Grouted Column end 5.3 6.2 12.5 13.3 GGSS-2 Column-Footing Grouted-Grouted Footing 3.9 5.5 11.1 13.5 FGSS-2 Column-Cap beam Fastened-Grouted Cap beam 3.9 5.2 10.3 10.3

bond failure associated with the grout crushing in the vicinity of the rebar deformations, inside the sleeves. A few inclined cracks developed during the 4% and 5% drift ratios for GGSS-2. This specimen failed when the east column rebar fractured during the first cycle of the 7% drift ratio. The fracture location was found to be 2 in. above the column end. A combined failure mode was ascertained for FGSS-2. At the 7% drift ratio, the west column rebar fractured 1 in. above the column end, during the first push. Another drop in the load capacity was observed during the subsequent pull caused by the east column bar pulling out from the sleeve. Fig. 4(b) shows the final damage state for all test specimens. Ductility and Energy Dissipation

The displacement ductility of the specimens was obtained based on the concept of equal energy of an idealized elasto-plastic system [9]. The average backbone curve was first constructed using the peak values of the first cycle for each drift ratio. The idealized elasto- plastic curve was then generated in order to calculate the displacement ductility. The ultimate displacement was taken as the displacement corresponding to a 20% or larger drop in the lateral load capacity [10]. For the specimens with sleeves in the columns, GGSS-1 and FGSS-1, the displacement ductility was

Figure 3. Force-Displacement response with damage states.

GGSS-1 FGSS-1 GGSS-2 FGSS-2

(a) Damage State at 3% drift ratio.

GGSS-1 FGSS-1 GGSS-2 FGSS-2

(b) Final damage state. Figure 4. Damage state; crack pattern and spalling.

found to be 5.9 and 4.4, respectively; however, for specimens with sleeves in the footing and cap beam, GGSS-2 and FGSS-2, the displacement ductility was 6.1 and 5.8, respectively—a considerable improvement. It is noted that the ductility levels achieved by all four specimens satisfy the Caltrans requirements [11]. The area enclosed by the hysteresis loops is referred to as the hysteretic energy of a system. Fig. 5(a) compares the hysteretic energy per drift level for all specimens. This is an indication of the quality of the hysteretic response, and the load-displacement characteristics. It is noted that all specimens dissipated nearly the same amount of energy up to a 3% drift ratio. Specimens GGSS-1 and GGSS-2 showed better hysteretic energy dissipation capabilities during the subsequent drift ratios. Equivalent viscous damping is another quantity used to evaluate the relative energy dissipation capacity of systems under cyclic loads. Equivalent viscous damping offers more information about the hysteretic response of the system since both hysteretic and strain energy are included in this method. The equivalent viscous damping ratio (ξeq) was obtained as the ratio of the hysteretic energy to the energy of the equivalent viscous system as defined in Eq.1 [12]:

ξ E

4πE (1)

where ED and ES0 are the area inside the hysteresis loop and the strain energy, respectively. Fig. 5(b) presents the average ξeq of both cycles for each drift ratio. In the inelastic region of the response, which begins after completion of the 1% drift ratio, ξeq for all specimens increases with an increase in the drift ratio. At 6% drift ratio, GGSS-2 with ξeq of 25% had a superior dissipation capacity when compared to the other test specimens; on the contrary, GGSS-1 and FGSS-1 had a lower ξeq of 15%, implying a lower energy dissipation capacity when the sleeves were located in the column.

(a) Dissipated energy (b) Equivalent Viscous Damping

Figure 5. Energy dissipation capacity. Column Curvature Linear variable differential transformers (LVDTs) installed on both extreme sides of the column end, were used to study the curvature distribution and curvature capacity of the specimens. Therefore, four curvature segments were specified by using four LVDTs on each side. The average curvature was computed as shown in Eq. 2:

φ A Bwh

(2)

where A and B are LVDT readings, and w and h are the segments width and height, respectively. Fig. 6 plots the average curvature profile over a 30 in. column height above the column end. Positive curvature values were associated with the push direction and negative values with the pull direction. The calculated curvature value was assumed to be an average over the whole segment height. Curvature values are included up to a 6% drift ratio, which was the last common drift ratio among all specimens. Dashed lines mark the top of the splice sleeves in the column end for GGSS-1 and FGSS-1. Lack of curvature was evident in these two specimens due mainly to the presence of the splice sleeves, with pronounced localized demand around the interface. This indicates a condition of concentrated inelasticity for the case of specimens with sleeves in the column end. By contrast, GGSS-2 and FGSS-2, with sleeves in the footing and cap beam, respectively, had a well-distributed curvature profile, although FGSS-2 had an asymmetric profile. GGSS-2 had a superior curvature distribution and its maximum curvature value at 6% drift ratio was 3.3 times larger than that of GGSS-1. Stiffness Degradation

The effective stiffness was calculated in each cycle using the peak displacement values and corresponding forces. The average of the stiffness values was then obtained for both cycles of every drift ratio. Fig. 7 displays the average effective stiffness at each drift level for all specimens. A similar trend was noted in the stiffness reduction per drift ratio for all specimens. The degradation rate was much higher during the first few cycles, mainly because of column rebar yielding. For example, there was a 60% reduction in GGSS-2 stiffness by the end of the 2% drift ratio. It is evident that GGSS-1 and FGSS-1 had a relatively higher stiffness than the other two, at every drift ratio.

Figure 6. Average curvature profiles.

Column Base Rotation Similar to the procedure established for column curvature, the rotation at the column base was computed. Two LVDTs were used to obtain the column base rotation. They were mounted on top of the footing and cap beam, and fastened to a rod extending out of the column at a section located 7 in. above the column end. The LVDTs provided valid readings for all specimens, except the east LVDT for FGSS-2 which had shifted as a result of severe damage during the test. Therefore, readings from the west side were used assuming a symmetric performance for FGSS-2.

It is evident that this measurement includes the effects of both rebar slip and flexural action in the bottom segment of the column. Assuming an approximately similar flexural behavior for specimens with the same splice sleeve location, bond-slip characteristics of the specimens could be compared. Fig. 8 displays the moment-rotation for the specimens at the interface, for drift ratios up to 6%. The results indicate that specimens had distinct rotational responses, especially when comparing GGSS-1 and FGSS-1, where the sleeves were accommodated in the column end. FGSS-1 showed an asymmetric performance with about 10% difference in the rotation corresponding to the push and pull directions during the 6% drift ratio. This could be attributed to greater rebar slip occurring in the longitudinal column bars of the east side rather than those of the west side. FGSS-1 had 10% and 28% larger rotations in the push and pull directions, respectively, compared to GGSS-1 during the last drift ratio.

GGSS-2 and FGSS-2 were similar in terms of the moment-rotation relationship. This indicates that bond-slip behavior was not as effective for these specimens as it was for the case that the splice sleeves were in the column end.

Figure 7. Stiffness degradation.

Conclusions

All four specimens performed well, with displacement ductility levels that satisfy the Caltrans requirements. In spite of more effort during the construction phase, specimens with the sleeves inside the footing or the cap beam achieved a higher curvature capacity and displacement ductility than the corresponding specimens with the sleeves in the column.

Column-to-footing specimens, with sleeves in which both bars were grouted at both ends (GGSS) had higher hysteretic energy compared with column-to-cap beam specimens in which one bar was fastened to one end and the other was grouted in the opposite end (FGSS). The equivalent viscous damping ratio for all specimens at 3% drift ratio was above 10%, and at 6% drift ratio was above 15%. The highest damping ratio was 25% for the column to footing specimen with GGSS in the footing (GGSS-2).

Figure 8. Moment-rotation at column base.

Stiffness degradation was similar for all specimens regardless of the type and location of the splice sleeves. More flexural cracks and larger crack widths, distributed damage, and a ductile failure for GGSS-2 and FGSS-2 implied a distributed plasticity and a more desirable performance, as opposed to the concentrated damage occurred in GGSS-1 and FGSS-1. The curvature profiles also affirmed the distinct plasticity distribution in the column ends. The curvature demand was minimal over the sleeve region for GGSS-1 and FGSS-1, with the sleeves located inside the columns. The performance of FGSS-1 was affected by the bond-slip behavior of the bars in the corresponding splice sleeves, more than the other specimens. This issue was partially mitigated when the sleeves were placed inside the cap beam, i.e. FGSS-2.

The experimental results discussed in this paper will be compared to the experimental results for two monolithic cast-in-place specimens, for each category, in the upcoming phases.

Acknowledgements

The authors would like to acknowledge the financial support of the Utah Department of Transportation, the New York State Department of Transportation and the Texas Department of Transportation. The authors wish to thank Professor Lawrence D. Reaveley, University of Utah, for his invaluable input throughout the project. Special thanks are extended to Mark Bryant, University of Utah, for his untiring efforts and support. The authors are also grateful for donation of materials by Splice Sleeve North America and ERICO.

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