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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Page 24
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Page 25
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Page 26
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Page 27
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
Page 27
Page 28
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Page 29
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Page 45
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
×
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2011. Development of a Precast Bent Cap System for Seismic Regions. Washington, DC: The National Academies Press. doi: 10.17226/14484.
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Below is the uncorrected machine-read text of this chapter, intended to provide our own search engines and external engines with highly rich, chapter-representative searchable text of each book. Because it is UNCORRECTED material, please consider the following text as a useful but insufficient proxy for the authoritative book pages.

92.1 Introduction This chapter provides a synopsis of key findings of the experimental and analytical research program. This research program investigated the following: • Emulative, nonintegral details – Grouted duct connection – Cap pocket connection • Hybrid, nonintegral details – Conventional system – Concrete filled pipe system – Dual steel shell system • Integral detail – Post-tensioned girder system detail Findings include a description of the experimental test program—including specimen design, fabrication, and test- ing protocol—and response of the specimens. Detailed find- ings can be found in the attachments to this report, available online at www.trb.org/Main/Blurbs/164089.aspx. 2.2 Description of Experimental Test Program This section describes the general development of the test- ing program and associated experimental specimens. Detailed drawings are provided in the attachments to this report. 2.2.1 Design of Nonintegral Prototype Bridge and Specimens In coordination with the NCHRP Project 12-74 panel, the prototype structure was selected as a two-span, nonintegral bridge with a three-column CIP bent cap supporting precast, prestressed girders, intended to represent a typical highway overcrossing located in an urban area. Design of the compo- nent specimens is based on a representative portion of the cen- ter column and bent cap of the bridge, as shown in Figure 2.1. The full design of the prototype and specimen are reported in Matsumoto et al. (15). This section summarizes key design features of the CIP prototype bridge and CIP specimen. The CIP prototype bridge and full ductility component specimens were designed in accordance with AASHTO LRFD Bridge Design Specifications (Third Edition with 2006 Interims) (2006 LRFD BDS) and Recommended LRFD Guidelines for the Seismic Design of Highway Bridges (2006) (2006 LRFD RSGS) prepared as part of NCHRP Project 20-07/Task 193 and provided to the research team (16, 2). It is important to note that the 2006 LRFD RSGS was superseded by the Proposed AASHTO Guide Specifications for LRFD Seismic Bridge Design (2007 LRFD PSGS) and later updated to the current 2009 AASHTO Guide Specifications for LRFD Seismic Bridge Design (2009 LRFD SGS) (17, 1). In addition, the 2006 LRFD RSGS contains different—and in some aspects more liberal (i.e., less conservative)—joint reinforcement requirements than the current 2009 LRFD SGS. For example, in contrast the 2006 LRFD RSGS, the 2009 LRFD SGS specifies vertical joint stirrups both inside and outside the joint region, a larger total area of joint stirrups, and a significant increase in bent cap longitudinal reinforcement. These provisions are compared in Chapter 3. For a major seismic event, the CIP prototype bridge was designed and detailed to exhibit ductile plastic hinging in the column adjacent to the bent cap (and footing). Prototype bridge drawings are shown in the attachments. Initial member sizing was based on input from design engineers and refined through application of the 2006 LRFD BDS (16). Seismic analysis and design were performed to finalize column and cap beam sec- tions. The prototype structure was considered nonessential and designed for the associated life-safety performance objec- tives defined in the 2006 LRFD RSGS (2). Because the system is nonintegral, the specified earthquake resisting system in the longitudinal direction consisted of cantilever response of the C H A P T E R 2 Findings

10 columns with plastic hinge formation at the base of columns. One-way soil springs were used to account for the seismic resistance of the abutment backfill. Transverse earthquake resistance was provided by the three-column bent cap with plastic hinge formation occurring at both the top and bottom of the columns. The design acceleration response spectrum (ARS) curve incorporated 5% damping and was developed using a 1.0-sec acceleration of 0.8 g, a 0.2-sec acceleration of 1.5 g, and coefficients for Site Class D soil. The resulting peak rock acceleration was 0.6 g. The ARS curve is representative of a site located in a high seismic region such as Southern California. If the design earthquake response spectral acceler- ation coefficient at a 1.0-sec period, SD1, was larger than 0.50, the structure was classified as SDC D. This category required a demand analysis, displacement capacity pushover analysis, capacity design, and SDC D detailing. Elastic dynamic analysis was performed according to the 2006 LRFD RSGS to estimate seismic displacement demands (2). The columns were assumed to be fixed at the base, and foundation design was not performed. Moment-curvature analysis was conducted to estimate the effective stiffness of the columns. Seismic demands were determined using the SEISAB program (18). Results from the seismic analysis indicated a first mode (longitudinal) period of 1.27 sec and associated displace- ment demand of 15.1 in. The second mode (transverse) period was 0.58 sec, with a displacement demand of 5.3 in, including magnification to account for demand underestimation for shorter period structures. Shear keys at the abutments were designed to fail during the design seismic event. In the longitudinal direction, the displacement demand to capacity ratio, D/C, was 0.85. For determining transverse capacity, overturning effects were considered using an iterative procedure to refine the estimated column plastic moment capacities based on column axial loads obtained from the push- over analysis. XSection was used to perform moment-curvature analysis, and WFrame was used for pushover analysis, includ- ing overturning effects and bent cap flexibility (19, 20). The transverse displacement D/C ratio was 0.57. Ductility demand ratios were approximately 5.0 for both the longitudinal and transverse directions, well below the limit of 8.0 (multicolumn bent caps). The prototype structure thus satisfied the require- ments for displacement and ductility. P-delta effects were checked in accordance with the 2006 LRFD RSGS (2). Capacity protection design principles were also applied. Flexural and shear demands on the bent cap were based on force levels associated with the columns reaching their over- strength capacity. The bent cap design considered axial load effects due to transverse response in combination with grav- ity loads and overstrength demands imposed by columns. According to the 2006 LRFD RSGS, the bent cap was required to remain “nominally elastic” (2). Bent cap transverse rein- forcement outside of the joint region was designed according to the 2006 LRFD BDS (16) using modified compression field theory and considered seismic plus gravity loading. Column transverse reinforcement was designed to resist overstrength demands due to transverse response. The joint region of the bent cap was designed based on transverse response using the external force transfer mecha- Figure 2.1. Portion of prototype used for specimen design.

11 nism assumed in the 2006 LRFD RSGS (2). Joint design included vertical stirrups with horizontal cross ties in the region adjacent to the column (not within the joint), joint transverse reinforcement (hoops), extension of column bars close to the top of the bent cap, and side face reinforcement. In addition, two 2-leg construction stirrups were used within the joint region, as explained by Matsumoto (21). However, the prototype bridge did not incorporate the more conser- vative joint reinforcement requirements of the 2009 LRFD SGS (1), such as placement of the required area of stirrups inside the joint or the additional area of longitudinal bent cap reinforcement. The CIP test specimen was designed using a 42% scale of the central portion of the prototype bridge (see Figure 2.1). As the prototype bridge would be expected to exhibit ductile plastic hinging in the column region adjacent to the bent cap due to transverse response, the scaled CIP control specimen— loaded in the transverse direction under quasi-static force control and displacement control sequences—was expected to perform similarly. Dead load plus seismic load governed the bent cap flexural reinforcement in the prototype. This reinforcement was scaled for use in the bent cap. Bent cap transverse reinforcement was designed per the 2006 LRFD BDS (16). The additional joint shear reinforcement was required per 2006 LRFD RSGS, although a larger principal tensile stress was found for the specimen than the prototype due to the relatively smaller column load, larger cap and column tension, and imperfect scaling of dimensions (2). However, this was more desirable for examining joint behavior. Matsumoto (21) provides a detailed comparison of column, bent cap, and joint reinforcement for the prototype bridge and CIP specimen. Full Ductility Emulative Specimens The GD (grouted duct) and CPFD (cap pocket full ductil- ity) specimens used the same full-ductility design basis as the CIP specimen (22, 23). The GD and CPFD specimens were intended to be directly compared with the CIP control spec- imen. Design of both precast specimens assumed emulative response would be achieved despite the following differences between these specimens and the CIP specimen: (1) separate precast elements, including the bent cap and column, and (2) use of a 1.5-in bedding layer between the bent cap soffit and column to accommodate tolerances. In addition, the GD specimen used closely spaced, 1.75-in diameter, 22-gage corrugated ducts in the bent cap and high- strength, non-shrink, cementitious grout to anchor the col- umn longitudinal reinforcement. Joint reinforcement matched that used for the CIP specimen, including two 2-leg construc- tion stirrups within the joint region. CIP and GD specimen assembly details and bent cap reinforcement details are shown in the attachments. The CPFD specimen used a single 18-in nominal diameter, 16-gage steel pipe in the bent cap to house the column bars and serve as a stay-in-place form as well as equivalent joint hoop reinforcement. Normal-weight concrete was placed in the bent cap void and bedding layer to anchor the column bars. The CPFD bent cap reinforcement details are shown in the attachments. Matsumoto (23) summarizes the approach used for selecting the readily available lock seam, helical cor- rugated steel pipe per ASTM A760, Standard Specification for Corrugated Steel Pipe, Metallic-Coated for Sewers and Drains (24). Figure 2.2 shows the corrugation and lock seam details for the pipe used in the specimen, and select joint details are summarized in Table 2.1. Seam (a) Corrugation and Lock Seam (b) Close–up of Lock Seam Figure 2.2. Corrugation and lock seam details for cap pocket specimen.

12 Pipe thickness was a specific design parameter, calculated to provide the same nominal circumferential hoop force in the joint as that required for the CIP specimen per 2006 LRFD RSGS (2). Hoop force calculations assumed pipe nominal ten- sile yield strength of 30 ksi and used the horizontal component of the helical pipe. Subsequent tensile coupon tests conducted on the pipe material indicated a tensile yield strength of approx- imately 58 ksi. Nevertheless, calculations using the assumed 30 ksi yield strength resulted in a pipe thickness that matched the thinnest readily available pipe (16 gage). In addition, a #3 hoop, matching the column hoop size, was placed approxi- mately 1 in from each end of the pipe to reinforce the pipe and limit dilation and potential unraveling (see Figure 2.3). This was considered a reasonably simple yet conservative measure, given the limited number of specimen tests and unknown perform- ance of this innovative detail. Table 2.1 shows a hoop force ratio (pipe/hoop) of 1.03 when supplementary hoops are neglected, and 1.38 when accounting for the hoops. Joint reinforcement for the CPFD specimen did not include the horizontal J-bars used for the CIP specimen, although two 2-leg construction stirrups were placed within the joint region. Emulative Precast Bent Cap Connections for Seismic Regions: Component Test Report-Grouted Duct Specimen (Unit 2) (22) and Emulative Precast Bent Cap Connections for Seismic Regions: Component Tests-Cap Pocket Full Ductility Specimen (Unit 3) (23) provide detailed comparisons of col- umn, bent cap, and joint reinforcement for the prototype bridge and full ductility specimens. Limited Ductility Emulative Specimen The cap pocket limited ductility (CPLD) specimen (26) was intended to aid investigation of the response of a precast cap pocket connection designed according to the principles of limited ductility per the 2006 LRFD RSGS rather than the principles of full ductility used for the other specimens (2). For direct comparison, the CPLD specimen used the CPFD design as its initial basis. Emulative performance of the CPLD specimen was to be examined, especially through a displace- ment ductility of 2.0, even though a limited ductility CIP Table 2.1. Comparison of specimen details—CPLD versus CPFD. Item CPLD CPFD Notes Helical Pipe: (24, 25) Pipe Diameter (nom) Pipe Thickness (gage) Corrugation Angle Corrugation Dimensions Lock Seam Strength Steel Yield Strength 18 in 0.065 in (16) 20 deg 2.67 in × 0.50 in 240 lb/in 57.5 ksi 18 in 0.065 in (16) 20 deg 2.67 in × 0.50 in 240 lb/in 57.9 ksi Same basis allowed direct comparison of specimens Hoop Force Ratio: CPFD Pipe / Design (#3 hoops) 1.03 (no end hoops) 1.38 (extra end hoops) 1.03 (no end hoops) Potential benefit of end hoops eliminated for CPLD Vertical Stirrups, Horizontal Cross Ties None External to Joint Only (2006 LRFD RSGS) (2) No joint reinforcement used for CPLD Joint Other Reinforcement None Two 2-leg construction stirrups placed in joint Potential benefit of construction stirrups eliminated for CPLD Longitudinal Reinforcement 16#5 (1.58%, Specimen/ Prototype ratio = 1.14) 16#5 (1.58%, Specimen/ Prototype ratio = 1.14) No note Column Transverse Reinforcement #3 hoops @ 2 in #3 hoops @ 2 in Same basis allows CPLD joint to poten- tially be challenged to greater extent Longitudinal Reinforcement 8#5 and 2#4 top and bottom (0.50%, Specimen/ Prototype ratio = 0.99) 12#5 top and bottom (0.65%, Specimen/ Prototype ratio = 1.27) Potential benefit of flexural reinforcement reduced for CPLD Bent Cap Transverse Reinforcement 2-leg #3 stirrups @ 8 in 2-leg #3 stirrups @ 6 in CPLD stirrups reduced to minimum requirement

13 specimen was not tested for direct comparison. The CPLD bent cap reinforcement details are shown in drawings pro- vided in the attachments. Table 2.1 summarizes select joint details for the CPLD spec- imen, including the specifications for the helical corrugated pipe with lock seams, which were the same as the CPFD spec- imen. A comparison of the overall CPLD and CPFD specimen details is also presented in Table 2.1. Table 2.2 summarizes the significant differences in SDC D and SDC B design and detail- ing provisions. Significant joint reinforcement, including transverse (hoop) reinforcement, is required for SDC D but not for SDC B (17). The same pipe size and thickness were used for the CPLD and CPFD specimens to allow a direct comparison of specimens. The pipe thickness was not consid- ered excessive and was the minimum size readily available for construction. Table 2.1 reveals important differences that were intention- ally incorporated into the CPLD joint detailing to severely chal- lenge the limited ductility specimen, in accordance with the intent and provisions of the 2006 LRFD RSGS for SDC B: (1) elimination of the construction stirrups within the joint region, (2) elimination of all joint-related stirrups and horizon- tal ties (As jh, As jv) placed external to the CPFD joint, and (3) elim- ination of the extra hoop at each end of the pipe (2). Table 2.1 also reveals that bent cap flexural reinforcement was reduced to eliminate potential strengthening of the joint due to higher bent cap flexural strength (and thus to allow potential yielding of flexural reinforcement adjacent to and within the joint) and to provide more accurate prototype scaling. In addition, bent cap transverse reinforcement, including that adjacent to the joint, was based on shear developed within the bent cap due to forces associated with plastic hinging of the column, not a joint force transfer mechanism. These modifications were imple- mented despite the possibility that principal tensile stresses in the joint could exceed the 2006 LRFD RSGS limit of 3.5 , at which the additional joint reinforcement is required (2). Thus, ′fc these measures were deemed conservative for testing to exam- ine potential failure modes, and it was understood that more stringent detailing could be adopted for design as required. In addition, CPLD column reinforcement (including con- fining reinforcement) was not reduced but designed to match the SDC D-based requirements of the CPFD design. This was intended to help ensure that the column would not prema- turely become a weak link in the system, but impose as large of a demand and as many cycles as possible on the joint so that potential failure modes associated with the joint could be fully investigated. Matsumoto (26) provides a detailed comparison of column, bent cap, and joint reinforcement for the prototype bridge and CPLD specimen. These measures were deemed reasonably conservative for testing to examine limited ductility performance and potential failure modes. The impact of these measures was unknown. However, it was anticipated that more extensive joint damage would be exhibited than for the CPFD as the specimen dis- placement ductility approached µ2 and could possibly result in joint failure at larger ductility levels. It was understood that more stringent detailing could be adopted for SDC B design as required. Hybrid Specimens Lateral performance of hybrid systems differs from CIP and emulative systems due to the presence of unbonded post- tensioning and reinforcement. The prototype bridge served as the basis for the design of the CIP and emulative systems; how- ever, differences in the design were required for the hybrid sys- tems. To achieve a somewhat comparable lateral response, the hybrid specimens were designed to have similar lateral force resistance when compared to the CIP specimen at a 1.0% drift ratio. However, during the erection of the conventional hybrid specimen, the actual anchor set losses for the post- tensioning were significantly less than expected, resulting in a greater effective post-tensioning force. This increase in effec- tive post-tensioning resulted in an appreciably greater lateral resistance as compared to the CIP and emulative specimens. Detailed descriptions of the performance objectives, design methodology theory, and specimen designs for all three spec- imens are highlighted in the attachments to this report as well as by Tobolski (5). Complete design drawings for the hybrid specimens are provided as an attachment to this report. Each hybrid detail uses half of the conventional reinforcement as compared to the CIP and emulative specimens connected to the bent cap using a grouted duct connection with a grouted bedding layer joint dimension of 1 in. The conventional hybrid specimen used closely spaced spiral reinforcement at the column end to provide lateral confinement of the concrete compression toe. The flexural reinforcement in the column extended full height and was Figure 2.3. Rebar cage with corrugated pipe and supplementary hoops—CPFD.

14 Table 2.2. Design and detailing provisions—SDC D versus SDC B. NCHRP Project 20-7/Task 193 Criteria SDC D (Full Ductility) SDC B (Limited Ductility) Force Demands (8.3.2, 8.3.3) Based on forces resulting from the overstrength plastic hinging moment capacity or the maximum connection capacity following the capacity design principles specified in Article 4.11.* The lesser of the forces resulting from the overstrength plastic hinging moment capacity or unreduced elastic seismic forces in columns or pier walls. Ductility Demands (8.3.4) The local displacement ductility demands, µD, of members shall be determined based on the analysis method adopted in Section 5. The local displacement ductility demand shall not exceed the maximum allowable displacement ductilities established in Article 4.9. N/A Column Shear Demand, Vu (8.6.1) Based on the force, Vpo, associated with the overstrength moment, Mpo, defined in Article 8.5 and outlined in Article 4.11. Based on the lesser of (1) force obtained from an elastic linear analysis and (2) force, Vpo, for plastic hinging of the column including an overstrength factor Concrete Shear Capacity (8.6.2) Using concrete shear stress for circular columns with hoops, modified by: where is 6 (multicolumn bent) or lower inside plastic hinging region, per Eq. 4.9-5 Using concrete shear stress for circular columns with hoops, modified by: where is 2 Minimum Column Shear Reinforcement (Spiral) (8.6.5) Minimum Longitudinal Reinforcement (8.8.2) Splicing of Longitudinal Reinforcement in Columns (8.8.2) Minimum Longitudinal Reinforcement (8.8.3) Outside plastic hinging region N/A Minimum Development Length into Cap Beams (8.8.4) extended as close as practically possible to opposite face N/A** Anchorage of Bundled Bars into Increased by 20% for a two-bar bundle and 50% for a three-bar bundle. Four-bar N/A Cap Beams (8.8.5) bundles are not permitted in ductile elements. Maximum Bar Diameter (8.8.6) N/A Lateral Reinforcement Inside Plastic Hinge Region (8.8.7) Butt-welded hoops or spirals N/A Lateral Reinforcement Outside Plastic Hinge Region (8.8.8) Volumetric ratio shall not be less than 50% of that determined in 8.8.7 and 8.6. Reinforcement shall be of the same type. Lateral reinforcement shall extend into bent caps a distance that is as far as is practical and adequate to develop the reinforcement for development of plastic hinge mechanisms. N/A ' ' , not to be reduced; , Columns , Columns , Columns , Columns

15 locally debonded across the bedding layer to facilitate dis- tributed straining of the reinforcement. The concrete filled pipe hybrid specimen used a full height, steel shell that provided enhanced confinement at the column end. The flexural reinforcement extending from the bent cap into the column terminated after a given distance required for development in the column. After the termination of the rein- forcement, the shell, concrete, and unbonded post-tensioning were the only elements in the column. Similar to the conven- tional specimen, the reinforcement was locally debonded across the joint to prevent premature reinforcement fracture. The dual shell hybrid specimen used a full height exterior steel shell that provided confinement at the column end. To form the interior void, a corrugated metal pipe that was in con- formance with ASTM A760 was used (24). This interior pipe was intended to act as a stay-in-place form during fabrication as well as to prevent the potential implosion of the concrete section during large compressive strains associated with lateral response. Similar to the concrete filled pipe specimen, the rein- forcement extending from the bent cap into the column was terminated following adequate development. The three bent caps for the hybrid specimens were identi- cal and were designed in accordance with the 2006 LRFD RSGS (2). This design and detailing considered the various increases in reinforcement required in the joint as well as the flexural reinforcement in the bent cap. Joint shear design was performed considering only the area of flexural steel when determining the required joint shear reinforcement. During early discussions with the NCHRP Project 12-74 panel, the decision was made to use stainless steel reinforcement Table 2.2. (Continued). NCHRP Project 20-7/Task 193 Criteria SDC D (Full Ductility) SDC B (Limited Ductility) Requirements for Lateral Reinforcement (8.8.9) Various detailing requirements. Capacity Protection Requirements (8.9) Capacity-protected members such as bent caps shall be designed to remain essentially elastic when the plastic hinge reaches its overstrength moment capacity, Mpo. The expected nominal capacity is used in establishing the capacity of essentially elastic members. Superstructure Capacity Design (8.10, 8.11) For longitudinal direction, the superstructure shall be designed as a capacity-protected member. For transverse direction, integral bent caps shall be designed as an essentially elastic member. Longitudinal flexural bent cap beam reinforcement shall be continuous. Splicing of reinforcement shall, at a minimum, be accomplished using mechanical couplers capable of developing 125% of the expected yield strength, fye, of the reinforcing bars. Superstructure Design for Non- Integral Bent Cap (8.12) For superstructure to substructure connections not intended to fuse, provide a lateral force transfer mechanism at the interface. For connections intended to fuse, minimum lateral force at interface shall be 0.40 times the dead load reaction plus the overstrength shear key(s) capacity. Non- integral cap beams supporting superstructures with expansion joints at the cap shall have sufficient support length to prevent unseating. N/A N/A N/A N/A Joint Design (8.13) Major joint design and detailing provisions, such as bent cap width, joint shear reinforcement (vertical stirrups inside and outside the joint, and horizontal cross ties/J-bars), transverse joint reinforcement, additional bent cap longitudinal reinforcement, and side face reinforcement. N/A *Articles, sections, and equation numbers cited in Table 2.2 refer to 2007 LRFD PSGS (17). **AASHTO LRFD Bridge Design Specifications (3) provisions (i.e., 5.10.11.4.3), in contrast to NCHRP Project 20-7/Task 193 provisions, require even longer lengths for column bar extension into joint.

16 across the joint for the experimental specimens. The reason- ing for this decision relates to the localized crack in the hybrid system at nominal yield as opposed to the distributed cracking in a conventional CIP column. The extent of cracking at the bedding layer is highly localized due to the intention debond- ing across the joint. This will result in slightly larger crack widths at the nominal yield point. The use of stainless steel reinforcement locally across the joint serves to provide added comfort in the durability of these systems during their expected service life. To consider the potential influence of stainless steel reinforcement, a variety of material tests were conducted. Figure 2.4 provides a summary of a series of uniaxial tension tests conducted on No. 5 reinforcement for A706 steel and 316LN stainless steel. These results indicate that the stainless steel reinforcement has more than three times the uniform strain capacity as A706 steel, indicating significantly greater material ductility and energy dissipation capacity. The effective yield for the two steel grades was similar with a slightly greater ultimate tension capacity observed for the 316LN steel. In addition to uniaxial tension testing, a series of cyclic rebar tests were conducted. Figure 2.5a depicts the cyclic response of the two rebar specimens that were tested with an unbraced length equal to six times the bar diameter. The bars were loaded to a compression strain equal to approximately one-third of the previously reached tension strain. From Figure 2.5, it is apparent that the A706 and 316LN reinforcing bars have sim- ilar cyclic response for realistic free lengths. This indicates that commonly accepted relationships for steel reinforcement may be acceptable for the use of stainless steel reinforcement. Further study is needed to fully investigate the potential seis- mic implications of using stainless steel reinforcement. Figure 2.5b provides the complete cyclic stress-strain response of the two reinforcing bars. From this figure, it is apparent that the 316LN reinforcing bar has significantly greater ultimate tension strain capacity. The recording of strain was terminated at the last point on the 316LN plot due to the limits of the recording instrumentation. 2.2.2 Fabrication and Assembly of Nonintegral Specimens All nonintegral emulative specimens were fabricated at the precast yard of Clark Pacific (West Sacramento, California). Precast bent cap and column segments were then assembled at California State University—Sacramento (CSUS). The fabrica- tion and assembly of the precast specimens replicated as much (a) (b) Figure 2.4. Comparison of A706 and 316LN uniaxial tension response (5). Figure 2.5. Comparison of A706 and 316LN cyclic stress-strain to (a) 3% strain and (b) failure (5).

17 as possible the expected field process so constructability issues could be examined. The construction sequence for precast specimens included the following: 1. Fabricate rebar cages for the bent cap and column at CSUS. 2. Transport rebar cages to Clark Pacific, prepare bent cap and column forms, and cast bent cap and column concrete. 3. Transport precast cap and column to CSUS. 4. Prepare column and bent cap for assembly and conduct cap setting operation in upright position. 5. Prepare connection. For grouted duct connection, pump grout into the bedding layer to fill the bedding layer and ducts. For cap pocket connections, fill pocket and bedding layer with concrete from top of cap. 6. Invert specimen and install in test area. The following sections provide a brief summary and select photos of the specimen fabrication. Further details are pro- vided by Matsumoto (21, 22, 23, 26). Cast-in-Place Specimen Fabrication of the CIP specimen required building special elevated forms for casting the bent cap on top of the column (see Figure 2.6 and Figure 2.7), as well as inverting the entire T-shaped specimen in the yard for transportation. The speci- men was fabricated accurately according to the drawings. Grouted Duct Specimen Figures 2.8 through 2.12 show the bent cap rebar cage, cap setting operation, and grouting of the GD specimen. Grout compressive strength was designed to exceed that of the bent cap by at least 500 psi to ensure the connection grout was not a weak link in the system. A hand pump system and collar were used for grouting the bedding layer and ducts. Grout was pumped from the bottom of the bedding layer up into the ducts, and an air vent system at the top of the bedding layer helped prevent air entrapment within the connection. Fluidity of grout was determined before grouting using a flow cone test in accordance with ASTM C939-02 (27). After the bedding layer form was attached and sealed, the bedding layer Figure 2.6. Lowering bent cap rebar cage into elevated formwork—CIP. Figure 2.7. Bent cap rebar cage in form during fabrication—CIP. Figure 2.8. Bent cap rebar cage in form during fabrication—GD.

18 dam at the top and bottom of the corrugated pipe and a column bar template were useful to form the cap pocket void full height of the bent cap, as the pipe is placed only between the top and bottom of the longitudinal rebar (see Figure 2.13b). Figure 2.14 shows concreting of the CPFD pocket. Fabrication and assem- bly operations for the CPLD and CPFD were the same. Concrete was placed in the pocket and bedding layer using a bucket at the top of the pocket. Concrete was cast into the pocket around the bent cap longitudinal reinforcement from above, and a collar with an air vent system was used to help remove entrapped air at the bedding layer. The concrete mix was selected to be close to that used for the bent cap and col- umn, with the intention of achieving a strength and stiffness at least 500 psi greater than that of the bent cap, ensuring that the connection would not be the weak link in the system. After the bedding layer form was attached and sealed (see Figure 2.15), the bedding layer was prewatered to ensure sealing and to pre- vent loss of moisture from the pocket concrete. Buckets were used to fill the pocket with concrete in several layers with vibration. Once concrete flowed through the air vents in the bedding layer, the vents were sealed. After hardening, curing compound was applied to the top surface. After the bedding layer form was removed, the bedding layer and top of the pipe were inspected. The first column hoop below the top of the CPFD column was placed approximately 2 in below its intended location dur- ing fabrication. This reduced the overall drift to some extent, but did not affect the maximum load induced in the joint. Conventional Hybrid Specimen Figures 2.16 through 2.19 show the rebar cage, cap set- ting operation, grouting, and post-tensioning of the con- Figure 2.10. Cap placement during and after cap setting operation—GD. was prewatered to ensure sealing and prevent loss of moisture from the grout. After mixing the grout using a paddle-type mortar mixer, grout with a flow cone efflux time of 20 to 30 sec was pumped into the connection. After grout flowed through air vents in the bedding layer, the vents were sealed. Grout was added manually to top off each duct. After the grout cured several days, the bedding layer form was removed and the bedding layer and the top of the ducts were inspected. Cap Pocket Specimen Figure 2.13 compares the bent cap rebar cage for the CPFD and CPLD specimens. The significant reduction of joint re- inforcement for the CPLD specimen is evident. A Sonotube Figure 2.9. Joint region of bent cap during fabrication—GD.

19 ventional hybrid specimen. Unlike the emulative grouted duct specimen, the grouting of the bedding layer and ducts for this specimen were performed by pumping the grout in from the top. The grout tube was inserted into a corrugated duct extending near the bottom of the bedding layer. As the grout filled the bedding layer, the grout tube was slowly extracted from the corrugated duct. The hydraulic head pressure of the column of grout in the one duct was used to fill the remaining ducts with some head loss. Similar to the grouted duct connection, each duct was then topped off in a way similar to what is shown in Figure 2.12. Once grout had adequate time to cure, the column and bent cap assembly was post-tensioned, inverted, and installed in the testing frame. Concrete Filled Pipe Hybrid Specimen The reinforcing cage and details of the bent cap for the con- crete filled pipe and dual shell specimens are the same as those presented in Figure 2.16 for the conventional hybrid specimen. A view down the inside of the column prior to casting is shown in Figure 2.20. In this figure, the installed curvature gages are apparent. Also in this photo, weld beads on the inside of the column can be observed. These weld beads were placed inside the column shell to promote reliable transfer of reinforcement tensile forces into the shell. The erection of this specimen was the same as the erection of the conventional specimen. The bedding layer form used during casting can be seen in Figure 2.21. During the casting of the bedding layer for this specimen, minor Figure 2.11. Mixing and pumping of grout—GD. Figure 2.12. Topping off ducts with grout and cap top post-grouting—GD.

20 (c) CPLD (West Face) (d) CPFD (West Face) (a) CPLD (b) CPFD Figure 2.13. Comparison of cap pocket bent cap rebar cages during fabrication.

21 (a) Bucketing Concrete into Pocket (b) Vibrating Concrete in Pocket Figure 2.14. Concreting of cap pocket connection—CPFD. (a) Plan View of Pocket (b) Bedding Layer Figure 2.15. Pocket and bedding layer before concreting—CPLD. Figure 2.16. Reinforcing cage installed in form— HYB1. bleeding of the bedding layer grout was observed. Following the removal of the formwork, no defects were noted in the bedding layer as only minor bleeding was seen. Dual Steel Shell Hybrid Specimen A view down the region between the internal and external shell for the dual shell specimen can be seen in Figure 2.22. Similar to the concrete filled pipe specimen, in this column detail weld beads were placed on the inside of the column to promote the transfer of forces from the reinforcing bar to the shell. In Figure 2.22, the debonding material on the rebar can also be seen, along with the termination of the bar within the

22 Figure 2.17. Bent cap setting operation—HYB1. Figure 2.18. Bedding layer and corrugated duct grouting operation—HYB1. Figure 2.19. Post-tensioning operation—HYB1. Figure 2.20. View inside steel shell showing weld beads—HYB2. column. A top view of the column after casting is shown in Figure 2.23. From this figure, the internal corrugated metal pipe is observed with a polyvinyl chloride (PVC) pipe in the center for threading of column post-tensioning. This PVC pipe was used to ensure the easy threading of tendons from anchorage to anchorage. For corrosion purposes, this PVC pipe can be grouted following stressing to prevent moisture from reaching the tendon during its service life. This grout- ing will not affect the unbonded nature of the tendon because the bond between the grout and PVC will break easily. 2.2.3 Nonintegral Testing Protocol and Instrumentation Emulative Specimens The specimen test setup, shown in Figure 2.24, included a simply supported inverted bent cap that allowed accurate

23 establishment of specimen forces. The test setup ensured accu- rate conditions at each end of the joint so that the force trans- fer mechanism in the joint could be investigated (15, 28). The specimen was tested in inverted position with a column stub that allowed biaxial loading of the specimen using a vertical hydraulic actuator to apply scaled gravity load and a horizon- tal hydraulic actuator to induce seismic response. As required, different axial force conditions in the bent cap were produced for the push and pull directions. Force control and displacement control sequences were applied to all specimens, similar to the force and control sequences shown in Figure 2.25 and Figure 2.26. Force control loading was used for an approximate determination of first yield of column longitudinal bars in the push and pull direc- tions, establishment of effective yield, and application of the displacement control sequence including quasi-static displace- ment in three cycles. Nominal displacement ductility demand, as multiples of system effective yield displacement, was applied at the following levels, or until the residual capacity of the spec- imen dropped below 30% of the maximum load: µ1, µ1.5, µ2, µ3, µ4, µ6, µ8, and µ10. Figure 2.24b shows the external gages, including linear and string potentiometers and linear variable differential trans- formers (LVDTs), mounted on the column, joint, and bent cap. Internal strain gages were placed on bent cap, joint, and column reinforcing bars, as well as on corrugated ducts or pipe. In addition to the approximately 100 channels of data, speci- men response was also monitored using digital photos, crack markings and measurements, video recording, and notes. Hybrid Specimens The test setup for the hybrid specimens is shown in Fig- ure 2.27. As shown in previous images, the hybrid specimens were constructed in an upright condition and then inverted for installation in the test setup. The vertical actuator was set to apply a constant load during testing to simulate gravity load- ing. This force varied between hybrid specimens in order to try and match the lateral response of the three hybrid tests. The horizontal actuator was actively controlled to apply specified forces or displacements during testing. The initial stage of loading consisted of force controlled loading protocols, which apply positive and negative lateral forces of increasing magnitude until the first yield of the extreme mild reinforcing bar is reached. This force control protocol is shown in Figure 2.28. Each force loading cycle was repeated three times in both directions. Following the first yield of the system, the lateral loading was applied to a speci- fied lateral drift ratio. The basic loading protocol is shown in Figure 2.29. At each cycle to a given drift ratio, the column was subjected to two cycles in both directions followed by one cycle to the previous lateral drift. This protocol was developed to help accurately calibrate nonlinear models of the system. Figure 2.21. Bedding layer form—HYB2. Figure 2.22. View down dual shell prior to casting with rebar—HYB3. Figure 2.23. Top view of column after casting—HYB3.

24 CAP ROTATION PANEL DEFORMATION COLUMN DISPLACEMENT HORIZONTAL ACTUATOR VERTICAL ACTUATOR COLUMN CURVATURE (NORTH) COLUMN CURVATURE (SOUTH) CAP ROTATION (a) Schematic (b) Specimen in Test Bay with External Instrumentation Shown—GD Figure 2.24. Test setup for emulative specimens.

25 -50 -40 -30 -20 -10 0 10 20 30 40 50 0 1 2 3 4 A pp lie d Fo rc e (ki ps ) Cycles Note: Vertical Force held constant at 38 kips Figure 2.25. Representative force controlled sequence for emulative specimens. -5.5 -4.8 -4.1 -3.4 -2.8 -2.1 -1.4 -0.7 0.0 0.7 1.4 2.1 2.8 3.4 4.1 4.8 5.5 -8 -7 -6 -5 -4 -3 -2 -1 0 1 2 3 4 5 6 7 8 0 3 6 9 12 15 18 21 D rif t R at io (% ) Cycles D is pl ac em en t D uc til ity , µ Δ Note: Vertical Force held constant at 38 kips. Figure 2.26. Representative displacement controlled sequence for emulative specimens.

26 External instrumentation mounted on the specimens is shown in Figure 2.30. Instrumentation consisted of linear potentiometers and inclinometers for measuring and isolating various modes of deformation in the member. In addition to the external instrumentation, many internal strain gages were employed to capture the local response of materials. 2.2.4 Design of Integral Prototype Bridge and Specimen An overall elevation of the prototype bridge is shown in Figure 2.31 with the connection detail shown in Figure 2.32. The design of the prototype bridge was completed in accor- dance with the AASHTO LRFD Bridge Design Specifications (29) and the 2009 LRFD SGS (1). The selection of initial member sizing was based on conven- tional design practices and span range tables for girder systems used. The 74-in deep Washington DOT post-tensioning beam was selected using recommended span limits published for these girders. This girder section was selected over a bulb-tee due to the increased bottom flange area desirable for nega- tive flexural demands at the bent cap. The design was refined through the application of LRFD design requirements, as needed. To minimize the neutral axis depth, a design 28-day compressive strength of 9 ksi was used for the prototype structure. Post-tensioning in the girders was designed so that the ulti- mate, extreme event and service limit states were satisfied. The design of the post-tensioning was governed by the Strength I limit state with the seismic demands of a similar magnitude. Two stages of post-tensioning were specified, with the first stage occurring prior to the deck casting and the second stage occurring after the deck casting. Service level performance of the structure was considered in the prototype design through Figure 2.27. Test setup for hybrid specimens. Cycle Fo rc e, k ip s 0 1 2 3 4 5 6 7 8 9 -60 -40 -20 0 20 40 60 Figure 2.28. Representative hybrid force controlled loading protocol. Cycle D rif t R at io 9 12 15 18 21 24 27 30 33 -10% -6% -2% 2% 6% 10% Figure 2.29. Representative hybrid displacement controlled loading protocol.

27 Figure 2.30. Representative hybrid external instrumentation (HYB1 shown). Figure 2.31. Portion of integral prototype bridge. Figure 2.32. Girder to bent cap prototype connection detail.

a construction staging analysis explicitly considering the devel- opment of stresses in the system at various stages. For seismic design, the prototype bridge was considered nonessential and designed to meet life safety requirements as defined by the 2009 LRFD SGS (1). The specified mechanism of inelastic deformation in the longitudinal direction consists of flexural plastic hinge formation at both the tops and bottoms of columns and knock off backwalls. Additionally, the super- structure to bent cap joint was allowed to open during seismic excitations as long as the response was essentially elastic. The transverse mechanism involved the development of flexural plastic hinging at both the tops and bottoms of columns and shearing of sacrificial shear keys at the abutments. The design ARS was developed in accordance with the 2006 LRFD RSGS (2). The ARS curve incorporated 5% damping and was developed using a 1-sec acceleration of 0.80 g, a 0.2-sec acceleration of 1.50 g, and site coefficients for Site Class D soil. The resulting peak rock acceleration for the prototype design for the study site was 0.60 g. The input seismic demand and site classification resulted in a bridge subject to SDC D require- ments. This ARS curve is representative of a site located in a high seismic region such as Southern California. The imposed demand levels required a seismic demand analysis, displace- ment capacity analysis with pushover, capacity design provi- sions, and SDC D detailing. Due to the assumed site location, the bridge was considered located within 6 miles of a fault. Therefore, vertical ground motion with a magnitude of 0.80 g of vertical excitation was considered. The structural system was modeled in the computer analysis program SAP2000 for service, strength, and seismic design. Modeling procedures for seismic analysis were performed based on provisions of the 2009 LRFD SGS (1). Effective sec- tion properties were modeled to accommodate the expected dynamic behavior of the bridge system, including column inelasticity. Dynamic analyses indicated that the dominant transverse period of vibration is equal to 0.73 sec and the dominant longitudinal period of vibration is equal to 0.69 sec. The longitudinal analysis considered the effects of the backwall stiffness in accordance with the provisions of the 2009 LRFD SGS. The resulting displacement demands in the longitudinal and transverse directions were 6.8 in and 7.5 in, respectively. In order to determine the displacement capacity of the sys- tem, an inertial pushover analysis was conducted in both the longitudinal and transverse directions. In the longitudinal direction, the bent cap to girder joint was allowed to open during seismic excitation. This decision was made based on extensive discussions with the research team and project panel. It was decided that allowing the bridge to flex and open at the joint was acceptable as long as the joint responded in an essen- tially elastic manner. To consider the potential joint opening, a superstructure moment-rotation hinge was modeled at the face of the bent cap. For the prototype design, these hinges were based on moment-curvature analyses of the superstructure and an assumed equivalent plastic hinge length. The original design considered an effective hinge length to be 1 ft; however, a more realistic length is approximately one-half of the structure depth. It is expected that allowing the superstructure joint to hinge will result in a redistribution of seismic moment demand due to the reduction in stiffness following hinging. The proto- type design resulted in seismic moment redistribution of approximately 20%. The observed response from analysis indi- cated that the system is expected to respond in an essentially elastic manner. In the longitudinal direction, the displacement capacity determined via pushover analysis is 11.0 in, which results in a demand-to-capacity ratio of 0.62. The results of the longi- tudinal pushover analysis indicated that positive joint open- ing is expected, but without appreciable rotation demand. The response of the joint is classified as essentially elastic because the calculated rotations are only slightly greater than the elastic rotation. For the transverse displacement capacity, overturning effects were considered in determining the column inelastic response. The transverse displacement capacity determined via pushover analysis was 10.6 in, which results in a demand-to-capacity ratio of 0.71. Ductility demands for both directions were approximately 5, well below the limit of 8 for multicolumn bent caps. Capacity design principles were applied to the design of the superstructure to ensure that the seismic overstrength demands could be resisted in a nominally elastic manner. Column transverse reinforcement was designed based on overstrength demands imposed by transverse response. Flexural and shear demands on the cap beam were based on the demands devel- oped with flexural hinging of the columns at overstrength demands. Design of bent cap reinforcement was based on over- strength demands in addition to the longitudinal and trans- verse force-transfer mechanisms assumed in the AASHTO LRFD Bridge Design Specifications (3). The main goal of the experimental effort was to determine the response of the girder to bent cap joint when subjected to simulated seismic demands. To satisfy this goal, a portion of the prototype structure was extracted for experimental testing. The test specimen selected for testing consisted of a girder, deck, and reaction block. This specimen was based on an extracted portion of the prototype bridge, as shown in Figure 2.33. The scaled length of the girder used in the experimental program was 31 ft, which is representative of 0.46 times the central span length. The bent cap was represented by a large reaction block in the testing of the specimen. The flexibility of the bent cap can be neglected in experimental efforts as the deformation from this member can be considered in analytical modeling. The extracted portion of the prototype bridge provides sufficient 28

29 information regarding the moment-rotation response of the joint and shear transfer across the joint. These two items are the major unknowns in the performance of this integral bridge detail. 2.2.5 Fabrication and Assembly of Integral Specimen The integral test specimen was fabricated and constructed at the University of California—San Diego (UCSD) Charles Lee Powell Structural Systems Laboratory. Labor for construction activities was provided by a combination of subcontracted construction firms and lab staff. A local steel fabrication company performed the majority of construction activities related to fabrication of the steel reinforcement cages. A local construction company experienced in bridge construction performed the majority of construction activities related to building formwork. All post-tensioning activities were per- formed by a post-tensioning manufacturer and contractor. Lab staff performed the casting of concrete and all activities related to erection of members. The first stage of fabrication consisted of the construction of the reinforcing cages for the girder and reaction block. The girder reinforcing cage can be seen in Figure 2.34 with the post-tensioning ducts installed and the cage installed in the formwork. The scaling of this specimen did not provide sig- nificant access for pencil vibrators in the duct. To ensure that adequate consolidation would be achieved during fabrication, a form vibrator was used in regions with limited pencil vibrator access. In addition, a superplasticized concrete mix was used to enhance flowability during casting. Concrete for the girder and reaction block was cast using a bucket attached to the overhead crane in the laboratory. The girder was cast first to ensure that the maximum flowability of the concrete mixture was obtained during the casting of the member with limited vibrator access. During casting of the girder, an external form vibrator was attached to the form- work near the location at which concrete was being poured. This vibrator was moved around the formwork as concrete was placed in different locations and was used on both sides of the formwork. At the end regions, where more sufficient vibrator access was provided, traditional pencil vibrators were used. The girder and reaction block were cast and allowed to harden until the girder had strength greater than 3 ksi. Inspec- tion of the girder after form removal indicated only one region of minor concrete segregation over a small portion of the bot- tom flange of the girder (approximately 4 in. in length). This region was patched by lab staff following placement on tempo- rary supports. The girder was moved away from the reaction block to facilitate the construction of two temporary support towers. Following the completion of these towers, the girder was lifted and placed on the towers in line with the reaction block (see Figure 2.35). The girder was leveled on the temporary supports and subsequently secured using chains to provide stability during construction activities. The girder was placed to maintain an approximate 1-in clo- sure joint between the reaction block and girder. This joint can be seen in Figure 2.36. Additionally, the alignment was checked to ensure the post-tensioning ducts were properly aligned. The careful activities carried out during the placement of the Figure 2.33. Test specimen representation— post-tensioned integral specimen (INT). Figure 2.34. Girder reinforcing cage.

30 girder resulted in a system in which the post-tensioning ducts and closure joint were properly aligned with no noticeable variations. The post-tensioning ducts were then jointed using industrial adhesive tape, which was applied by hand. The scaled specimen made the joining of these ducts slightly cumbersome, as hand access was tight. However, the splicing of these ducts was performed without any major complications. The girder formwork was modified and reused as the closure joint formwork by drilling new holes in the form and reusing the original form tie holes in the girder. After installation of the side forms, the bottom of the joint was closed using a single piece of plywood. A drain hole was placed in the bottom of the form to allow for draining of excess water (water is used to moisten the faces of the reaction block and girder prior to grouting). The edges of the formwork were sealed using a com- mercially available sealant and allowed to set prior to grouting. Grout material was mixed on the laboratory floor and lifted onto the top of the specimen. The grout was then gravity fed into the closure joint, as shown in Figure 2.37. The grout material was Masterflow 928 high-strength, non-shrink grout containing a 0.2% volume fraction of polypropylene fibers. This grout matrix was mixed to be flowable based on manu- facturer’s recommended water content and considering the presence of the fibers. The relatively low volume fraction of fibers did not greatly affect the flowability of the matrix. No noticeable leakage was observed during the grouting activities. The grouting activities were completed without any observed complications. Formwork was removed from the girder the following day. Observations after removal of the formwork indicated no observable voids in the closure joint and overall a very success- ful grouting operation. The grout was allowed to cure for 3 days prior to the first stage post-tensioning. The first portion of the post-tensioning operation consisted of setting the wedges for the bottom tendons. Each strand was stressed to approximately 5% guaranteed ultimate tensile strength (GUTS) to allow for sufficient seating of the wedge on the live end. The middle ten- don was then stressed to a target stress of 75% GUTS. Each strand was individually stressed using a monostrand jack. Both the bottom and middle ducts were then grouted using SikaGrout 300PT. Following 2 days of curing in the post- Figure 2.35. Girder on falsework prior to grouting closure joint. (a) (b) Figure 2.36. Girder post-tensioning duct (a) prior to splicing and (b) after splicing.

31 tensioning grout, the temporary support near the reaction block was removed, simulating the removal of the strong back in the prototype structure. The post-tensioning force in the middle tendon at this stage provides a sufficient shear friction mechanism in the system for casting of the deck and associated construction activities. Formwork for the reinforced concrete deck was constructed to react off the girder, as is customary in precast concrete bridge construction (see Figure 2.38). The construction of the deck formwork utilized the existing holes in the girder from the form ties to secure the forms in place. With the formwork in place, the deck reinforcing cage was fabricated. Similar to cast- ing of the deck and reaction block, the deck was cast using a bucket attached to the overhead crane. The deck formwork was removed after the minimum uniaxial compressive strength of the deck concrete was 3 ksi. The top post-tensioning tendon was stressed following the casting and curing of the deck. This tendon was stressed to 75% GUTS, similar to the middle tendon. Additionally, each strand in the tendon was individually stressed using a monostrand jack. Following post-tensioning, the duct was grouted using SikaGrout 300PT and allowed to cure. The specimen was then painted white to aid in the identification of cracking during testing. Loading frames and external instrumentation were subsequently installed on the specimen, in addition to the installation of vertical actuators in preparation of testing. 2.2.6 Integral Specimen Testing Protocol and Instrumentation The general testing system is shown in Figure 2.39. From this figure, the nomenclature for actuator reference and plan location reference can be observed. For the integral testing protocol, the first stage of loading consisted of relieving the reaction from the temporary support installed between Actuator 1 and Actuator 2. The goal of this stage was to relieve the reaction while minimizing the associ- ated displacement. Actuators 2 and 3 were set to force control with zero force. Actuator 1 was controlled in manual displace- ment control and applied upward displacements until the load on the temporary support tower was relieved. The next stage of loading was designed to apply the simu- lated dead loading. Actuator 1 was controlled in displacement control and applied upward displacements until a specified force was reached. Actuators 2 and 3 were slaved to Actuator 1 in force control to apply moment and shear profiles as shown in the prototype specimen section. For all additional stages, Actuator 1 was controlled in dis- placement until either a specified force or joint rotation limit was reached. A modified equation relating the force in Actuators 2 and 3 to the force in Actuator 1 was used. This new Figure 2.37. Grouting of girder to reaction block closure joint. Figure 2.38. Construction of concrete deck reacting off girder. Figure 2.39. Integral specimen loading system.

32 loading equation is based on the seismic flexural and shear demand profiles determined via lateral analysis of the proto- type structure. Following the application of the simulated dead loading, 100 cycles of essentially elastic loading were imposed on the system primarily in the negative flexural direction. This load- ing was meant to allow investigation of the potential response of the system in service and ultimate loading. The system was loaded initially in the negative flexural direction until the system was nearing the expected limit of proportionality. Following this, the system was loaded to 90% of the initial dead load demand. This was repeated for 100 cycles in contin- uous operation. The next stage of loading was simulated seismic demands. The loading demands generated for the seismic stage were based on a combination of lateral seismic load demands and vertical seismic shear demand. The vertical seismic shear demand was held constant during all phases, with constant loading applied at the actuator nearest the joint. Flexural moment and shear demands were based on scaled flexural demands caused by simulated column overstrength demands. The simulated flexural demands impose flexural moments at the girder to reaction block interface that are applicable to vertical, lateral, or seismic settlement demands. The cyclic load- ing is conservative for loadings generated by seismically induced settlement. Additionally, the additional shear demand applied at the joint is conservative for lateral loading scenarios. This loading program was developed to conservatively encom- pass the various potential loading cases. The loading was controlled while operating Actuator 1 in displacement control set to hold when a specified joint rotation limit was reached. The rotation targets were initially specified using the linear potentiometers at the joint. However, it became apparent during testing that these calculated rotations were not correct during increasing displacements due to cracking at potentiometer target supports and spalling of concrete. The actual achieved rotations were reassessed based on more reliable inclinometer readings, which better match the observed rotations during testing. The actual loading protocol used during the seismic cycle is shown in Figure 2.40. As a main goal of large joint rotation cycles is to determine the overall rotation capacity of the connection due to relative settlement potential, a reversed cyclic loading protocol, which produces a highly sever case, was used. This is because relative settlement demands on the connection will only occur in one loading direction and not the reverse. This protocol probably caused a reduction in the actual ultimate rotation capacity when subjected to loading in a single direction. Another driving factor in the development of this testing protocol was the desire to determine the inelastic rotation response in the event of superstructure inelasticity. The test specimen was instrumented to capture the major response characteristics of the specimen when subjected to applied loadings. This instrumentation includes strain gages mounted on rebar and post-tensioning and external gages mounted onto the specimen. External instrumentation consists of linear potentiometers, string potentiometers, and inclinometers mounted on the exterior of the specimen. A summary of external instrumenta- tion is shown in Figure 2.41. Linear potentiometers are placed to capture opening of the joint, slip between the girder and reaction block, and estimated rotation of the girder at the joint. String potentiometers are installed to capture the displacement of the specimen at the actuator locations. Inclinometers are installed to capture the rotation of the reaction block and girder at the joint. Strain gages were installed on deck longitudinal Cycle Jo in t R ot at io n, ra d * 1 00 0 9 18 27 36 45 54 -1.25 -0.75 -0.25 0.25 0.75 1.25 Figure 2.40. Integral specimen realized loading protocol.

33 reinforcement, girder longitudinal reinforcement at the base, and girder shear reinforcement. 2.3 Test Results This section summarizes key aspects of specimen response for the emulative, hybrid and integral experimental tests. Detailed results are provided by Matsumoto (30, 21, 22, 23, 26) and Tobolski (5). In reporting specimen response, displacement ductility (µ) and drift ratio are both used. The drift ratio is the column dis- placement divided by the column height and is reported as a percentage. This is a more consistent basis for comparison of specimen response than displacement ductility. However, sys- tem ductility levels are also reported, although these values should be considered nominal (i.e., approximate) due to the approximate determination of first yield. The terms “drift” and “drift ratio,” are used interchangeably. 2.3.1 Nonintegral Emulative Connections This section summarizes primary aspects of specimen response, including column hysteretic response (lateral force displacement), displacement decomposition, and joint response. Comparisons are made between the CIP and precast connections, as well as between the full and limited ductility specimens. The lateral force displacement (hysteretic) response of the column is used to characterize the fundamental performance of the specimen. Displacement decomposition refers to the separation of the column displacement into the components that contribute to the overall lateral displacement of the column (column flexure, fixed end rotation due to plastic hinging and bar slip, bent cap flexibility, and joint shear). Comparisons are made between decomposition for analytical predictions and experimental measurements. Joint response includes a sum- mary of joint cracking, principal stresses, joint deformation, and strain records. Cast-in-Place Specimen CIP specimen response was dominated by plastic hinging of the column adjacent to the bent cap, as shown in Figure 2.42 and Figure 2.43. The specimen exhibited excellent ductility to a large drift of 5.9% (nominal displacement ductility of 10), and the load-displacement response indicated stable hysteretic behavior without appreciable strength degradation. Post-test inspection revealed that the core remained primarily intact with several column bars buckling and fracturing at ultimate. Initial spalling of the column occurred at 1.8% drift (µ3), with progressive spalling at higher drifts. In contrast to significant Figure 2.41. Integral specimen external instrumentation plan. Figure 2.42. Specimen response at a 2.3% drift ratio (µ4)—CIP.

column flexural and shear cracks and spalling associated with increasing lateral force, relatively minor cracking occurred in the joint region (0.025 in maximum) without spalling. This corresponded to stiff joint shear response and limited soften- ing, with the contribution of joint shear to column displace- ment averaging 3.4%. Principal tensile stresses significantly exceeded 3.5 psi and justified the use of additional joint′fc reinforcement required for development of a force transfer mechanism. Bent cap longitudinal bars reached only 46% of yield. However, the north construction stirrup within the joint yielded, indicating its contribution to the stable joint performance. Column Lateral Force versus Lateral Displacement. The lateral force displacement (hysteretic) response of the column, shown in Figure 2.44, indicates stable hysteretic behavior with loops of increasing area without appreciable strength degrada- tion. A comparison of the load-displacement envelope to the predicted envelope showed a good correlation. The hysteretic response also portrayed appropriate stiffness, strength, duc- tility, and features such as crack distribution and width rep- resentative of appropriate response for a CIP beam-column connection. The dominance of ductile plastic hinging in the column and minimal damage in the capacity-protected joint and bent cap satisfied the performance goal for the CIP control specimen. Thus, the specimen provided an appropriate base- line for comparison with the precast specimens. Column Displacement Decomposition. Column dis- placement decomposition, summarized in Figure 2.45, confirmed the dominance of plastic hinging and showed that displacement components were reasonably determined and predictions were reasonably made. The joint shear displace- ment was minor, contributing only 3.4% on average to the overall column displacement, and was consistent with visual observations of minor joint cracking. Splitting cracks formed 34 -6.09 -5.22 -4.35 -3.48 -2.61 -1.74 -0.87 0.00 0.87 1.74 2.61 3.48 4.35 5.22 6.09 -80 -70 -60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60 70 80 -3.5 -3.0 -2.5 -2.0 -1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 Drift Ratio (%) La te ra l F or ce (k ip) Lateral Displacement (in) Actual Response Actual Envelope Predicted Response Maximum actuator pull stroke reached Figure 2.43. Specimen response at a 5.9% drift ratio (µ10)—CIP. Figure 2.44. Lateral force versus lateral displacement—CIP.

in the bent cap and column, and the top surface of the bent cap (as tested) exhibited splitting cracks and local spalling; how- ever, column bars were well anchored within the joint, with bar slip contributing less than 4% on average to fixed end rotation. Joint Response. As shown in Table 2.3, CIP joint distress was limited. Analysis of the joint indicated that the principal tensile stress was limited to 5.4 psi, less than half of the 2006 LRFD RSGS (2) limit of 12 psi, but about 50% larger than 3.5 psi, the level at which more extensive (additional) joint reinforcement is required for development of the assumed force transfer mechanism. Principal compressive stresses did ′fc ′fc ′fc not exceed 0.09f ′c, less than a third of the 2006 LRFD RSGS limit of 0.25f ′c. These values correspond well with the inten- tions of the design and the observed joint performance. The joint shear stress-strain response was appropriately stiff and exhibited minor softening at increasing drift (see envelope in Figure 2.46). This correlated well with the maximum surface crack width in the joint region that was limited to 0.025 in (with no surface spalling), as shown in Figure 2.47, as well as displacement decomposition results. Joint deformation was very small, with maximum change in panel area limited to less than 0.2%. Bent cap longitudinal bars did not yield, reaching only 46% of yield, even though additional bent cap longitudi- nal reinforcement (0.245Ast) required by 2009 LRFD SGS was not included (1). Stirrup strain outside the joint remained well 35 9.1% 4.6% 4.2% 3.7% 3.4% 3.0% 2.8% 2.8% 18.4% 16.2% 14.3% 10.9% 9.0% 7.0% 6.3% 5.9% 20.9% 18.3% 20.4% 25.0% 21.9% 26.4% 28.1% 24.7% 51.6% 60.9% 61.1% 60.5% 65.7% 63.6% 62.9% 66.6% -2.4% -8.0% -7.6% -5.2% -4.0% -3.2% -2.8% -2.3% -16.1% -18.9% -15.8% -11.0% -8.4% -6.3% -5.4% -4.5% -26.5% -23.5% -22.4% -27.1% -28.1% -28.1% -31.2% -34.3% -55.1% -49.5% -54.2% -56.7% -59.4% -62.5% -60.6% -58.9% -100% -80% -60% -40% -20% 0% 20% 40% 60% 80% 100% 0.69 (FC48) 0.93 (µ1.5) 1.19 (µ2) 1.76 (µ3) 2.34 (µ4) 3.60 (µ6) 4.59 (µ8) 5.88 (µ10) D is pl ac em en t C om po ne nt Drift Ratio (%) (Displacement Ductility) Fixed End Rotation Pull Column Flexure Pull Bent Cap Flexure Pull Joint Shear Pull hs uP ll uP Figure 2.45. Displacement decomposition component percentages—CIP. Parameter CIP GD CPFD CPLD Joint Shear Stress (psi) 328 (4.86 ) 312 (4.62 323 (4.31 371 (6.32 Principal Tensile Stress (psi) 363 (5.38 ) 343 (5.09 356 (4.75 411 (6.99 Principal Compressive Stress (psi) 401 (0.088 ) 370 (0.081 ) 398 (0.071 ) 460 (0.13 ) Angle of Principal Plane (deg) 45.0 45.0 44.2 44.8 Joint Rotation (rad) 1.95 × 10-3 2.25 × 10-3 1.73 × 10-3 2.87 × 10-3 Change in Panel Area (%) 0.16 0.19 0.13 0.46 ) ) ) ) ) ) Table 2.3. Maximum joint response—all specimens.

36 -0.86 -0.57 -0.29 0.00 0.29 0.57 0.86 -500 -400 -300 -200 -100 0 100 200 300 400 500 -0.015 -0.010 -0.005 0.000 0.005 0.010 0.015 Joint Shear Strain (deg) Jo in t S he ar St re s s (p si ) Joint Shear Strain (rad) CIP GD CPFD CPLD (g) East Side—GD(e) East Side—CIP (c) East Side—CPLD(a) East Side—CPFD (h) West Side—GD(f) West Side—CIP (d) West Side—CPLD(b) West Side—CPFD Figure 2.46. Joint shear stress versus joint shear strain envelopes—all specimens. Figure 2.47. Joint region cracking post test—emulative specimens.

below yield, but the north construction stirrup within the joint yielded, as shown in Figure 2.48, indicating its contribution to the stable joint performance. Grouted Duct Specimen GD specimen response was dominated by plastic hinging of the column adjacent to the bent cap (Figure 2.49 and Fig- ure 2.50), as intended by the emulative assumption in the design. Similar to the CIP specimen, the GD specimen exhibited excellent ductility to a large drift of 5.2% (nomi- nal displacement ductility of 8), and load-displacement response indicated stable hysteretic behavior without apprecia- ble strength degradation. Post-test inspection revealed that the core and bedding layer remained primarily intact with several column bars buckling and two bars fracturing at ultimate. Initial spalling of the column developed at the column- bedding layer interface at 1.2% drift (µ1.5), with progressive spalling at higher drifts. As for the CIP specimen, significant column flexural and shear cracks and spalling developed, but relatively minor cracking occurred in the joint region (0.040 in maximum). This corresponded to stiff joint shear response and limited softening, with the contribution of joint shear to col- umn displacement averaging 4.9%. Column bars were well anchored within the ducts, with only minor bar slip evident. Principal tensile stresses significantly exceeded 3.5 and justified the use of additional joint reinforcement required for development of a force transfer mechanism. Similar to the CIP specimen, bent cap longitudinal bars reached only 53% of yield. The south construction stirrup reached 75% of yield, indicating its contribution to the stable joint performance. ′fc Column Lateral Force versus Lateral Displacement. The lateral force displacement (hysteretic) response of the GD col- umn, shown in Figure 2.51, indicates stable hysteretic behav- ior with loops of increasing area without appreciable strength degradation, as well as stiffness, strength, ductility, and fea- tures such as crack distribution anticipated for an emulative 37 -500 -200 100 400 700 1000 1300 1600 1900 2200 2500 2800 3100 3400 -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 M ic ro st ra in Location (in) Mu 1 Mu 1.5 Mu 2 Mu 3 Mu 4 Mu 6 Mu 8 Push Pull Bent Cap North Bent Cap South Joint Yield Strain Figure 2.48. Strain profile—stirrups in bent cap (midheight) and joint (bottom), displacement control—CIP. Figure 2.49. Specimen response at a 2.6% drift ratio (µ4)—GD.

beam-column connection test. A comparison of the load- displacement envelope to the predicted envelope showed a good correlation. In addition, Figure 2.52 reveals a very simi- lar load-displacement response for the GD and CIP specimens. The dominance of ductile plastic hinging in the column and minimal damage in the capacity-protected joint and bent cap satisfied the emulation performance goal for the GD specimen. Column Displacement Decomposition. GD column displacement decomposition, summarized in Figure 2.53, confirmed the dominance of plastic hinging and showed that displacement components were reasonably determined and predictions were reasonably made. The joint shear displace- ment was minor, contributing 4.9% on average to the overall column displacement, and was consistent with visual observa- tions of minor joint cracking. Column bars were well anchored within the ducts, and although splitting cracks developed between ducts (at the top and bottom of the bent cap as tested), there was no evidence of grout splitting within ducts, initiation of pullout failure, significant bar slip or duct slip. Displacement component magnitudes and percentages for the GD and CIP specimens compared very favorably. Joint Response. As shown in Table 2.3 and Table 2.4, GD joint distress was limited and joint behavior compared very favorably with the CIP specimen. Analysis of the joint indicated that the principal tensile stress was limited to 5.1 , less than half of the 2006 LRFD RSGS (2) limit of 12 , but about 50% larger than 3.5 , the level at which more extensive (addi- tional) joint reinforcement is required for development of the assumed force transfer mechanism. Principal compressive stresses did not exceed 0.08f ′c, less than a third of the 2006 LRFD RSGS limit of 0.25f ′c. These values correspond well with the intentions of the design and the observed joint perfor- mance. The joint shear stress-strain response compared closely with the CIP, with limited joint softening evident at increasing drift ratios (see Figure 2.46). This correlated well with the max- ′fc ′fc ′fc 38 -6.1 -5.2 -4.3 -3.5 -2.6 -1.7 -0.9 0.0 0.9 1.7 2.6 3.5 4.3 5.2 6.1 -70 -60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60 70 -3.5 -3.0 -2.5 -2.0 -1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 Drift Ratio (%) La te ra l F or ce (k ip) Lateral Displacement (in) Actual Response Actual Envelope Predicted Envelope Figure 2.50. Specimen response at a 5.1% drift ratio (µ8)—GD. Figure 2.51. Lateral force versus lateral displacement—GD.

imum surface crack width in the joint region that was limited to 0.040 in, as shown in Figure 2.47, as well as displacement decomposition results. Diagonal joint crack patterns were reasonably consistent for the GD and CIP specimens, as were flexural crack patterns. Although maximum joint crack widths for the GD specimen were somewhat larger (0.040 in versus 0.025 in), they were consistent with the level of joint stresses. Minor surface spalling developed on the east face of the bent cap for GD, whereas no spalling developed for CIP. Joint deformation was very small, with maximum change in panel area limited to less than 0.2%. The GD bedding layer performed integrally with 39 -80 -70 -60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60 70 80 -3.5 -3.0 -2.5 -2.0 -1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 )pik( ecr oF lare taL Lateral Displacement (in) Actual Envelope - CIP Actual Envelope - GD Actual Envelope - CPFD Actual Envelope - CPLD Figure 2.52. Applied lateral force versus lateral displacement envelopes— all specimens. Figure 2.53. Displacement decomposition component percentages—GD. 7.6% 6.6% 4.9% 4.5% 25.7% 17.8% 13.6% 12.0% 24.7% 24.3% 24.8% 29.2% 42.0% 51.3% 56.7% 54.4% -8.1% -6.4% -6.2% -5.9% -15.9% -11.8% -11.5% -10.8% -31.1% -27.4% -21.2% -26.2% -44.8% -54.4% -61.1% -57.0% -100% -80% -60% -40% -20% 0% 20% 40% 60% 80% 100% 0.56 (FC45) 1.12 (FC55) 1.88 (µ3) 2.55 (µ4) D is pl ac em en t C om po ne nt Drift Ratio (%) (Stage) Fixed End Rotation Column Flexure Bent Cap Flexibility Joint Shear hs uP ll uP Note: Cell 1 Curvature gages removed after µ4.

the column, and crushing of the column concrete above the bedding layer confirmed the preferable condition that grout was not a weak link in the system. Similar to the CIP specimen, bent cap longitudinal bars reached only 53% of yield, even though the additional bent cap longitudinal reinforcement (0.245Ast) required by 2009 LRFD SGS was not included (1). Stirrup strain outside the joint reached 68% of yield, and, although the construction stirrups within the joint did not yield, the south construction stirrup reached 75% of yield (see Figure 2.54), indicating its contribution to the stable joint performance. The CIP specimen exhibited a similar trend of large stirrup strains (exceeding yield) within the joint. Cap Pocket Full Ductility Specimen CPFD specimen response was dominated by plastic hinging of the column adjacent to the bent cap (see Figure 2.55 and Figure 2.56), as intended by the emulative assumption in the design. Similar to the CIP specimen, the CPFD specimen exhib- ited excellent ductility to a large drift of 4.3% (nominal displace- ment ductility of 8), and load-displacement response indicated stable hysteretic behavior without appreciable strength degra- dation. Post-test inspection revealed that two column bars fractured after buckling at ultimate. Initial spalling of the column just above the bedding layer formed at a drift of 40 Table 2.4. Maximum joint response—comparison ratios for all specimens. Parameter GD/CIP CPFD/CIP CPLD/CIP CPLD/CPFD Joint Shear Stress 0.95 0.89 1.30 1.47 Principal Tensile Stress 0.94 0.88 1.30 1.47 Principal Compressive Stress 0.92 0.81 1.48 1.86 Angle of Principal Plane 1.00 0.98 1.00 1.01 Joint Rotation 1.15 0.89 1.47 1.66 Change in Panel Area 1.19 0.81 2.82 3.26 -2500 -2000 -1500 -1000 -500 0 500 1000 1500 2000 2500 -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 M ic ro st ra in Location (in) Mu 1 Mu 1.5 Mu 2 Mu 3 Mu 4 Mu 6 Mu 8 Push Bent Cap North Bent Cap South Joint Yield Strain Yield Strain Figure 2.54. Strain profile—stirrups in bent cap (midheight) and joint (bottom), displacement control—GD.

0.9% (µ1.5), with spalling much more evident at a drift of 3.2% (µ6). Significant column flexural and shear cracks and spalling developed; however, a distinctive crack pattern in the joint developed, different from that observed for the CIP spec- imen. Diagonal cracks formed above and below the corru- gated pipe through a drift of 3.1% (µ6, pull), at which stage diagonal cracks (limited to 0.009 in) passed through the cen- tral portion of the joint. Joint shear contributed only 5% to col- umn displacement, and joint shear stiffness compared closely to that of the CIP, with limited joint softening evident at increasing drift ratios. Column bars were well anchored within the pipe, with only minor bar slip. Principal tensile stresses significantly exceeded 3.5 and justified the use of additional joint reinforcement, including the pipe, for development of a force transfer mechanism. Stirrup strains within the joint reached only 25% of yield for the CPFD, but yielded for the CIP. Bent cap longitudinal bar strains exhibited a pattern similar to the CIP bottom bar, but the CPFD longitudinal bars yielded within the joint. In addi- tion, supplementary hoops that were placed at the ends of the pipe to reinforce the pipe and limit dilation and potential unraveling reached up to 52% of yield, indicating their contri- bution to joint performance. Pipe strains were limited to 37% of yield. The bedding layer appeared to perform integrally with the column, did not produce unusual behavior in the joint or specimen, and was not a weak link in the system. In addition, integral behavior between the pocket concrete, pipe, and sur- rounding concrete was evident. Column Lateral Force versus Lateral Displacement. The lateral force displacement (hysteretic) response of the CPFD column, shown in Figure 2.57, indicates stable hysteretic behavior with loops of increasing area without appreciable strength degradation, as well as stiffness, strength, ductility, and features such as crack distribution anticipated for an emu- lative beam-column connection test. A comparison of the load-displacement envelope to the predicted envelope showed a good correlation. In addition, Figure 2.52 reveals a very sim- ilar load-displacement response for the CPFD and CIP speci- mens. The dominance of ductile plastic hinging in the column and minimal damage in the capacity-protected joint and bent cap satisfied the emulation performance goal for the CPFD specimen. Column Displacement Decomposition. CPFD column displacement decomposition, summarized in Figure 2.58, confirmed the dominance of plastic hinging and showed that displacement components were reasonably determined and predictions were reasonably made. The joint shear displace- ment was minor, contributing only 4.1% to the overall column displacement, and was consistent with visual observations of minor joint cracking. Column bars were well anchored within the pipe, contributing less than 7% to fixed end rotation. Although two flexural cracks extended across the pipe, there was no evidence of concrete splitting within the pipe, initiation of pullout failure, or significant bar slip or pipe slip. Displace- ment component magnitudes and percentages for the CPFD and CIP specimens compared very favorably. ′fc 41 Figure 2.55. Specimen response at a 2.1% drift ratio (µ4)—CPFD. Figure 2.56. Specimen response at a 4.2% drift ratio (µ8)—CPFD.

Joint Response. As shown in Table 2.3 and Table 2.4, CPFD joint distress was limited and joint behavior compared very favorably with the CIP specimen. Analysis of the joint indi- cated that the principal tensile stress was limited to 4.4 , less than half of the 2006 LRFD RSGS (2) limit of 12 , but 37% larger than 3.5 , the level at which more extensive (addi-′fc ′fc ′fc tional) joint reinforcement is required according to the 2006 LRFD RSGS. Principal compressive stresses did not exceed 0.07f ′c, less than a third of the 2006 LRFD RSGS limit of 0.25f ′c. These values correspond well with the intentions of the design and the observed joint performance. Accounting for the differ- ent concrete strengths, the CPFD stresses were 11% to 19% smaller than those for CIP. The joint shear stress-strain 42 Figure 2.57. Lateral force versus lateral displacement—CPFD. Figure 2.58. Displacement decomposition component percentages—CPFD. -5.93 -5.08 -4.24 -3.39 -2.54 -1.69 -0.85 0.00 0.85 1.69 2.54 3.39 4.24 5.08 5.93 -80 -70 -60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60 70 80 -3.5 -3.0 -2.5 -2.0 -1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 Drift Ratio (%) La te ra l F or ce (k ip) Lateral Displacement (in) Actual Response Actual Envelope Predicted Envelope 6.5% 6.2% 5.8% 4.4% 3.8% 3.5% 20.2% 18.2% 15.7% 11.1% 9.5% 8.6% 13.4% 11.1% 8.9% 13.4% 15.0% 8.4% 59.9% 64.5% 69.6% 71.2% 71.7% 79.6% -7.3% -6.1% -5.3% -4.3% -3.3% -3.1% -15.1% -13.2% -11.2% -9.2% -8.0% -6.6% -33.4% -31.8% -33.9% -30.0% -24.4% -17.8% -44.3% -48.9% -49.6% -56.5% -64.3% -72.5% -100% -80% -60% -40% -20% 0% 20% 40% 60% 80% 100% 0.63 (µ1) 0.85 (µ1.5) 1.10 (µ2) 1.62 (µ3) 2.17 (µ4) 3.20 (µ6) D is pl ac em en t C om po ne nt Drift Ratio (%) (Lateral Displacement Ductility) Fixed End Rotation Column Flexure Bent Cap Flexibility Joint Shear hs uP ll uP Note: Curvature gages unreliable after µ6.

response compared closely to the CIP, with limited joint soft- ening evident at increasing drift ratios (see Figure 2.46). The maximum change in the CPFD panel area was approx- imately 20% less than that for the CIP specimen, correspon- ding with fewer diagonal cracks in the CPFD joint region and a significantly smaller maximum diagonal crack width (0.009 in) compared to the CIP joint (0.025 in). In addition, only at a 3.2% drift (µ6, pull) did diagonal cracks pass through the central portion of the CPFD joint itself. The CIP joint exhibited a more extensive pattern of diagonal cracks through the joint region for both push and pull loading. The different CPFD crack pattern and widths and strain distribution sug- gest a somewhat different load path in the joint region due to the presence of the corrugated pipe. Differences in joint behavior were also evident in strain dis- tributions. Stirrup strains within the joint reached only 25% of yield for the CPFD (see Figure 2.59), but yielded for the CIP. Bent cap longitudinal bar strains exhibited a pattern similar to the CIP bottom bar, but the CPFD longitudinal bars yielded within the joint. In addition, supplementary hoops that were placed at the ends of the pipe to reinforce the pipe and limit dilation and potential unraveling reached up to 52% of yield, indicating their contribution to joint performance. Pipe strains were largest at midheight, where principal strains were limited to 37% of yield. Cap Pocket Limited Ductility Specimen CPLD specimen response was characterized by a combina- tion of plastic hinging of the column adjacent to the bent cap and joint shear cracking and deformation (see Figure 2.60, Figure 2.61, Figure 2.47, and Figure 2.46). However, the system achieved an unexpectedly large drift ratio of 5.1% (nominal displacement ductility of 8), and load- displacement response indicated stable hysteretic behavior without appreciable strength degradation. Failure was due to buckling and fracture of two column bars rather than joint failure. These characteristics were similar to the full ductility specimens. 43 -2500 -2000 -1500 -1000 -500 0 500 1000 1500 2000 2500 -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 M ic ro st ra in Location (in) 13 kips 20 kips 30 kips 48 kips Push Pull Bent Cap North Bent Cap South Joint Yield Strain Yield Strain Figure 2.59. Strain profile—stirrups in bent cap (midheight) and joint (bottom), force control—CPFD. Figure 2.60. Specimen response at a 2.5% drift ratio (µ4)—CPLD.

Nevertheless, significant effects of joint shear associated with the SDC B limited ductility design developed, including the following: • More severe joint distress including crack widths as large as 0.080 in (with minor joint spalling); • Much softer joint shear response and large joint shear strains after an initial stiff response; • Similar initial spalling of the column at a drift of 1.2% (µ2) but delay of significant spalling and plastic hinging to a much larger drift (3.7%, µ6) due to the initial dominance of joint shear; • A flexure/shear displacement component ratio that aver- aged 2.2, nearly an order of magnitude smaller than that for the CPFD and CIP specimens (16.5 and 20.0, respectively); and • Much larger bar slip, but without loss of anchorage. Principal tensile stresses significantly exceeded 3.5 , jus- tifying the use of joint reinforcement. However, joint re- inforcement other than the corrugated pipe was not used, thus allowing joint shear cracks to open and grow without restraint. Joint crack patterns were more similar to the CIP specimen than to the CPFD specimen. Although column longitudinal bars remained anchored within the pipe, significant bar slip developed. Bottom bent cap longitudinal reinforcement exhibited a pattern similar to the CPFD, reaching yield, but pipe strains were larger for the CPLD. Column Lateral Force versus Lateral Displacement. The lateral force-lateral displacement (hysteretic) response of the CPLD column, shown in Figure 2.62, indicates stable hyste- retic behavior with loops of increasing area without appre- ciable strength degradation, as well as stiffness and strength anticipated for an emulative beam-column connection test. The level of ductility exceeds that expected for a limited ductility connection. A comparison of the load-displacement envelope to the predicted envelope—which assumed full flex- ural capacity without any limitation based on limited ductil- ity performance—showed a good correlation. In addition, ′fc 44 Figure 2.61. Specimen response at a 5.0% drift ratio (µ8)—CPLD. -5.93 -5.08 -4.24 -3.39 -2.54 -1.69 -0.85 0.00 0.85 1.69 2.54 3.39 4.24 5.08 5.93 -80 -70 -60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60 70 80 -3.5 -3.0 -2.5 -2.0 -1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 Drift Ratio (%) La te ra l F or ce (k ip) Lateral Displacement (in) Actual Response Actual Envelope Predicted Envelope Figure 2.62. Lateral force versus lateral displacement—CPLD.

Figure 2.52 reveals a very similar overall load-displacement response for the CPLD, CPFD, and CIP specimens. The eventual dominance of ductile plastic hinging in the col- umn satisfied the performance goal for the limited ductility specimen. Column Displacement Decomposition. Limited ductil- ity emulative bridge bent caps are expected to exhibit flexural plastic hinging, but also are expected to achieve a significantly lower displacement ductility capacity (in the range of µ2) due to less stringent joint and column detailing requirements (2, 1). The CPLD column detailing matched that of the CPFD design, allowing the CPLD to develop plastic hinging and exhibit large flexural displacement components at increasing drift levels. However, the less stringent SDC B joint detailing permitted more extensive joint damage to occur; thus, joint shear com- ponents were expected to contribute significantly. In agreement with visual observations (see Figure 2.47), the CPLD displacement decomposition summarized in Figure 2.63 demonstrated that the column displacement due to joint shear was nearly an order of magnitude larger than and the flexural component averaged approximately 25% less than that for the full ductility specimens (see Figure 2.64). The CPLD flexure/ shear ratio (2.2 average) was also nearly an order of magnitude smaller than for CPFD and CIP (16.5 and 20.0, respectively). In addition, the CPLD bar slip component of column dis- placement was approximately 11 times larger than the bar slip component for the CIP and CPFD specimens. A significant increase in CPLD slip toward ultimate was observed, even as the load decreased. Although this may indicate bar pullout was impending, pullout did not mobilize before column bar buck- ling failure occurred. Joint Response. As shown in Figure 2.47 and Table 2.3, the joint region for the CPLD specimen exhibited a significant level of distress that increased throughout the test. The CPLD specimen did not include 2006 LRFD RSGS (2) joint reinforce- ment (for full ductility specimens) or supplemental construc- tion stirrups in the joint. Analysis of the joint indicated that the principal tensile stress reached was 7.0 , twice the 3.5 limit at which extensive (additional) joint reinforcement is required according to the 2006 LRFD RSGS. Table 2.4 shows significantly larger joint stresses and deformation for CPLD specimens compared to the full ductility specimens. However, no joint reinforcement other than the corrugated pipe was used, which allowed joint shear cracks to open and grow without restraint. Principal compressive stresses reached 0.13f ′c, approximately half the limit of 0.25f ′c. The joint shear stress-strain response shown in Figure 2.46 shows much softer joint shear response and larger joint shear strains for CPLD specimens compared to other specimens, after initially stiff response. CPLD diagonal cracks were as wide as 0.050 in at µ2 (push, both faces), and although cracks increased to large widths at ultimate (0.070 in, east face; 0.080 in, west face), joint failure did not occur. Joint stirrups were not used for the CPLD; how- ever, CIP analysis demonstrated that construction stirrups were highly effective, reaching yield, and contributed to resist- ing joint stresses and limiting crack opening. These stirrups were less effective for the CPFD specimen, which exhibited ′fc′fc 45 Figure 2.63. Displacement decomposition component percentages—CPLD. 33.0% 30.6% 27.6% 23.8% 26.2% 20.7% 15.7% 19.3% 20.9% 20.0% 16.1% 10.4% 7.5% 7.6% 7.2% 7.7% 8.1% 13.0% 16.5% 22.3% 40.1% 40.5% 40.7% 44.3% 47.2% 46.9% 49.5% 36.6% -34.3% -35.6% -34.6% -31.4% -23.2% -22.3% -20.3% -16.3% -15.6% -13.5% -11.3% -10.2% -6.4% -1.2% -12.9% -12.2% -12.9% -13.4% -13.5% -19.6% -31.7% -36.6% -36.6% -39.0% -43.9% -53.1% -51.6% -46.8% -100% -80% -60% -40% -20% 0% 20% 40% 60% 80% 100% 0.52 (µ1) 0.79 (µ1.5) 1.09 (µ2) 1.62 (µ3) 2.47 (µ4) 3.67 (µ6) 4.96 (µ8) D is pl ac em en t C om po ne nt Drift Ratio (%) (Displacement Ductility) Fixed End Rotation Column Flexure Bent Cap Flexibility Joint Shear hs uP ll uP

more limited cracking. Although different joint crack patterns correspond to different load paths, the most important effect of the joint cracking on overall specimen response was the sig- nificant increase in joint shear displacements due to softening of the CPLD joint. Larger corrugated pipe strains developed for the CPLD, up to 70% of yield, compared to the CPFD (37% of yield), and strain distributions also differed. Strain patterns for the bent cap longitudinal bars were reasonably consistent among spec- imens, especially for the bottom bars, with both CPLD and CPFD bars yielding at the centerline. Although column lon- gitudinal bars remained anchored within the pipe, the bar slip component of column displacement was approximately 11 times that of the CIP and CPFD specimens. However, a bar anchorage equation from prior research on grout pockets indicated a larger development length for the CPLD column bars (beyond that required by 2006 LRFD RSGS) and may have helped reduce slip (7, 2). The bedding layer appeared to perform integrally with the column, did not produce unusual behavior in the joint or spec- imen, and was not a weak link in the system. 2.3.2 Nonintegral Hybrid Connections This section summarizes primary aspects of specimen response, including column hysteretic response (lateral force- displacement), displacement decomposition, and joint response. Comparisons are made between the CIP and pre- cast connections, as well as between the full and limited duc- tility specimens. Conventional Hybrid Specimen For the conventional hybrid specimen (HYB1), the primary lateral response was dominated by the localized joint rotation occurring at the bedding layer, as shown in Figure 2.65 and Figure 2.66. This specimen achieved drift ratios in excess of 6.0% with no noticeable reduction in lateral capacity. Localized spalling of concrete within the compression toe was observed related to the large strains expected with large drift ratios. Fracture of the first reinforcing bar was noted by auditory observation following two loading cycles reaching a 6.0% drift ratio. Review of the specimen indicated that appreciable buckling of the longitudinal reinforcement occurs within the compression region followed by premature fracture on the following tension cycle. Minimal damage to the column was observed outside of the base of the column with only a few minor flexural and tension cracks noted. The overall perfor- mance of the bent cap joint indicated only minor flexural cracking, and small crack widths indicated that a reliable joint design methodology was used. Column Lateral Force versus Lateral Displacement. The complete force-displacement curve obtained for this specimen is shown in Figure 2.67. The lateral force presented is the actual lateral force considering the effects of system deformation dur- ing testing. Stable lateral response is observed up to and includ- ing drift levels of 6.0%. For the loading cycles reaching 8.0% drift, a considerable drop in the lateral force resistance is observed. During testing, it was noted by auditory observation that the first reinforcing bar fractured during the push cycle to 4.0% drift ratio following the two cycles to 6.0% drift ratio. This 46 Figure 2.64. Comparison of flexural and joint shear displacement components—all specimens. 3.4% 67.7% 20.0 4.9% 72.8% 14.9 4.1% 68.3% 16.5 28.3% 63.0% 2.2 Joint Shear (Measured) Flexure (Measured) Flexure/Shear CIP GD CPFD CPLD

can be observed in the experimental data, which show the sud- den short drop in lateral force just prior to reaching a 4.0% drift ratio in the push direction. The nominal capacity calculated using the simplified analysis technique and the complete force- displacement prediction is also provided. Review of Figure 2.67 indicates that the nominal capacity predicted using the simpli- fied procedure provides a reasonable estimate of the nominal lateral capacity of the specimen. Additionally, the complete force-displacement prediction matches very well with the recorded response. The predicted failure of the section was underpredicted, indicating appreciable conservatism in the ultimate displacement capacity prediction. The force-displacement envelopes for all three hybrid spec- imens along with the CIP specimen are shown in Figure 2.68. It is apparent that all hybrid specimens have greater lateral capacity than the CIP control specimen. The larger-than- anticipated effective post-tensioning force in the conventional hybrid specimen resulted in this increase, and the other hybrid specimens were designed to be similar to the conventional hybrid specimen. Column Displacement Decomposition. Figure 2.69 pro- vides a graphical breakdown of the key components of the lateral deformation captured with instrumentation during testing. This plot provides a summary of the relative contribu- tion of a given mode of deformation as compared to the total displacement recorded at the same instant of time. This plot shows that with increasing lateral deformation, the relative contribution of end rotations increases and the relative contri- butions of column flexure and beam rotation decrease. This trend is expected as the system facilitates larger deformations through concentrated end rotations. The reduction in total dis- placement modes recorded at larger drift ratios indicates the presence of additional modes of response occurring at large drift ratios. The difference between the sum of the relative con- tributions and 100% is due to additional system deformations not explicitly isolated with instrumentation during testing. Joint Response. Observed bent cap joint damage follow- ing the testing is shown in Figure 2.70 for all hybrid specimens. Figure 2.70a shows that only minor damage occurred within the joint during the entirety of the testing. The diagonal crack- ing patterns indicate that joint shear cracking occurred but that 47 (a) (b) Figure 2.65. Specimen response at 2% drift (a) column base and (b) joint—HYB1. Figure 2.66. Specimen response at 6% drift—HYB1.

the joint reinforcement design was adequate to resist extensive crack growth and subsequent joint damage. Column Compression Strain Profile. Small-diameter, No. 2 reinforcing bars were embedded within the confined concrete core near the spiral reinforcement to try and capture the maximum confined concrete strains in the section. These bars were aimed at determining (1) the level of straining in the concrete compared to the expected failure strain and (2) the vertical distribution of strains. Results from these strain gages are shown in Figure 2.71, which shows that the maximum recorded compression strain is less than the predicted ultimate compression strain of the confined concrete core as predicted by Mander, Priestley, and Park (1988) (31). Additionally, the spread of the compression strain within the column is slightly less than the assumed distance equal to the neutral axis depth. The recorded resulting strains were less than the expected strain, which indicates that there is sectional nonlinearity at the column base, which results in a reduction in the experienced maximum straining. The assumptions presented in Improving the Design and Performance of Concrete Bridges in Seismic Regions (5) are conservative and reasonable for design but may be subject to future improvements. Residual Drift. One of the major aims of hybrid bridge systems is the reduction of residual displacements. Figure 2.72 provides a plot of the ratio of recorded residual drift to maxi- mum drift during that cycle. This plot includes data for the 48 Figure 2.67. Lateral force versus lateral displacement—HYB1. Figure 2.68. Lateral force versus lateral displacement envelopes—hybrids and CIP. Displacement, inches Drift Ratio, % La te ra l F or ce , k ip s -5 -4 -3 -2 -1 0 1 2 3 54 -8 -6 -4 -2 0 2 4 6 8 -100 -75 -50 -25 0 25 50 75 100 Nominal Capacity HYB1 Prediction Displacement, inches Drift Ratio, % La te ra l F or ce , k ip s 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 0.8 1.6 2.4 3.2 4 4.8 5.6 6.4 7.2 8 0 20 40 60 80 100 HYB1 HYB2 HYB3 Cast-in-place

49 (a) HYB1 (b) HYB2 (c) HYB3 Figure 2.69. Lateral displacement decomposition—HYB1. Figure 2.70. Joint region cracking post test—hybrid specimens. Strain, με H ei gh t A bo ve B en t C ap , i nc he s H ei gh t / C ol um n Di am et er 8 10 0.5 12 -20000 -15000 -10000 -5000 0 5000 10000 15000 20000 0 2 0.1 4 0.2 6 0.3 0.4 0.6 0 Bedding Layer Push (South Gages)Pull (North Gages) 0.50% 0.75% 1.00% 1.50% 2.00% 3.00% 4.00% 6.00% Figure 2.71. Compression strain distribution—HYB1.

three hybrid specimens as well as the CIP control specimen. Only the first cycle residual drift ratios are shown; however, the second cycle exhibited only slightly greater residual drifts. In general, for the conventional hybrid specimen the residual drift ratio increases with the applied lateral drift. However, the recorded residual drift is significantly less in comparison to the CIP specimen, indicating an overall improvement in the post- earthquake performance of the system. Concrete Filled Pipe Hybrid Specimen Similar to the conventional hybrid specimen, the primary lateral response of the concrete filled pipe specimen (HYB2) is dominated by the localized end rotations at the bedding layer, as shown in Figure 2.73 and Figure 2.74. Up to the 2.0% drift level, the overall response of the system was as anticipated. However, following the drift cycles to 2.0%, noticeable degra- dation of the grout bedding layer was observed. Deterioration continued with increasing lateral drifts. The degradation in the bedding layer resulted in a continual loss of lateral strength due to a reduction in the effective column dimension. No damage was observed in the column outside of the bedding layer. Fracture of the reinforcement was noted on 6.0% drift ratio cycles with similar observed buckling leading to fracture. The bent cap responded as anticipated and similarly to the con- ventional hybrid specimen even with the increase in lateral demand recorded. The overall performance of the bent cap joint indicated only minor flexural cracking, and small crack widths indicated a reliable joint design methodology was used. Column Lateral Force versus Lateral Displacement. The complete force-displacement curve obtained for this specimen is shown in Figure 2.75. The lateral force presented is the actual lateral force considering the effects of system deformation dur- ing testing. Hysteretic response was stable up to a 6.0% drift ratio in terms of the stability of the hysteresis loops under repeated cycles. However, loss of lateral strength was observed in both the positive and negative directions following loading cycles to a 2.0% drift. This loss in lateral strength is attributa- ble to the accumulation of damage within the grout bedding layer, which resulted in a continual decrease in the effective col- umn diameter. According to the commonly accepted defini- tion of failure as when the system lateral strength is 80% of the maximum, the concrete filled pipe specimen failed at a 5.0% drift ratio. 50 Drift Ratio, % R es id ua l / M ax im um D rif t, % 1 2 3 4 65 0 20 40 60 80 100 1.0% Residual Drift Ratio 1.5% Residual Drift Ratio 2.0% RDR HYB1 HYB2 HYB3 Cast-in-place Figure 2.72. Residual drift ratio versus applied drift ratio—three hybrid systems. (a) (b) Figure 2.73. Specimen response at 2% drift (a) column base and (b) joint—HYB2.

The nominal capacity calculated using the simplified analy- sis technique and the complete force-displacement prediction are also provided. Figure 2.75 indicates that the nominal capac- ity predicted using the simplified procedure provides a reason- able and slightly conservative estimate of the nominal lateral capacity of the specimen. Additionally, the complete force- displacement prediction matches very well with the recorded response up to the 2.0% drift level. Following the cycles to 2.0% drift, the degradation in the bedding layer was not captured by the prediction; thus, the expected lateral resistance continued to grow. The force-displacement envelopes for all three hybrid spec- imens along with the CIP specimen, are shown in Figure 2.68. Comparison of the conventional (HYB1) and concrete filled pipe (HYB2) envelopes shows the stability of the lateral resis- tance for the conventional specimen whereas a continual reduc- tion in strength is observed for the concrete filled pipe specimen. Column Displacement Decomposition. Figure 2.76 pro- vides a graphical breakdown of the key components of the lateral deformation captured with instrumentation during testing. This plot shows that with increasing lateral defor- mation, the relative contribution of end rotations increases and the relative contributions of column flexure and beam rotation decrease. This trend is expected as the system facili- tates larger deformations through concentrated end rotations. The reduction in total displacement modes recorded at larger drift ratios indicates the presence of additional modes of response occurring at large drift ratios. The difference between the sum of the relative contributions and 100% is due to addi- tional system deformations not explicitly isolated with instru- mentation during testing. It is noted that an appreciable amount of deformation was not captured during the lower level loading cycles. Joint Response. Observed bent cap joint damage follow- ing testing of the concrete filled pipe hybrid specimen is shown in Figure 2.70b. Figure 2.70b indicates that only minor damage occurred within the joint during the entirety of the testing, sim- ilar to what was observed in the conventional specimen. The level of observed damage is also of a similar magnitude even though the lateral demands, and therefore joint demands, were greater for this specimen. Diagonal cracking patterns indicate that joint shear cracking occurred, but the joint reinforcement design was adequate to resist extensive crack growth and sub- sequent joint damage. Residual Drift. Review of Figure 2.72 shows the ratio of residual drift to maximum drift during that cycle for this spec- imen. The observed residual drift for this specimen is greater than that recorded for the conventional hybrid specimen, 51 Figure 2.74. Specimen response at 6% drift—HYB2. Displacement, inches Drift Ratio, % La te ra l F or ce , k ip s -5 -4 -3 -2 -1 0 1 2 3 54 -8 -6 -4 -2 0 2 4 6 8 -100 -75 -50 -25 0 25 50 75 100 Nominal Capacity HYB2 Prediction Figure 2.75. Lateral force versus lateral displacement—HYB2.

resulting from the increased damage in the bedding layer dur- ing this specimen’s testing. Similar to the conventional hybrid specimen, only slightly greater residual drifts were recorded during the second cycle to a given drift. Even though the resid- ual drifts were greater than those of the conventional hybrid specimen, the recorded residual drift was significantly less than the residual drift of the CIP specimen, indicating an over- all improvement in the post-earthquake performance of the system. Dual Steel Shell Hybrid Specimen Similar to concrete filled pipe hybrid specimen, the primary lateral response of the dual steel shell hybrid specimen was dominated by the localized end rotations at the bedding layer, as shown in Figure 2.77 and Figure 2.78. Up to the 2.0% drift level, the overall response of the system was as anticipated. However, similar to the concrete filled pipe hybrid specimen, following the drift cycles to 2.0%, noticeable degradation of the grout bedding layer was observed, with deterioration continu- ing with increasing lateral drifts. The degradation in the bed- ding layer resulted in a continual loss of lateral strength due to a reduction in the effective column dimension. No damage was observed in the column outside of the bedding layer. Fracture of the reinforcement was noted on 6.0% drift ratio cycles with similar observed buckling leading to fracture. The bent cap responded as anticipated even with the increase in lateral demand recorded, similar to the conventional hybrid speci- 52 Figure 2.76. Lateral displacement decomposition—HYB2. (a) (b) Figure 2.77. Specimen response at 2% drift (a) column base and (b) joint—HYB3.

men. The overall performance of the bent cap joint indicated only minor flexural cracking with small crack widths indicat- ing that a reliable joint design methodology was used. The overall condition of the bedding layer following testing is shown in Figure 2.79. The post-test consistency of much of the bedding layer grout was a very fine material indicating sig- nificant crushing and degradation of the grout matrix. The specimen was also observed to have decreased in overall height following seismic testing due to the reduction in bedding layer thickness associated with a reduction in the bearing area of the grout. Column Lateral Force versus Lateral Displacement. The complete force-displacement curve obtained for this specimen is shown in Figure 2.80. The lateral force presented is the actual lateral force considering the effects of system deformation dur- ing testing. Hysteretic response was stable up to a 4.0% drift ratio in terms of stability of the hysteresis loops under repeated cycles. However, loss of lateral strength was observed in both the positive and negative directions following loading cycles to 2.0% drift. This loss in lateral strength is attributable to the accumulation of damage within the grout bedding layer, which resulted in a continual decrease in the effective column diam- eter. Considering the commonly accepted practice that failure is defined when the system lateral strength is 80% of the max- imum, the dual steel shell hybrid specimen is said to have failed at 5.0% drift ratio. The nominal capacity calculated using the simplified analy- sis technique and the complete force-displacement prediction is also provided. Review of Figure 2.80 indicates that the nominal capacity predicted using the simplified procedure provides a reasonable and slightly conservative estimate of 53 Figure 2.78. Specimen response at 6% drift—HYB3. Figure 2.79. Bedding layer grout deterioration at end of test—HYB3. Displacement, inches Drift Ratio, % -5 -4 -3 -2 -1 0 1 2 3 4 5 -8 -6 -4 -2 0 2 4 6 8 Nominal Capacity HYB3 Prediction La te ra l F or ce , k ip s -100 -75 -50 -25 0 25 50 75 100 Figure 2.80. Lateral force versus lateral displacement—HYB3.

the nominal lateral capacity of the specimen. Additionally, the complete force-displacement prediction matches very well with the recorded response up to the 2.0% drift level. Following the cycles to 2.0% drift, the degradation in the bedding layer was not captured by the prediction, thus the expected lateral resis- tance continued to grow. The force-displacement envelopes for all three hybrid spec- imens along with the CIP specimen are shown in Figure 2.68. Comparison of the conventional and dual shell envelopes shows the stability of the lateral resistance for the conventional specimen. A continual reduction in strength is observed for both the dual shell specimen and the concrete filled pipe specimen. Column Displacement Decomposition. Figure 2.81 pro- vides a graphical breakdown of the key components of lateral deformation captured with instrumentation during testing. From this plot, it can be seen that with increasing lateral defor- mation, the relative contribution of end rotations increases as the relative contribution of column flexure and beam rotation decreases. This trend is expected because the system facilitates larger deformations through concentrated end rotations. The reduction in total displacement modes recorded at larger drift ratios indicates the presence of additional modes of response occurring at large drift ratios. The difference between the sum of the relative contributions and 100% is due to additional sys- tem deformations not explicitly isolated with instrumentation during testing. It is noted that an increased amount of error accumulated during the testing, which resulted in the greatest amount of error at the end of testing. Bedding Layer Response. As was mentioned in the gen- eral summary of the specimen response, the overall dimension of the bedding layer reduced during testing. This bedding layer deformation was captured using the lower curvature cages shown in Figure 2.82. The growth of the bedding layer compared with the lateral deformation followed a linear rela- tionship of centroid joint growth during lateral loading and zero displacement upon return to zero drift up to the 3% drift cycles. Following this point, a noticeable reduction in stiffness of the column growth versus drift was observed. In addition, following this drift level, a continual reduction in the overall dimension was observed as the column passed through the zero drift point. This loss in bedding layer dimension also resulted in a loss of effective post-tensioning force due to a reduction in the length of tendon. This loss in effective tendon force also contributed to the continual reduction in lateral capacity of the specimen. Joint Response. Observed bent cap joint damage follow- ing the testing is shown in Figure 2.70c. Review of this figure indicates that only minor damage occurred within the joint during the entirety of the testing, similar to what was observed in the other hybrid specimens. The level of observed dam- age is of a similar magnitude as the conventional hybrid specimen even though the lateral demands, and therefore joint demands, were greater for this specimen. Diagonal cracking patterns are observed, indicating that joint shear cracking occurred but that the joint reinforcement design was adequate to resist extensive crack growth and sub- sequent joint damage. Residual Drift. Review of Figure 2.72 shows the ratio of residual drift to maximum drift during that cycle for this spec- imen. The observed residual drift for the dual steel shell hybrid specimen is similar to that observed for the concrete filled pipe hybrid specimen, which was greater than that recorded for the conventional hybrid specimen. This increase compared to the 54 Figure 2.81. Lateral displacement decomposition—HYB3.

conventional hybrid specimen is attributable to the increased damage in the bedding layer during the dual steel shell hybrid specimen’s testing. Similar to the conventional hybrid speci- men, only slightly greater residual drifts were recorded during the second cycle to a given drift. Even though the residual drifts are greater than those of the conventional hybrid specimen, in comparison to the CIP specimen, the recorded residual drift is significantly less, indicating an overall improvement in the post-earthquake performance of the system. 2.3.3 Integral Connection The integral experimental specimen (INT) was subjected to a combination of elastic loading cycles and simulated seismic loadings. These loadings were developed to apply flexural demands nearing the anticipated point of nonlinearity in the negative flexural response. At this level, distributed cracking with crack widths less than 0.005 inches was evident. The over- all response was characterized as essentially elastic, with no noticeable accumulation of seismic damage. Seismic loading cycles subjected the girder to positive and negative flexural demands. Photographic records of certain loading cycles are shown in Figures 2.83 through 2.86. In the negative loading cycles, flexural response was representative of traditional CIP superstructure response. A defined compres- sion fan was observed at the girder web at the end with the sta- bilization of cracking at 45 deg, a distance approximately equal to the superstructure depth. Distributed flexural cracking was observed within the deck with a larger crack width observed at the girder to reaction block joint. During increasing levels of seismic loading, the crack in the deck at the joint separated into two cracks a couple of inches apart. The lack of continu- ous reinforcement extending from the girder into the reaction block resulted in the observed concentrated opening at the joint during negative flexure; however, the presence of the deck flexural reinforcement served to reduce the concentra- tion of cracking within the deck. During positive loading cycles, flexural cracking was con- centrated at the girder to reaction block joint. Essentially elas- tic response was observed within the section up to the point of joint opening. As the joint began to open, the concentrated rotations about the end resulted in a reduction in the positive flexural stiffness; however, the increase in flexural resistance continued. During reversed cycling, the fiber-reinforced clo- sure joint performed well, with no observed reduction in 55 Displacement, inches Drift Ratio, % Be dd in g la ye r g ro wt h, in ch es -5 -4 -3 -2 -1 0 1 2 3 4 5 -8 -6 -4 -2 0 2 4 6 8 -0.4 -0.3 -0.2 -0.1 0 0.1 0.2 Figure 2.82. Bedding layer centerline axial deformation— HYB3. Figure 2.83. Girder end block region at 0.29% joint rotation.

joint integrity. Furthermore, at large rotation cycles, initial spalling of concrete in the bottom flange was observed with no observed damage to the joint, an indication of the excep- tional joint performance. During loading cycles past about a 0.6% joint rotation, a horizontal crack was observed between the top flange of the girder and the deck, as shown in Figure 2.85. Subsequent load- ing cycles caused a continued increase in the dimension of this crack, ultimately leading to a reduction in shear stiffness across the joint. This reduction in stiffness resulted in the slip between the girder and reaction block at large rotations, as shown in Figure 2.86b. The shear slip was caused by inadequately developed shear reinforcement within the girder end when subjected to flexural joint opening. Although a reduction of stiffness, and therefore an increase in shear slip, was observed, the ability to resist the applied seismic shear was not reduced. Moment versus Rotation Response The complete moment-rotation hysteretic response is shown in Figure 2.87. This plot indicates that there is appre- ciable energy dissipation capacity in the negative flexural direction with significantly less in the positive direction. This response characteristic is expected because the negative flex- ural direction has a significantly greater amount of mild rein- forcement present, which is expected to yield and dissipate seismic energy under increasing load cycles. Under increasing 56 Figure 2.84. Girder bottom flange joint opening at 0.19% joint rotation. Figure 2.85. Girder to deck interface crack at 0.79% joint rotation. (a) (b) Figure 2.86. (a) Bottom of closure joint and (b) shear slip at 1.03% joint rotation.

levels of rotation demand at the joint, a noticeable reduction in the negative flexural stiffness is observed. This is caused by the yielding of mild reinforcement in the concrete deck, which decreases the effective stiffness of the reinforcement. In the positive flexural direction, the reduction in post-yield stiffness under increasing cycles is not as significant as in the negative direction. The moment-rotation predicted envelope is also shown in Figure 2.87. The predicted response shows good agreement with the recorded results assuming an effective plastic hinge length equal to one-half times the superstructure depth includ- ing deck. Although the envelope captures the inelastic response with accuracy, the ultimate rotation capacity is over-predicted. The observed failure of the system occurred at approxi- mately 1.3% drift in both the positive and negative directions. However, the predicted failures in the positive and negative directions were at joint rotations equal to 1.46% and −1.69%, respectively. The error in ultimate rotation is approximately 12% in the positive direction and 30% in the negative direc- tion. Both the prediction and observed failure were controlled by fracture of the post-tensioning tendons. The failure strain in the post-tensioning tendon was equal to 0.03 in/in, per the 2009 LRFD SGS (1). The over-estimation of the ultimate rota- tion is caused by the observed kinking action in the tendon due to shear slip under large rotations. The recommended modifi- cation to the shear reinforcement detailing at the girder end is expected to alleviate much of this issue and thus result in an increase in the ultimate rotation capacity of the connection. Even with the reduction in ultimate rotation capacity due to the kinking action, the ultimate rotation capacity results in a system that can safely undergo relative settlements between adjacent bent caps in excess of 1 ft for a structure 100 ft long. This level of geometric demand is greater than would be expected in a properly designed bridge structure. The simplified nominal section capacity is also shown on the moment-rotation plots. This capacity prediction provides a relatively accurate prediction of the nominal capacity in both positive and negative directions. The negative flexural capacity was predicted using standard design equations in the fifth edi- tion of the AASHTO LRFD Bridge Design Specifications. This calculated capacity shows excellent agreement with the capac- ity determined using a strain compatibility method. For the positive flexural direction, capacity was calculated using a moment-curvature program that considers strain compati- bility across the section. The decision to use a strain com- patibility approach is due to the presence of unstressed post-tensioning in the bottom of the girder. In addition, it was observed that the moment-rotation prediction is highly sensitive to the tensile strength of the concrete, which is not accounted for in traditional design equations. While the use of simplified capacity equations for positive flexural capacity will be conservative, it is recommended to also perform a capacity calculation using strain compatibility to determine a better estimate of the connection capacity. The recorded moment-rotation response at the joint is shown in Figure 2.88 for the 100 cycles of elastic loading. This response indicates that there is no noticeable degradation in stiffness or strength within this loading range. These loading cycles confirm the elastic response of the joint region when subjected to loading within the service load range. Figure 2.88 also overlays the elastic loading cyclic response over the lower level seismic response to provide a visual comparison of the relative elastic demand compared with the section capacity. The joint rotation in these plots is based on a zero rotation at the beginning of elastic loading and does not include the orig- inal rotation imposed during the application of simulated dead loading. The moment-rotation predication is also shown in Figure 2.88. The predicted response indicates the system was 57 Joint Rotation, % M om en t, ki p- ft -1.5 -1 -0.5 0 0.5 1 1.5 -1200 -800 -400 0 400 800 Nominal Capacity Figure 2.87. Moment versus rotation response.

58 Joint Rotation, % M om en t, ki p- ft -0.5 -0.4 -0.3 -0.2 -0.1 0 0.1 0.2 0.3 0.4 0.5 -1200 -800 -400 0 400 800 Nominal Capacity Seismic Loading Elastic Cycles Figure 2.88. Moment versus rotation response at low level seismic and elastic loading. Joint Rotation, % G ird er S he ar S lip , i nc he s -1.5 -1 -0.5 0 0.5 1 1.5 -0.1 0 0.1 0.2 0.3 0.4 0.5 0.6 Figure 2.89. Recorded girder shear slip during seismic loading. loaded in the negative direction just prior to a predicted reduc- tion in the stiffness of the system. Girder Shear Slip Figure 2.89 shows the recorded girder shear slip history during all loading stages. Results from this loading indicate that the maximum relative slip between the girder and reac- tion block is less than four-hundredths of an inch for the entirety of the elastic loading cycles. Interestingly, these results indicate that during the elastic loading cycles, the girder also slipped upwards during many cycles. This recorded response does not match the expected response as downward shear loading is applied to the system during all stages. The relatively minor differential movement between the girder and reaction block is not considered a significant response characteristic and is not expected to cause adverse impacts in structural response or functionality of a bridge structure. All loading cycles below approximately −0.6% joint rota- tion have less than five-hundredths of an inch slip. As applied joint rotations increased, the recorded drifts continued to increase. Review of the recorded results indicate that during the larger joint rotation cycles, the positive loading cycles have less slip than the negative cycles. This trend is expected due to the decrease in applied shear loading during the positive cycles. Observations made during testing indicate a significant portion of the observed shear slip is due to the separation between the girder and the deck. This separation is caused by

the inadequately anchored shear reinforcement in the girder, which cannot develop the required shear within the deck. The use of headed reinforcing bars is expected to greatly reduce the observed shear slip by fully anchoring the shear reinforcement within the reinforced concrete deck. As shown in Figure 2.90, the observed horizontal cracking between the deck and girder provided a length of embedded shear reinforcement that was less than the required develop- ment length. Therefore, although the shear strength of the sys- tem was maintained, there was continued slip of the bar during repeated cycles of testing. To mitigate this issue, the use of well- anchored shear reinforcement in the deck is recommended. 2.4 Analytical Results 2.4.1 Nonintegral Hybrid Connections A series of analyses was conducted on the hybrid specimens in order to assess the adequacy of these systems for implemen- tation in seismic regions. The first set of analyses relates to the adequacy of the presented simplified and complete prediction methodologies that are discussed in more detail in Tobolski 2010 (5). The second set of analyses relates to the investigation of the potential inelastic displacement demands for hybrid systems. Analysis Prediction Methodologies The design and implementation of hybrid systems relies heavily on the ability to predict the response of these systems. The lateral force-displacement response of the hybrid mem- bers is provided in Figure 2.67, Figure 2.75 and Figure 2.80. Each of these figures also includes the lateral force-displacement envelope prediction and the predicted nominal yield demand using the simplified procedure. For the conventional hybrid specimen, both the complete and simple prediction methods provide very good agreement with the recorded response from experimental testing. The lateral displacement capacity for this specimen was underpredicted due to the conservative estimate of experienced maximum concrete strain. This agreement indicates that the prediction methods shown in the attachments to this report are adequate for the implementation of the conventional hybrid detail. For the concrete filled pipe and dual shell specimens, the prediction methods used provide good agreement with the observed response up to a lateral drift ratio of 2.0%. Above this level of lateral demand, the observed response had continual reduction in lateral capacity due to grout bedding layer degradation. These details still provided acceptable lateral response up to a 5% drift ratio when the lat- eral capacity approached 80% of the maximum recorded capacity. Within the realm of design demands, the prediction methodology is reasonable and conservative. Future work is required to verify the benefits of modifications to the grout bedding layer for improving performance of the second and third hybrid specimens. Nonlinear Time History Analyses In presenting a new structural system for use in seismic regions, the potential implications of realized displacement demands during strong ground shaking must be investigated. A series of nonlinear time history analyses were conducted on a calibrated model to determine the level of displacement ampli- fication in inelastic systems as compared to similar elastic sys- tems. The results from the conventional hybrid specimen test were used to calibrate a lumped plasticity model for dynamic analysis. A comparison between the recorded experimental results and the calibrated model is shown in Figure 2.91. The calibrated model was developed in the analysis package RUAUMOKO (32) by combining a modified Takeda model and a bilinear elastic model. The input parameters were based on the response predictions, including the relative contribution of the post-tensioning and conventional reinforcement, and then finetuned based on quasi-static simulations in the analysis model. The nonlinear time history analyses were performed for records developed for Site Class B, C, and D. A total of 30 ground motions recorded from California earthquakes were modified using the wavelet modification program WAVGEN (33). Each record was modified to be consistent with a speci- fied design spectrum developed in accordance with the 2009 LRFD SGS (1). The records were manipulated and developed for each of the site classes, resulting in a total of 90 spectrum compatible records. The resulting response spectra for each site class, in addition to the actual modified response spectra, are shown in Figure 2.92. Review of Figure 2.92 indicates that the achieved mean response spectrum for each site class matches well with the target spectrum with variability between actual time history records. 59 V C T V M < ld Tested Detail Recommended Detail Figure 2.90. Girder shear slip mechanism.

Nonlinear time history analyses were conducted using the calibrated lumped plasticity model for periods ranging from 0.1 sec to 3.0 sec at 0.1 sec intervals. Specified viscous damping was equal to 5% using tangent stiffness damping. The analyses were performed for inelastic force reduction factors ranging from 2 to 6 for single degree-of-freedom systems. The initial runs were performed with elastic response in order to deter- mine the expected yield force in the system. Considering the multiple site classes, earthquakes, and force reduction factors, a total of 13,500 analyses were performed. The results of the analysis for Site Class D are presented in Figure 2.93, with individual “x” marks representing a single inelastic displacement modification factor from a specific earthquake. The line labeled “HYB Mean” represents the mean response parameters over a range of periods. Additionally, a plot titled “EPT Mean” is presented that represents the results from a similar series of analyses conducted on elastic-plastic single degree-of-freedom oscillators. The hatched region on the plots represents the region in which the experienced dis- placement demand results in ductility values in excess of the maximum code limit of 6. Results from these analyses indi- cate that the hybrid systems investigated have displacement demands similar to those of more conventional systems. Thus, these systems are not expected to experience displace- ment demands significantly greater than those experienced by CIP or emulative systems, and provisions published in the code for these systems can be used for hybrid systems with a similar level of safety. For all systems, the overall trend observed is that the mean inelastic displacement factor approaches unity as the period approaches infinity. This trend agrees with the commonly accepted equal displacement principle (34). 60 Displacement, inches Drift Ratio, % La te ra l F or ce , k ip s -2 -1.5 -1 -0.5 0 0.5 1 1.5 2 -3 -2 -1 0 1 2 3 -100 -75 -50 -25 0 25 50 75 100 Experimental Analytical Figure 2.91. Comparison of analytical and experimental hysteresis (5). Figure 2.92. Acceleration response spectra (5).

61 Period, seconds In el as tic d isp la ce m en t f ac to r, C R 0.1 0.2 0.3 0.5 0.7 1 2 3 3 3 3 0 2 4 6 8 10 R = 2 - Site Class D HYB Mean EPT Mean μD > 6 Period, seconds In el as tic d isp la ce m en t f ac to r, C R 0.1 0.2 0.3 0.5 0.7 1 2 0 2 4 6 8 10 R = 3 - Site Class D HYB Mean EPT Mean μD > 6 Period, seconds In el as tic d isp la ce m en t f ac to r, C R 0.1 0.2 0.3 0.5 0.7 1 2 0 2 4 6 8 10 R = 4 - Site Class D HYB Mean EPT Mean μD > 6 Period, seconds In el as tic d isp la ce m en t f ac to r, C R 0.1 0.2 0.3 0.5 0.7 1 2 0 2 4 6 8 10 R = 6 - Site Class D HYB Mean EPT Mean μD > 6 Figure 2.93. Hybrid system inelastic displacement modification factor (5).

Next: Chapter 3 - Interpretation, Appraisal, and Applications »
Development of a Precast Bent Cap System for Seismic Regions Get This Book
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TRB’s National Cooperative Highway Research Program (NCHRP) Report 681: Development of a Precast Bent Cap System for Seismic Regions explores the development and validation of precast concrete bent cap systems for use throughout the nation’s seismic regions.

The report also includes a series of recommended updates to the American Association of State Highway and Transportation Officials (AASHTO) Load and Resistance Factor Design (LRFD) Bridge Design Specifications, Guide Specification for LRFD Seismic Bridge Design, and AASHTO LRFD Bridge Construction Specifications that will provide safe and reliable seismic resistance in a cost-effective, durable, and constructible manner.

A number of deliverables are provided as attachments to NCHRP Report 681, including design flow charts, design examples, example connection details, specimen drawings, specimen test reports, and an implementation plan from the research agency’s final report. These attachments, which are only available online, are titled as follows:

Attachment DS—Design Specifications

Attachment DE—Design Examples

Attachment CS—Construction Specifications

Attachment ECD—Example Connection Details

Attachment SD —Specimen Drawings

Attachment TR—Test Reports

Attachment CPT—Corrugated Pipe Thickness

Attachment IP—NCHRP 12-74 Implementation Plan

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