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33 Figure 2.41. Integral specimen external instrumentation plan. reinforcement, girder longitudinal reinforcement at the base, mary of joint cracking, principal stresses, joint deformation, and girder shear reinforcement. and strain records. 2.3 Test Results Cast-in-Place Specimen This section summarizes key aspects of specimen response CIP specimen response was dominated by plastic hinging of for the emulative, hybrid and integral experimental tests. the column adjacent to the bent cap, as shown in Figure 2.42 Detailed results are provided by Matsumoto (30, 21, 22, 23, 26) and Figure 2.43. The specimen exhibited excellent ductility to and Tobolski (5). a large drift of 5.9% (nominal displacement ductility of 10), In reporting specimen response, displacement ductility () and the load-displacement response indicated stable hysteretic and drift ratio are both used. The drift ratio is the column dis- behavior without appreciable strength degradation. Post-test placement divided by the column height and is reported as a inspection revealed that the core remained primarily intact with several column bars buckling and fracturing at ultimate. percentage. This is a more consistent basis for comparison of Initial spalling of the column occurred at 1.8% drift (3), with specimen response than displacement ductility. However, sys- progressive spalling at higher drifts. In contrast to significant 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 Figure 2.42. Specimen response at a 2.3% drift ratio experimental measurements. Joint response includes a sum- (4)--CIP.

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34 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 Figure 2.43. Specimen response at a 5.9% drift ratio specimen. Thus, the specimen provided an appropriate base- (10)--CIP. line for comparison with the precast specimens. Column Displacement Decomposition. Column dis- column flexural and shear cracks and spalling associated with placement decomposition, summarized in Figure 2.45, increasing lateral force, relatively minor cracking occurred in confirmed the dominance of plastic hinging and showed that the joint region (0.025 in maximum) without spalling. This displacement components were reasonably determined and corresponded to stiff joint shear response and limited soften- predictions were reasonably made. The joint shear displace- ing, with the contribution of joint shear to column displace- ment was minor, contributing only 3.4% on average to the ment averaging 3.4%. Principal tensile stresses significantly overall column displacement, and was consistent with visual exceeded 3.5 fc psi and justified the use of additional joint observations of minor joint cracking. Splitting cracks formed Drift Ratio (%) -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 Lateral Force (kip) 20 Maximum actuator 10 pull stroke reached 0 -10 -20 -30 -40 Actual Response -50 Actual Envelope -60 Predicted Response -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 Lateral Displacement (in) Figure 2.44. Lateral force versus lateral displacement--CIP.

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35 100% Fixed End Rotation Pull Column Flexure Pull 80% Bent Cap Flexure Pull Joint Shear Pull 51.6% 60.9% 61.1% 60.5% 63.6% 62.9% 65.7% 66.6% 60% Displacement Component 40% 20.9% 18.3% 20.4% 25.0% Push 21.9% 26.4% 28.1% 20% 18.4% 24.7% 16.2% 14.3% 10.9% 9.0% 6.3% 5.9% 9.1% 7.0% 4.6% 4.2% 3.7% 3.4% 3.0% 2.8% 2.8% 0% -2.4% -7.6% -5.2% -4.0% -3.2% -2.8% -2.3% -8.0% -16.1% -8.4% -6.3% -5.4% -4.5% -11.0% Pull -15.8% -18.9% -20% -28.1% -31.2% -34.3% -28.1% -26.5% -27.1% -22.4% -40% -23.5% -60% -59.4% -62.5% -60.6% -58.9% -55.1% -54.2% -56.7% -49.5% -80% -100% 0.69 0.93 1.19 1.76 2.34 3.60 4.59 5.88 (FC48) (1.5) (2) (3) (4) (6) (8) (10) Drift Ratio (%) (Displacement Ductility) Figure 2.45. Displacement decomposition component percentages--CIP. in the bent cap and column, and the top surface of the bent cap not exceed 0.09f c, less than a third of the 2006 LRFD RSGS (as tested) exhibited splitting cracks and local spalling; how- limit of 0.25f c. These values correspond well with the inten- ever, column bars were well anchored within the joint, with bar tions of the design and the observed joint performance. The slip contributing less than 4% on average to fixed end rotation. joint shear stress-strain response was appropriately stiff and exhibited minor softening at increasing drift (see envelope in Joint Response. As shown in Table 2.3, CIP joint distress Figure 2.46). This correlated well with the maximum surface was limited. Analysis of the joint indicated that the principal 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 tensile stress was limited to 5.4 fc psi, less than half of the displacement decomposition results. Joint deformation was 2006 LRFD RSGS (2) limit of 12 fc psi, but about 50% larger very small, with maximum change in panel area limited to less than 0.2%. Bent cap longitudinal bars did not yield, reaching than 3.5 fc psi, the level at which more extensive (additional) only 46% of yield, even though additional bent cap longitudi- joint reinforcement is required for development of the assumed nal reinforcement (0.245Ast) required by 2009 LRFD SGS was force transfer mechanism. Principal compressive stresses did not included (1). Stirrup strain outside the joint remained well Table 2.3. Maximum joint response--all specimens. Parameter CIP GD CPFD CPLD 328 312 323 371 Joint Shear Stress (psi) (4.86 ) (4.62 ) (4.31 ) (6.32 ) Principal Tensile 363 343 356 411 (psi) Stress (5.38 ) (5.09 ) (4.75 ) (6.99 ) Principal 401 370 398 460 (psi) Compressive Stress (0.088 ) (0.081 ) (0.071 ) (0.13 ) Angle of Principal (deg) 45.0 45.0 44.2 44.8 Plane -3 -3 -3 Joint Rotation (rad) 1.95 10 2.25 10 1.73 10 2.87 10-3 Change in Panel (%) 0.16 0.19 0.13 0.46 Area

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36 Joint Shear Strain (deg) -0.86 -0.57 -0.29 0.00 0.29 0.57 0.86 500 400 300 Joint Shear Stress (psi) 200 100 0 -100 -200 CIP -300 GD CPFD -400 CPLD -500 -0.015 -0.010 -0.005 0.000 0.005 0.010 0.015 Joint Shear Strain (rad) Figure 2.46. Joint shear stress versus joint shear strain envelopes--all specimens. (a) East Side--CPFD (b) West Side--CPFD (c) East Side--CPLD (d) West Side--CPLD (e) East Side--CIP (f) West Side--CIP (g) East Side--GD (h) West Side--GD Figure 2.47. Joint region cracking post test--emulative specimens.

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37 3400 Bent Cap Joint Bent Cap 3100 South North 2800 Push 2500 Pull 2200 Yield Strain 1900 Microstrain Mu 1 Mu 1.5 1600 Mu 2 Mu 3 Mu 4 Mu 6 1300 Mu 8 1000 700 400 100 -200 -500 -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 Location (in) Figure 2.48. Strain profile--stirrups in bent cap (midheight) and joint (bottom), displacement control--CIP. below yield, but the north construction stirrup within the joint Column Lateral Force versus Lateral Displacement. The yielded, as shown in Figure 2.48, indicating its contribution to lateral force displacement (hysteretic) response of the GD col- the stable joint performance. umn, shown in Figure 2.51, indicates stable hysteretic behav- ior with loops of increasing area without appreciable strength Grouted Duct Specimen degradation, as well as stiffness, strength, ductility, and fea- tures such as crack distribution anticipated for an emulative 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 (1.5), 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 fc 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, Figure 2.49. Specimen response at a 2.6% drift ratio indicating its contribution to the stable joint performance. (4)--GD.

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38 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 Figure 2.50. Specimen response at a 5.1% drift ratio favorably with the CIP specimen. Analysis of the joint indicated (8)--GD. that the principal tensile stress was limited to 5.1 fc , less than half of the 2006 LRFD RSGS (2) limit of 12 fc , but about 50% beam-column connection test. A comparison of the load- larger than 3.5 fc , the level at which more extensive (addi- displacement envelope to the predicted envelope showed a good correlation. In addition, Figure 2.52 reveals a very simi- tional) joint reinforcement is required for development of the lar load-displacement response for the GD and CIP specimens. assumed force transfer mechanism. Principal compressive The dominance of ductile plastic hinging in the column and stresses did not exceed 0.08f c, less than a third of the 2006 minimal damage in the capacity-protected joint and bent cap LRFD RSGS limit of 0.25f c . These values correspond well with satisfied the emulation performance goal for the GD specimen. the intentions of the design and the observed joint perfor- mance. The joint shear stress-strain response compared closely Column Displacement Decomposition. GD column with the CIP, with limited joint softening evident at increasing displacement decomposition, summarized in Figure 2.53, drift ratios (see Figure 2.46). This correlated well with the max- Drift Ratio (%) -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 Lateral Force (kip) 20 10 0 -10 -20 -30 -40 Actual Response -50 Actual Envelope -60 Predicted Envelope -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 Lateral Displacement (in) Figure 2.51. Lateral force versus lateral displacement--GD.

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39 80 70 60 50 40 30 20 10 Lateral Force (kip) 0 -10 -20 Actual Envelope - CIP -30 -40 Actual Envelope - GD -50 Actual Envelope - CPFD -60 -70 Actual Envelope - CPLD -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 Lateral Displacement (in) Figure 2.52. Applied lateral force versus lateral displacement envelopes-- all specimens. imum surface crack width in the joint region that was limited were somewhat larger (0.040 in versus 0.025 in), they were to 0.040 in, as shown in Figure 2.47, as well as displacement consistent with the level of joint stresses. Minor surface decomposition results. spalling developed on the east face of the bent cap for GD, Diagonal joint crack patterns were reasonably consistent for whereas no spalling developed for CIP. Joint deformation was the GD and CIP specimens, as were flexural crack patterns. very small, with maximum change in panel area limited to less Although maximum joint crack widths for the GD specimen than 0.2%. The GD bedding layer performed integrally with 100% Fixed End Rotation Column Flexure 80% 42.0% Bent Cap Flexibility Joint Shear 51.3% 56.7% 54.4% 60% Displacement Component 24.7% 40% 24.3% Push 24.8% 29.2% 20% 25.7% 17.8% 13.6% 12.0% 7.6% 6.6% 4.9% 4.5% 0% -6.4% -5.9% -8.1% -6.2% Pull -11.8% -11.5% -10.8% -15.9% -20% -21.2% -26.2% -27.4% -40% -31.1% -60% -57.0% -61.1% -54.4% -44.8% -80% Note: Cell 1 Curvature gages removed after 4. -100% 0.56 1.12 1.88 2.55 (FC45) (FC55) (3) (4) Drift Ratio (%) (Stage) Figure 2.53. Displacement decomposition component percentages--GD.

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

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41 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 fc 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 Figure 2.55. Specimen response at a 2.1% drift ratio specimen, and was not a weak link in the system. In addition, (4)--CPFD. integral behavior between the pocket concrete, pipe, and sur- rounding concrete was evident. 0.9% (1.5), with spalling much more evident at a drift of Column Lateral Force versus Lateral Displacement. The 3.2% (6). Significant column flexural and shear cracks and lateral force displacement (hysteretic) response of the CPFD spalling developed; however, a distinctive crack pattern in the column, shown in Figure 2.57, indicates stable hysteretic joint developed, different from that observed for the CIP spec- behavior with loops of increasing area without appreciable imen. Diagonal cracks formed above and below the corru- strength degradation, as well as stiffness, strength, ductility, gated pipe through a drift of 3.1% (6, pull), at which stage and features such as crack distribution anticipated for an emu- diagonal cracks (limited to 0.009 in) passed through the cen- 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- Figure 2.56. Specimen response at a 4.2% drift ratio ment component magnitudes and percentages for the CPFD (8)--CPFD. and CIP specimens compared very favorably.

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42 Drift Ratio (%) -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 Lateral Force (kip) 20 10 0 -10 -20 -30 -40 Actual Response -50 Actual Envelope -60 -70 Predicted Envelope -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 Lateral Displacement (in) Figure 2.57. Lateral force versus lateral displacement--CPFD. Joint Response. As shown in Table 2.3 and Table 2.4, tional) joint reinforcement is required according to the 2006 CPFD joint distress was limited and joint behavior compared LRFD RSGS. Principal compressive stresses did not exceed very favorably with the CIP specimen. Analysis of the joint indi- 0.07fc, less than a third of the 2006 LRFD RSGS limit of 0.25f c. cated that the principal tensile stress was limited to 4.4 fc , less These values correspond well with the intentions of the design and the observed joint performance. Accounting for the differ- than half of the 2006 LRFD RSGS (2) limit of 12 fc , but 37% ent concrete strengths, the CPFD stresses were 11% to 19% larger than 3.5 fc , the level at which more extensive (addi- smaller than those for CIP. The joint shear stress-strain 100% Fixed End Rotation Column Flexure 80% Bent Cap Flexibility Joint Shear 59.9% 64.5% 69.6% 71.2% 71.7% 60% 79.6% Displacement Component 40% 13.4% 11.1% 8.9% 13.4% 20% 15.0% Push 20.2% 18.2% 8.4% 15.7% 11.1% 9.5% 8.6% 6.5% 6.2% 5.8% 4.4% 3.8% 3.5% 0% -6.1% -5.3% -4.3% -3.3% -3.1% -7.3% -9.2% -8.0% -6.6% -13.2% -11.2% Pull -15.1% -20% -17.8% -24.4% -30.0% -31.8% -33.9% -40% -33.4% -60% -72.5% -64.3% -56.5% -48.9% -49.6% -44.3% -80% Note: Curvature gages unreliable after 6. -100% 0.63 0.85 1.10 1.62 2.17 3.20 (1) (1.5) (2) (3) (4) (6) Drift Ratio (%) (Lateral Displacement Ductility) Figure 2.58. Displacement decomposition component percentages--CPFD.

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43 2500 Joint 2000 Bent Cap Yield Strain Bent Cap South North 1500 1000 500 Microstrain 0 Push -500 Pull -1000 13 kips 20 kips -1500 30 kips -2000 48 kips Yield Strain -2500 -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 Location (in) Figure 2.59. Strain profile--stirrups in bent cap (midheight) and joint (bottom), force control--CPFD. response compared closely to the CIP, with limited joint soft- and joint shear cracking and deformation (see Figure 2.60, ening evident at increasing drift ratios (see Figure 2.46). Figure 2.61, Figure 2.47, and Figure 2.46). However, the The maximum change in the CPFD panel area was approx- system achieved an unexpectedly large drift ratio of imately 20% less than that for the CIP specimen, correspon- 5.1% (nominal displacement ductility of 8), and load- ding with fewer diagonal cracks in the CPFD joint region displacement response indicated stable hysteretic behavior and a significantly smaller maximum diagonal crack width without appreciable strength degradation. Failure was due (0.009 in) compared to the CIP joint (0.025 in). In addition, to buckling and fracture of two column bars rather than only at a 3.2% drift (6, pull) did diagonal cracks pass through joint failure. These characteristics were similar to the full the central portion of the CPFD joint itself. The CIP joint ductility specimens. 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- Figure 2.60. Specimen response at a 2.5% drift ratio tion of plastic hinging of the column adjacent to the bent cap (4)--CPLD.

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

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49 Figure 2.69. Lateral displacement decomposition--HYB1. (a) HYB1 (b) HYB2 (c) HYB3 Figure 2.70. Joint region cracking post test--hybrid specimens. 12 0.6 0.50% 0.75% Height Above Bent Cap, inches 10 0.5 Height / Column Diameter 1.00% 1.50% 8 2.00% 0.4 3.00% 6 4.00% 0.3 6.00% 4 0.2 2 0.1 Pull (North Gages) Push (South Gages) Bedding Layer 0 0 -20000 -15000 -10000 -5000 0 5000 10000 15000 20000 Strain, Figure 2.71. Compression strain distribution--HYB1.

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50 100 as shown in Figure 2.73 and Figure 2.74. Up to the 2.0% drift 1.0% Residual Drift Ratio level, the overall response of the system was as anticipated. Residual / Maximum Drift, % 80 1.5% Residual Drift Ratio However, following the drift cycles to 2.0%, noticeable degra- dation of the grout bedding layer was observed. Deterioration 2.0% RDR continued with increasing lateral drifts. The degradation in the 60 bedding layer resulted in a continual loss of lateral strength due 40 to a reduction in the effective column dimension. No damage was observed in the column outside of the bedding layer. HYB1 20 HYB2 Fracture of the reinforcement was noted on 6.0% drift ratio HYB3 cycles with similar observed buckling leading to fracture. The Cast-in-place 0 bent cap responded as anticipated and similarly to the con- 1 2 3 4 5 6 ventional hybrid specimen even with the increase in lateral Drift Ratio, % demand recorded. The overall performance of the bent cap Figure 2.72. Residual drift ratio versus applied drift joint indicated only minor flexural cracking, and small crack ratio--three hybrid systems. widths indicated a reliable joint design methodology was used. Column Lateral Force versus Lateral Displacement. The three hybrid specimens as well as the CIP control specimen. complete force-displacement curve obtained for this specimen Only the first cycle residual drift ratios are shown; however, the is shown in Figure 2.75. The lateral force presented is the actual second cycle exhibited only slightly greater residual drifts. In lateral force considering the effects of system deformation dur- general, for the conventional hybrid specimen the residual drift ing testing. Hysteretic response was stable up to a 6.0% drift ratio increases with the applied lateral drift. However, the ratio in terms of the stability of the hysteresis loops under recorded residual drift is significantly less in comparison to the repeated cycles. However, loss of lateral strength was observed CIP specimen, indicating an overall improvement in the post- in both the positive and negative directions following loading earthquake performance of the system. 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- Concrete Filled Pipe Hybrid Specimen umn diameter. According to the commonly accepted defini- Similar to the conventional hybrid specimen, the primary tion of failure as when the system lateral strength is 80% of the lateral response of the concrete filled pipe specimen (HYB2) is maximum, the concrete filled pipe specimen failed at a 5.0% dominated by the localized end rotations at the bedding layer, drift ratio. (a) (b) Figure 2.73. Specimen response at 2% drift (a) column base and (b) joint--HYB2.

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51 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 Figure 2.74. Specimen response at 6% drift--HYB2. amount of deformation was not captured during the lower level loading cycles. The nominal capacity calculated using the simplified analy- sis technique and the complete force-displacement prediction Joint Response. Observed bent cap joint damage follow- are also provided. Figure 2.75 indicates that the nominal capac- ing testing of the concrete filled pipe hybrid specimen is shown ity predicted using the simplified procedure provides a reason- in Figure 2.70b. Figure 2.70b indicates that only minor damage able and slightly conservative estimate of the nominal lateral occurred within the joint during the entirety of the testing, sim- capacity of the specimen. Additionally, the complete force- ilar to what was observed in the conventional specimen. The displacement prediction matches very well with the recorded level of observed damage is also of a similar magnitude even response up to the 2.0% drift level. Following the cycles to 2.0% though the lateral demands, and therefore joint demands, were drift, the degradation in the bedding layer was not captured by greater for this specimen. Diagonal cracking patterns indicate the prediction; thus, the expected lateral resistance continued that joint shear cracking occurred, but the joint reinforcement to grow. design was adequate to resist extensive crack growth and sub- The force-displacement envelopes for all three hybrid spec- sequent joint damage. imens along with the CIP specimen, are shown in Figure 2.68. Comparison of the conventional (HYB1) and concrete filled Residual Drift. Review of Figure 2.72 shows the ratio of pipe (HYB2) envelopes shows the stability of the lateral resis- residual drift to maximum drift during that cycle for this spec- tance for the conventional specimen whereas a continual reduc- imen. The observed residual drift for this specimen is greater tion in strength is observed for the concrete filled pipe specimen. than that recorded for the conventional hybrid specimen, Drift Ratio, % -8 -6 -4 -2 0 2 4 6 8 100 75 Nominal Capacity 50 Lateral Force, kips 25 0 -25 -50 -75 HYB2 Prediction -100 -5 -4 -3 -2 -1 0 1 2 3 4 5 Displacement, inches Figure 2.75. Lateral force versus lateral displacement--HYB2.

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52 Figure 2.76. Lateral displacement decomposition--HYB2. resulting from the increased damage in the bedding layer dur- dominated by the localized end rotations at the bedding layer, ing this specimen's testing. Similar to the conventional hybrid as shown in Figure 2.77 and Figure 2.78. Up to the 2.0% drift specimen, only slightly greater residual drifts were recorded level, the overall response of the system was as anticipated. during the second cycle to a given drift. Even though the resid- However, similar to the concrete filled pipe hybrid specimen, ual drifts were greater than those of the conventional hybrid following the drift cycles to 2.0%, noticeable degradation of the specimen, the recorded residual drift was significantly less grout bedding layer was observed, with deterioration continu- than the residual drift of the CIP specimen, indicating an over- ing with increasing lateral drifts. The degradation in the bed- all improvement in the post-earthquake performance of the ding layer resulted in a continual loss of lateral strength due to system. 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 Dual Steel Shell Hybrid Specimen similar observed buckling leading to fracture. The bent cap Similar to concrete filled pipe hybrid specimen, the primary responded as anticipated even with the increase in lateral lateral response of the dual steel shell hybrid specimen was demand recorded, similar to the conventional hybrid speci- (a) (b) Figure 2.77. Specimen response at 2% drift (a) column base and (b) joint--HYB3.

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

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

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55 Drift Ratio, % -8 -6 -4 -2 0 2 4 6 8 0.2 0.1 Bedding layer growth, inches 0 -0.1 -0.2 -0.3 -0.4 -5 -4 -3 -2 -1 0 1 2 3 4 5 Displacement, inches Figure 2.82. Bedding layer centerline axial deformation-- HYB3. conventional hybrid specimen is attributable to the increased ous reinforcement extending from the girder into the reaction damage in the bedding layer during the dual steel shell hybrid block resulted in the observed concentrated opening at the specimen's testing. Similar to the conventional hybrid speci- joint during negative flexure; however, the presence of the men, only slightly greater residual drifts were recorded during deck flexural reinforcement served to reduce the concentra- the second cycle to a given drift. Even though the residual drifts tion of cracking within the deck. are greater than those of the conventional hybrid specimen, in During positive loading cycles, flexural cracking was con- comparison to the CIP specimen, the recorded residual drift is centrated at the girder to reaction block joint. Essentially elas- significantly less, indicating an overall improvement in the tic response was observed within the section up to the point of post-earthquake performance of the system. 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 2.3.3 Integral Connection continued. During reversed cycling, the fiber-reinforced clo- The integral experimental specimen (INT) was subjected to sure joint performed well, with no observed reduction in 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 Figure 2.83. Girder end block region at 0.29% joint into two cracks a couple of inches apart. The lack of continu- rotation.

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

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

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