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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 (µ) 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 (µ3), 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- (µ4)--CIP.
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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-
(µ10)--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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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) (µ1.5) (µ2) (µ3) (µ4) (µ6) (µ8) (µ10)
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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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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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 (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 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. (µ4)--GD.
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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
(µ8)--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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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 µ4.
-100%
0.56 1.12 1.88 2.55
(FC45) (FC55) (µ3) (µ4)
Drift Ratio (%)
(Stage)
Figure 2.53. Displacement decomposition component percentages--GD.
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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.
OCR for page 41
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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,
(µ4)--CPFD.
integral behavior between the pocket concrete, pipe, and sur-
rounding concrete was evident.
0.9% (µ1.5), with spalling much more evident at a drift of Column Lateral Force versus Lateral Displacement. The
3.2% (µ6). 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% (µ6, 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
(µ8)--CPFD. and CIP specimens compared very favorably.
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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 µ6.
-100%
0.63 0.85 1.10 1.62 2.17 3.20
(µ1) (µ1.5) (µ2) (µ3) (µ4) (µ6)
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 (µ6, 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 (µ4)--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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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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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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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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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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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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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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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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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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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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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.
