Performance Evaluation of Longitudinal Ultrahigh-Performance Concrete Closure Pour Connection for Use in Modular Bridge Construction Pairwise Comparison of Capacity and Ductility at Failure Limit State Yaohua Deng, Brent M. Phares, Andrew J. Putz, Curtis Carter, Michael Nop, and Dean Bierwagen Because of increased traffic congestion, work zone safety concerns, and economic constraints, rapid renewal techniques are desirable in replacing the nation’s aging infrastructure (2). The development and study of accelerated bridge construction (ABC) techniques to address such needs under SHRP 2 is evidence of this. ABC techniques taking advantage of prefabricated bridge elements and high-performance materials are being used more commonly for bridge replacement projects. They result in minimal road closure time and traffic interruption and in reconstruction of long-lasting highway bridges. These techniques have been used in several demonstration bridge projects and are increasingly being adopted. For example, the goal of SHRP 2 Project R04 was to develop standards and codified language for ABC and to provide for the construction of demonstration bridges, such as the Keg Creek Bridge in Pottawattamie County, Iowa, which consisted of several prefabricated steel beam and concrete deck components connected with both transverse and longitudinal closure pours. To address design concerns in the negative moment region, laboratory tests were conducted to evaluate the ultrahigh-performance concrete (UHPC) transverse full-depth deck joint over the pier of the demonstration bridge (3). For this study, a project consisting of the replacement of a bridge located on IA-92 over Little Silver Creek in Pottawattamie County, Iowa, is presented (see Figure 1). The concepts used in the Little Silver Creek Bridge are, for all intents and purposes, a further iteration of those used in the Keg Creek Bridge. For both bridges, prefabricated bridge elements were placed adjacently on the substructure and connected by using longitudinal closure pour connections, as shown in Figure 1. The use of prefabricated components improves construction quality and reduces traffic disruption time on the bridge site. The longitudinal connections are an important deck-level component for transferring shear and moment between the prefabricated components. The longitudinal connections are heavily stressed by traffic loads and environmental effects, which gives rise to concerns for their durability. A literature search indicated that three types of longitudinal connections had been introduced for modular bridge elements: welded steel connections, distributed reinforcement connections, and UHPC connections (4–6). The welded steel connection was detailed by using welded steel
Accelerated bridge construction techniques taking advantage of pre fabricated bridge elements and high-performance materials are being used more frequently for bridge replacement projects. They result in minimal road closure times and traffic interruption and in the recon struction of long-lasting highway bridges. Longitudinal closure pour connections are an important deck-level component for modular bridge elements that are heavily stressed by traffic loadings and environmen tal effects and whose durability is a concern. To address cracking and leakage issues in such connections, the strength and failure modes of the longitudinal ultrahigh-performance concrete (UHPC) closure pour connection between adjacent prefabricated deck units were evaluated. First, specimens with and without a longitudinal UHPC closure pour connection were fabricated, instrumented, and tested. Finite element (FE) models were established to improve understanding of the behavior of the specimens under the loading condition. In addition, strut-and-tie models (STMs) were developed on the basis of FE model predictions to estimate the strength of the specimens. The jointed specimens were found not to have any cracks or leakage at the early stage but had lower cracking loads than did the jointless specimens. The strength and ductility of the jointed specimens were comparable with those of the jointless speci mens. On the basis of the FE models and STMs, the ultimate strength of the specimens was accurately predicted.
Numerous years of underfunding have resulted in the need for major replacement initiatives to improve the state of the nation’s infrastructure. According to ASCE’s 2013 Report Card for America’s Infrastructure—Bridges (1), approximately 11% of the nation’s 607,380 bridges are “structurally deficient,” and many bridges are approaching the completion of their design life and are in need of replacement. Y. Deng, B. M. Phares, and A. J. Putz, Bridge Engineering Center, Institute for Transportation, Iowa State University, 2711 South Loop Drive, Suite 4700, Ames, IA 50010-8664. C. Carter, M. Nop, and D. Bierwagen, Office of Bridges and Structures, Iowa Department of Transportation, 800 Lincoln Way, Ames, IA 50010. Corresponding author: Y. Deng,
[email protected]. Transportation Research Record: Journal of the Transportation Research Board, No. 2592, Transportation Research Board, Washington, D.C., 2016, pp. 65–75. DOI: 10.3141/2592-08 65
66
Transportation Research Record 2592
Specimen design
Longitudinal connection
Prefabricated deck unit (a)
(b) 10 in.
1 ft–3 in.
8 in. W40 × 149
W40 × 149
W40 × 149
MC18 × 42.7
4 ft–6 in.
W40 × 149 MC18 × 42.7
3 ft–4 in.
4 ft–6 in.
(c) FIGURE 1 Little Silver Creek Bridge on IA-92: (a) cross section, (b) plan view, and (c) details of segment for specimen design.
connectors and was popularly used for prestressed concrete bridges (3). However, because of a low flexural load resistance, this connection results in cracks over the longitudinal joint. Furthermore, these cracks commonly cause joint leakage, which induces deterioration of other below-deck structural components and limits the wider application of such connections for bridges (4). To address the cracking issue, three details based on a distributed reinforcement concept using U-shaped bars, headed bars, and spiral wire were proposed by Li et al. (4), and one similar detail using hook bars was introduced by Gull et al. (5). The connections using U-shaped bars and spiral wire were not evaluated extensively because of disadvantages in terms of constructability and economics. Experimental results indicated that the connections using headed bars and hook bars have sufficient strength and ductility (4, 5). However, their strength and serviceability need further evaluation in real bridges using these connections. FHWA has recently investigated the use of UHPC as a closure material for purposes of improving the performance of the longitudinal connection, reducing connection width, and simplifying the reinforcement configurations in the connection (6). As a result, a connection detail was developed that has the merits of UHPC (i.e., high durability, high strength, and superior bond action) (6). Fatigue load test results and the performance of two bridges constructed in New York State indicated the viability of the longitudinal UHPC connection for modular bridge elements (6).
To facilitate use of the UHPC connection for modular bridge elements, the flexural behavior and strength and failure modes of longitudinal UHPC connections between adjacent prefabricated deck units used in a second generation detail based on the SHRP 2 research results were evaluated. To do so, a pairwise comparison of the capacity and ductility at the failure limit state between a UHPC connection and a nonjointed detail was made.
Experimental Program An experimental program consisting of ponding and strength tests was designed and implemented. For comparison purposes, three specimens with and three specimens without the UHPC longitudinal connection (named jointed and jointless specimens, respectively) were fabricated, instrumented, and tested.
Specimen Design and Fabrication The test specimens were designed on the basis of the details of the Little Silver Creek Bridge, which has six prefabricated deck modules and five longitudinal UHPC closure pour connections, as shown in Figure 1. The details of the three jointed specimens were established by exactly replicating the details for two modules and one connection
Deng, Phares, Putz, Carter, Nop, and Bierwagen
67
For comparison purposes, three specimens without a connection consisting of a single deck panel were also evaluated. These jointless specimens are, as much as anything, a normal baseline with which the jointed test results can be compared. Each specimen has a width of 14 ft 10 in., a length of 3 ft, and a depth of 8 in. The configuration of these specimens was established to represent a standard bridge deck with the same amount of reinforcement as the jointed specimens (see Figure 4). All of the reinforcing steel bars have the same covers and spacing. To represent the rebar layout in a standard bridge deck, the No. 6 bars run the continuous width of the specimen with no overlap. The No. 6 bars are spaced at 1 ft with 6 in. to the edge of the specimen, and the No. 5 bars are evenly spaced at 11 in. across the whole deck panel with 3.75 in. of clear distance to the edges, as shown in Figure 4b. Four No. 5 bars at a spacing of 5.5 in. are designed at the connection location, as shown in Figure 4a. The specimens were fabricated in the Structural Engineering Laboratory at the Iowa State University. The steel bars were installed and placed into the formwork, as shown in Figure 3, a and b. The form liner or retarder was applied to the formwork at the connection interface, as shown in Figure 3b. The placement and surface finish of the concrete for the deck panels are as shown in Figure 3, c and d, respectively. After 28 days, the two deck panels of each jointed specimen were properly oriented, the connection formed up, and the UHPC placed into the formwork, as shown in Figure 3e. The finished UHPC connection is shown in Figure 3f. The jointless specimens were tested after 28-day curing, and the 28-day compressive strength of the normal-strength concrete was measured as 5.2 kips per square
in the yet-to-be-constructed Little Silver Creek Bridge shown in Figure 1c. Each of the jointed specimens has two deck panels with a width of 7 ft, a length of 3 ft, a depth of 8 in., and a 10-in.-wide UHPC longitudinal connection between the two deck panels. The jointless specimens have the same overall geometry as the jointed specimens, except for the joint. The cross sections and reinforcement details of the jointed specimens are shown in Figure 2. As shown in Figure 2, a and b, a clear cover of 2.625 in. and 1 in. was used for the top and bottom steel bars, respectively. As shown in Figure 2, a through d, the longitudinal bars are No. 5 bars spaced at 11 in., and the transverse bars are No. 6 bars spaced at 1 ft. The top and bottom longitudinal bars are staggered at a distance of 5.5 in. The transverse bars of two deck panels are staggered panel to panel at a distance of 6 in. and fit into the deck joint connection shown in Figure 3, c and d. All No. 6 bars project 9 in. into the 10-in. UHPC connection with an overlap length of 8 in., as shown in Figure 2e. Two No. 7 bars are also placed in the center of the connection above and below the No. 6 bars, as shown in Figure 2e. Three surface preparation techniques were used to texturize the joint surface of the three jointed specimens [two types of form liners (known as rubber sandblast medium and plastic sandblast) and a form retarder]. The three surface treatments met the designer’s criteria for achieving a specified concrete surface profile level. As will be seen, since the loads considered in this work were much higher than the strength limit loads considered for design, the surface treatments selected did not provide an interface bond higher than the concrete tensile capacity.
5.50 in.
11.00 in.
5.50 in.
11.00 in.
3.50 in. 9.00 in.
3.50 in. 9.00 in.
5.50 in.
5.50 in.
11.00 in.
11.00 in.
2.625 in.
2.625 in.
1.50 in.
1.50 in. 1.00 in.
1.50 in. 11.00 in. No. 6 Bars
No. 5 Bars
11.00 in.
1.50 in. 11.00 in. No. 6 Bars
No. 5 Bars
(a) 3.50 in.
5.50 in.
(b) 3.50 in.
11.00 in.
11.00 in.
1.00 in.
11.00 in.
5.50 in.
3.50 in.
3.00 in.
5.50 in.
11.00 in.
3.50 in.
9.00 in.
1.50 in. 1 ft
9.00 in.
9.00 in.
1.50 in. 1 ft
1 ft 6.00 in.
6.00 in.
1 ft 9.00 in. T&B
T
B
T
B
T
B
T
B
T
B
T
B
T
3.00 in.
1.50 in.
1.50 in. T&B
T&B
No. 5 Bars = All Bars Running Vertical No. 6 Bars = All Bars Running Horizontal
T
B
T
B
T
B
T
B
T
B
T
B
T
T&B
No. 5 Bars = All Bars Running Vertical No. 6 Bars = All Bars Running Horizontal
(c)
(d)
Longitudinal Joint Detail No. 7 Bars
6 in. Bar Stagger Between Modules
10 in. 5 in.
5 in.
UHPC
8 in. No. 6 Bars
(e) FIGURE 2 Details of jointed specimens: (a) cross section, left deck panel; (b) cross section, right deck panel; (c) rebar layout, left deck panel; (d) rebar layout, right deck panel; and (e) UHPC connection detail (T 5 top; B 5 bottom).
Form liner or retarder
(a)
(b)
(c)
(d)
(e)
(f)
FIGURE 3 Fabrication sequence of specimens: (a) formwork and rebar, jointless specimen; (b) formwork and rebar, jointed specimen; (c) placement of normal-strength concrete; (d) finish of normal-strength concrete; (e) placement of UHPC; and (f ) finish of UHPC.
Deng, Phares, Putz, Carter, Nop, and Bierwagen
2.25 in.
69 3.75 in.
5.50 in.
5.50 in.
11.00 in.
2.625 in. 1.00 in. 1.50 in.
5.50 in.
No. 6 Bars
11.00 in.
1.50 in.
No. 5 Bars
Connection location
(a) No. 5 Bars
1.50 in.
No. 6 Bars 6.00 in.
1.50 in.
1 ft-0 in.
1 ft-0 in. 6.00 in. T&B T
B
T
B
T
B
T
B
T
B
T
No. 6 Bars
B
T
B
T&B T&B
B
T
B
T
B
T
B
T
B
T
B
T
B
T
T&B
14 ft–10 in.
(b) FIGURE 4 Details of jointless specimens: (a) cross section and (b) rebar layout.
inch (ksi). The jointed specimens were tested after a 6-day cure of the UHPC, which had a measured 6-day compressive strength of 18.2 ksi. Curing and Ponding Tests A ponding test for a jointed specimen (J2) was conducted to check whether the UHPC connection developed cracks during curing. During the placement and curing of the UHPC pour, vertical restraining forces (as shown in Figure 5a) were applied at the specimen ends to simulate the transverse restraint provided by the girders. The test setups were determined on the basis of the bridge details shown in Figure 1c. During the curing process, the interface between the UHPC and the normal concrete was visually observed to record any crack formation. At the 5th day after UHPC placement, a pond was formed on the top of the connection for 6 h to check whether any leakage occurred at the connection and interface, as shown in Figure 5a. This experimental ponding regimen followed the ponding regimen specified for construction of the actual bridge. Strength Tests Strength tests were conducted to check whether the jointed specimens had the same strength as did the jointless specimens. All specimens were tested under the loading and boundary conditions shown in Figure 5b. Each specimen was supported in a manner that simulates the girders in each adjacent precast unit. The center span is 3 ft 4 in. long, and the two outside spans are 4 ft 4 in. long. These dimensions replicate the girder spacing from the bridge plans. The supports run the entire length of the specimen. For each jointed specimen, four strain gauges were installed on the bottom layer of reinforcing bars and embedded 2 in. away from the outermost line of the connection interface before concrete placement (see Figure 5c). A deflection transducer was installed on the bottom of
the specimen at the midcenter span (see Figure 5c). For each jointless specimen, the instrumentation layout was the same as for the jointed specimens (see Figure 5d). Two-line loads were applied to the jointed specimen 3.5 in. away from the outermost interface surface. The same loading approach was used for the jointless specimens. The loads were continuously applied on the specimens by two hydraulic actuators each fitted with load cells to record the applied loading. The loading continued until it was decided that each specimen had failed. Each specimen was visually observed multiple times throughout the process. Visual crack mapping techniques were used to monitor and document crack formation in the deck panels, joint material, and interfaces during loading. Finite Element Analysis Three-dimensional nonlinear finite element (FE) models (see Figure 6) were established to investigate the performance of the specimens further. The concrete deck and longitudinal connection were both modeled by using an eight-node solid element with three translational degrees of freedom at each node plus cracking and crushing capabilities. The Poisson’s ratio of the concrete was set to 0.2. The concrete material properties were also assigned with multilinear isotropic hardening in combination with the von Mises yield criterion. The stress–strain relationships of the normal-strength concrete and UHPC can be expressed by Equations 1 and 2, respectively, according to Hognestad (7) and Schmidt and Fehling (8), respectively. 2
ε ε fc = f c′ 2 − ε on ε on
(1)
n
ε fc = f c′ 1 − 1 − ε ou
(2)
70
Transportation Research Record 2592
Pond
(a) Loading
1 ft–5.00 in.
4 ft–4.00 in.
3 ft–4.00 in.
4 ft–4.00 in.
7 ft
10.00 in.
7 ft
1 ft–5.00 in.
(b)
3 3
Strain gauge
1 1
Deflection transducer
4
4
Strain gauge Deflection transducer
2 2
Joint line location for jointed specimens
Joint line location for jointed specimens (c)
(d)
FIGURE 5 Test setups and instrumentation: (a) test setup of curing and ponding tests, (b) test setup of strength tests, (c) instrumentation of jointed specimens, and (d) instrumentation of jointless specimens.
where fc = stress on concrete, f c′ = concrete compressive strength, ε = strain on concrete, n = parameter (1.4 for the UHPC compressive strength of 18.2 ksi), εon = strain at peak stress for normal-strength concrete, and εou = strain at peak stress for UHPC. Equations 3 and 4 provide expressions for εon and εou, respectively. ε on = 0.00048 ( f c′ )
14
( ksi )
(3)
To Line Restraint
ε ou =
(4)
where Ecu is the elastic modulus of UHPC, which can be expressed by Equation 5 (9): Ecu = 1,550 f c′
( ksi )
(5)
The smeared fixed crack model and Rankine maximum stress criterion were used to determine the initiation and development of concrete cracking. The tensile strength of the normal-strength
Line Loads Connection
f c′ Ecu
Deck Panel
Line Support FIGURE 6 FE model of jointed (or jointless) specimens.
Deng, Phares, Putz, Carter, Nop, and Bierwagen
71
concrete and UHPC can be derived by Equations 6 and 7 on the basis of AASHTO (10) and Russell and Graybeal (9), respectively. f t′= 0.24 f c′
( ksi )
f t′= 0.21 f c′
( ksi )
shapes and tensile stress relaxation after cracking were applied to solid elements (11–14).
(6) Results (7)
The steel bars were modeled by using a uniaxial tensioncompression element with three translational degrees of freedom at each node. A perfect elastic–plastic uniaxial material model was used for the steel. The yield strength, elastic modulus, Poisson’s ratio, and tangent modulus of the steel were set to 60 ksi, 29,000 ksi, 0.3, and 0, respectively. The steel bars were perfectly connected to the concrete through the sharing of common nodes. The concrete deck was also perfectly connected to the UHPC connection through the sharing of common nodes. No relative displacement between the deck panels and connection was observed, and only cracking was found in the interface or normal-strength concrete during testing. Accordingly, full bond was assumed at the interface between the concrete deck and connection, and the maximum bond strength should be equal to the tensile strength of the normal-strength concrete. Because of the asymmetry of the reinforcing detail in the jointed specimens, full models were established for all specimens shown in Figure 6. Similar meshes were used for the two types of specimens, but the material properties of the normal-strength concrete and UHPC were appropriately assigned. Line loads were applied on the nodes of the FE model at the loading location, and boundary conditions were defined to simulate the four support locations shown in Figure 6. Convergence criteria and tolerances were set for the displacement and force. The following strategies were used to facilitate convergent computations: (a) concrete compressive stress was constrained to a constant value after reaching its peak value; (b) shear transfer coefficients of 0.3 and 0.6 were used for open and closed cracks, respectively; (c) capability of concrete crushing was deactivated in the analysis, but model failure was determined when the maximum concrete compressive stress reached the concrete compressive strength; and (d) suppression of extra displacement
Curing and Ponding Tests During curing of the UHPC pour, no cracks were visually identified at the interface (top and bottom surfaces and two side surfaces). During the ponding test, no leakage was found. It was concluded that a good bond was achieved at the interface between the concrete and the UHPC and that the deformation due to early-age drying shrinkage and temperature change did not cause any cracks in the connection and interface.
Strength Tests Before failure of each specimen, cracks occurred above the two center supports, between the two loading lines, and in the first and third spans, as shown in Figure 7, a, c, and d. For the jointed specimens, cracks also formed at the connection interface because of the separation of the normal concrete and UHPC, as shown in Figure 7b. A flexural-shear failure occurred at the center span of the jointless and jointed specimens, as shown in Figure 7, c and d. The concrete crushed near a loading line, and large diagonal cracks formed and were extended from a center support to the loading line as shown in Figure 7, c and d. No concrete crushing or cracks were found in the UHPC. Load–strain relationships were developed on the basis of the measured strains in the four gauges for each of these specimens. Typical load–strain relationships for Specimen C1 (jointless) and Specimen J2 (jointed) are shown in Figure 8, a and b, respectively. For each load–strain relationship, a plateau can be observed at an early loading stage because of the concrete cracking shown in Figure 8, a and b. For each specimen, the load at concrete cracking was determined on the basis of the smallest plateau of the four load–strain
Cracks
Cracks
(a)
(b)
FIGURE 7 Cracking patterns and failure modes of specimens: (a) cracking pattern on Specimen C1 and (b) cracking at connection interface of Specimen J2. (continued on next page)
Flexural-shear failure
Flexural-shear failure Concrete crushing
Concrete crushing
Cracks
Cracks
(c)
(d)
FIGURE 7 (continued) Cracking patterns and failure modes of specimens: (c) failure modes of Specimen C1 and (d) failure modes of Specimen J2. 270 240
Load (kips)
210 180 150 120 90 60
Test results, Strain Gauge 1 Test results, Strain Gauge 2 Test results, Strain Gauge 3 Test results, Strain Gauge 4 FE results, Strain Gauge 1, 2, 3, or 4
30 0
Plateau 0
1,000
2,000
3,000
4,000
5,000
6,000
7,000
8,000
9,000 10,000
Strain (microstrain)
(a) 240 210 180
Load (kips)
150 120 90
Test results, Strain Gauge 1 Test results, Strain Gauge 2
60
Test results, Strain Gauge 3 Test results, Strain Gauge 4
30 0
FE results, Strain Gauge 1 or 4 FE results, Strain Gauge 2 or 3
Plateau 0
1,000
2,000
3,000
4,000
5,000
6,000
7,000
8,000
Strain (microstrain)
(b) FIGURE 8 Comparison of relationships obtained from test results and FE models: (a) load–strain relationships of jointless Specimen C1 and (b) load–strain relationships of jointed Specimen J2. (continued)
9,000
Deng, Phares, Putz, Carter, Nop, and Bierwagen
73
300
250
Load (kips)
200
150
Test results, Specimen C1
100
Test results, Specimen C2 Test results, Specimen C3 Test results, Specimen J2
50
Test results, Specimen J4 FE results, Jointless specimens FE results, Jointed specimens
0
0.0
0.1
0.2
0.3
0.4
0.5
0.6
Deflection (in.)
(c) FIGURE 8 (continued) Comparison of relationships obtained from test results and FE models: (c) load–deflection relationships.
The stresses and loads at failure of the specimens as predicted by the FE models were determined on the basis of concrete crushing on the top of specimens for the jointless and jointed specimens as shown in Figure 9, a and b, respectively. That is, the concrete compressive stress exceeds the normal concrete strength of 5,200 pounds per square inch. The load–strain relationships obtained from FE models were compared with test results and are shown in Figure 8, a and b. The crack loads and yield loads were also determined on the basis of the same approaches used for processing the experimental test results. The load–deflection relationships obtained from FE models were compared with test results, as shown in Figure 8c. Figure 8 indicates that the relationships for the jointed specimen follow patterns similar to those for the jointless specimens, and the predictions from the FE models indicate higher stiffness and less ductility as compared with the test results. Table 1 indicates that cracking and yield loads were slightly overpredicted by the FE models compared with the test results and that the failure deflections were slightly underestimated. The discrepancies arise because (a) the smeared crack model used in the FE models did not provide accurate prediction of the deformation caused by the cracks and (b) the bond stress–slip relationship between the concrete and steel bars was not taken into account. However, the failure loads predicted
relationships; a summary appears in Table 1. Likewise, the load at steel yield was determined on the basis of the smallest load when the steel bars reached the yield strain (i.e., 2,069 microstrain for 60-ksi steel), as summarized in Table 1. The maximum load at failure and the deflection at the maximum load are also summarized in Table 1 for each specimen. As Table 1 indicates, the crack loads for the jointed specimens are slightly smaller than those for the jointless specimens. This is because the bond strength at the longitudinal connection interface is less than the normal-strength concrete tensile strength. In addition, the form liner plastic sandblast produced the least bond between the normal-strength concrete and the UHPC because of the lack of roughness at the interface. Table 1 also indicates that the yield loads and the failure loads on the jointed specimens are slightly less than those on the jointless specimens. However, the deflections at failure of the jointed specimens are close to those of the jointless specimens. Furthermore, the load–deflection relationships of the jointed specimens are comparable with those of the jointless specimens. The conclusion is that the longitudinal closure pour connection has good performance in connecting the two deck panels and that the strength and ductility of the jointed specimens are comparable with those of the jointless specimens. TABLE 1 Summary and Comparisons of Test and FE Model Results Load at Concrete Cracking (kips)
Load at Steel Yield (kips)
Load at Specimen Failure (kips)
Deflection at Maximum Load (in.)
Type of Specimen
Slab No.
Joint Surface Preparation Technique
Test
FE Analysis
Test
FE Analysis
Test
FE Analysis
STM
Test
FE Analysis
Jointless
C1 C2 C3 J1 J4 J2
na na na Rubber sandblast medium Plastic sandblast Retarder
52 56 39 34 20 31
57
115 123 126 110 112 114
164
250 235 230 195 225 210
232
225
0.27
210
205
0.37 0.42 0.39 — 0.39 0.38
Jointed
Note: na = not applicable; — = bad data.
49
180
0.16
74
Transportation Research Record 2592
Y X Z
MX Lateral restraining action –5,175
–3,777
–4,475
Arching action –2,379
–3,078
–1,680
–981
–283
416
1,115
(a)
MN
Z X MX Lateral restraining action
Arching action –5,196
–4,537
–3,878
–2,559
–3,218
–1,899
–1,240
–581
79
738
(b)
Z
Top layer
MN
X
MX Bottom layer
Steel yield –2,873
4,114
11,101
18,088
25,075
39,049
32,062
46,036
53,023
60,010
(c) Y
MN
Top layer X MX Bottom layer
Steel yield
–3,363
3,678
10,719
17,760
24,801
31,842
38,883
45,923
52,964
(d) Load (P)
Horizontal arching strut (Ch )
P/2
P/2
Diagonal arching strut (Cl ) θ Restraining strut (Cr )
Diagonal arching strut (Cl ) Tie (Th )
Restraining strut (Cr )
(e) FIGURE 9 Stresses in concrete and rebar at failure predicted by FE models and forces in developed STM: (a) normal-strength concrete of jointless specimen [pounds per square inch (psi)], (b) normal-strength concrete of jointed specimen (psi), (c) steel bars of jointless specimen (microstrain), (d) steel bars of jointed specimen (microstrain), and (e) STM (MX 5 maximum; MN 5 minimum).
60,005
Deng, Phares, Putz, Carter, Nop, and Bierwagen
from the FE models are in good agreement with the test results, as shown in Table 1. Because the shear span-to-depth ratio (1.44) is less than 2, the center span falls into a D-region, and the load-carrying capacity should be evaluated with a strut-and-tie model (STM). On the basis of the concrete compressive stress flow shown in Figure 9, a and b, arching action was found between the two load lines and between the loads and supports; lateral restraining action was also found at the bottom regions of the first and third spans. Figure 9, c and d indicates that the steel bars were all in tension between the two center supports. For the jointless specimens, the top and bottom layers of steel bars yielded under the loading location and the top layer of steel bars also yielded above the center supports; for the jointed specimens, the bottom layer of steel bars yielded under the loading location and the top layer of steel bars only yielded above the center supports and has an average stress of 32 ksi under the loading location. An STM was developed on the basis of the FE model predictions, as shown in Figure 9e. The STM consists of a horizontal arching strut (Ch), two diagonal arching struts (Cl), two restraining struts (Cr), and a tie (Th). The force in the restraining strut is equivalent to the tension force in the top layer of steel bars above the center supports. The tie force is equal to the tension force in the steel bars between the loading lines. The force in the horizontal arching strut is equal to the sum of the forces in the restraining strut and tie. The angle between the diagonal strut and tie is estimated on the basis of the distance from the center supports to the nearest loading line and the center-to-center distance from the horizontal aching strut to the tie. The estimated ultimate load capacities of the jointless and jointed specimens using the STM are 225 and 205 kips, respectively, which agree well with the test results and FE model predictions shown in Table 1. Because of arching action in the deck, a viable design approach based on strut-and-tie methodologies rather than pure flexural bending methodologies may be used. In this approach, variable beam spacings can be strategically used in designs requiring such geometries, which allow the designer to limit tension in the longitudinal joint by inducing large compressive arching forces. Conclusions Ponding and strength tests were conducted to evaluate the behavior of the longitudinal closure pour connection planned to be used in the yet-to-be-constructed Little Silver Creek Bridge. For comparison purposes, specimens with and without a UHPC longitudinal connection were fabricated, instrumented, and tested. An FE model was established to carry out performance evaluation of the specimens with and without a connection and to aid in the interpretation of the test results. In addition, an STM was developed on the basis of the predictions from the FE models to estimate the strength of the specimens. The following conclusions were drawn: • The UHPC connections had no cracks or leakage in the joint or at the interface due to early-age drying shrinkage and temperature changes. • Under strength loading conditions, the jointed specimens had slightly lower cracking loads than did the jointless specimens no matter which joint surface preparation technique was used (i.e., form retarder or form liners). The bond at the longitudinal connection interface was deemed to be less than the normal-strength concrete tensile strength. Of the various surface preparation techniques, the form liner plastic sandblast achieved the worst surface roughness and bond at the interface.
75
• A flexural-shear failure mode (i.e., concrete crushed near the loading location and diagonal cracks extended from a center support) was found in the jointless and jointed specimens. Cracks formed at the connection interface, and no concrete crushing or cracks were found in the UHPC pour. • The strength and ductility of the jointed specimens with a longitudinal closure pour connection are comparable with those of the jointless specimens. • On the basis of the developed FE models, the failure loads were accurately predicted, but the cracking and yield loads were overestimated and the deflections at failure were underestimated. • On the basis of the developed STM, the ultimate load capacity of the specimens was accurately estimated. • Because of arching action in the deck, a viable design approach based on strut-and-tie methodologies may be used to limit tension in the longitudinal joint. References 1. 2013 Report Card for America’s Infrastructure—Bridges. ASCE, Washington, D.C. http://www.infrastructurereportcard.org/bridges/. Accessed July 18, 2014. 2. HNTB Corporation, Genesis Structures, Inc., Structural Engineering Associates, and Iowa State University. SHRP 2 Report S2-R04-RR-1: Innovative Bridge Designs for Rapid Renewal. Transportation Research Board of the National Academies, Washington, D.C., 2014. http:// onlinepubs.trb.org/onlinepubs/shrp2/SHRP2_S2-R04-RR-1.pdf. 3. Hartwell, D. R. Laboratory Testing of Ultra High Performance Concrete Deck Joints for Use in Accelerated Bridge Construction. MS thesis. Iowa State University, Ames, 2011. 4. Li, L., Z. J. Ma, M. E. Griffey, and R. G. Oesterle. Improved Longitudinal Joint Details in Decked Bulb Tees for Accelerated Bridge Construction: Concept Development. Journal of Bridge Engineering, Vol. 15, No. 3, 2010, pp. 327–336. 5. Gull, J. H., A. Yakel, and A. Azizinamini. Experimental Investigation of Longitudinal Closure Pour Detail for Prefabricated Slabs Used in Modular Construction. Presented at 93rd Annual Meeting of the Transportation Research Board, Washington, D.C., 2014. 6. Graybeal, B. A. Behavior of Ultra-High Performance Concrete Connections Between Precast Bridge Deck Elements. Presented at Concrete Bridge Conference: Achieving Safe, Smart, and Sustainable Bridges, Phoenix, Ariz., 2010. http://www.ductal.com/SLib/33-Graybeal-UHPC _Connections-048.pdf. 7. Hognestad, E. Study of Combined Bending and Axial Load in Reinforced Concrete Members. Bulletin No. 399, Engineering Experiment Station, University of Illinois at Urbana–Champaign, 1951. 8. Schmidt, M., and E. Fehling. Ultra-High-Performance Concrete: Research, Development and Application in Europe. American Concrete Institute Special Publication 228, 2005. http://citeseerx.ist.psu.edu/view doc/download?doi=10.1.1.472.5557&rep=rep1&type=pdf. 9. Russell, H. G., and B. A. Graybeal. Ultra-High Performance Concrete: A State-of-the-Art Report for the Bridge Community. FHWA-HRT-13-060. FHWA, McLean, Va., 2013. 10. AASHTO LRFD Bridge Design Specifications, Customary U.S. Units, 5th Edition, with 2010 Interim Revisions. AASHTO, Washington, D.C., 2010. 11. Deng, Y. Efficient Prestressed Concrete–Steel Composite Girder for Medium-Span Bridges. PhD dissertation. University of Nebraska, Lincoln, 2012. 12. Deng, Y., and G. Morcous. Efficient Prestressed Concrete–Steel Composite Girder for Medium-Span Bridges: II—Finite Element Analysis and Experimental Investigation. Journal of Bridge Engineering, Vol. 18, No. 12, 2013, pp. 1358–1372. 13. Deng, Y., T. R. Norton, and C. Y. Tuan. Numerical Analysis of ConcreteFilled Circular Steel Tubes. Structures and Buildings, Vol. 166, No. 1, 2013, pp. 3–14. 14. Deng, Y., B. Phares, H. Dang, and J. Dahlberg. Impact of Concrete Deck Removal on Horizontal Shear Capacity of Shear Connections. Journal of Bridge Engineering, Vol. 21, No. 3, 2016. The Standing Committee on General Structures peer-reviewed this paper.
