Structural Behavior of Hybrid Composite Beam Bridges in Missouri, USA

Transcription

1 ACIC 13 Advanced Composites in Construction 10 th 12 th September 2012, Queens University, Belfast, UK Structural Behavior of Hybrid Composite Beam Bridges in Missouri, USA Mohamed A. Abeol Seoud 1, C. Renee Earley 2, John J. Myers 3* 1,2,3 Missouri University of Science and Technology, Rolla, MO, USA ABSTRACT This project involves the Field Evaluation of three Hybrid-Composite Beam Bridges constructed in Missouri, USA. These hybrid composite beams (HCBs) are comprised of three main sub-components that are a composite shell, compression reinforcement, and tension reinforcement. In the preferred embodiment, the shell is comprised of a fiber reinforced polymer (FRP) box beam. The compression reinforcement consists of self-consolidating concrete (SCC) which is pumped into a profiled conduit within the beam shell. The tension reinforcement consists of galvanized steel tendons anchored at the ends of the compression reinforcement. The objectives of this study include the evaluation of the in-situ structural behavior of the bridge to conform to design assumptions, the development of a quality control / quality assurance testing of the bridge members, and the evaluation of potential serviceability and maintenance challenges. In order to accomplish these objectives, beam elements have been instrumented with various sensors and instrumentation. This paper discusses the behavior of the HCB elements in-situ under load testing and presents finite element modeling results. INTRODUCTION In the last decade there has been significant progress in the use of the fiber reinforced polymer (FRP) composites as construction materials in structural engineering applications. FRP composites have desirable characteristics, including high strength-to-weight ratio, fatigue resistance, dimensional stability, excellent durability because of their corrosion resistance, and low to moderate tooling costs. These advantages make FRP composites excellent candidates for use as internally reinforcement bars, externally bonded laminates for strengthening and upgrading of existing infrastructures, or main load-carrying members in civil engineering applications. The most common FRP composites that are used as main load-carrying members in civil engineering applications are glass fiber reinforced polymers (GFRP) and carbon fiber reinforced polymers (CFRP). However, these materials are still expensive in terms of initial costs in comparison with classical materials, such as concrete and steel. GFRP is less expensive than CFRP; however the use of GFRP sections independently as flexural members has some shortcomings. These shortcomings include: (1) the compressive flange is considerably weaker than the tensile flange, this is attributed to the local buckling phenomena and to the low compressive strength of GFRP relative to its tensile strength; (2) the catastrophic failure nature of GFRP sections, due to the fact that GFRP is linear elastic up to failure; and (3) the low stiffness of GFRP makes the design of GFRP sections governed by stiffness, resulting in the use of larger sections to satisfy serviceability requirements of design codes. The concept of hybrid FRP sections has emerged to overcome these problems. Some researchers combined the FRP composites with conventional materials to reduce the cost and to optimize the structure based on material properties of each constituent material. Among those are Deskovic et al. 1 Graduate Research Assistant, Dept. of Civil, Arch. & Envir. Engr. 2 Graduate Research Assistant, Dept. of Civil, Arch. & Envir. Engr. 3 Associate Professor, Dept. of Civil, Arch. & Envir. Engr., 1401 North Pine Street, 325 Butler Carlton Hall, Rolla, Missouri 65409, jmyers@mst.edu, Phone (573) * Corresponding Author

2 [1] and Van Erp et al. [2]. In their proposed hybrid beams, they replaced the upper GFRP flange with a layer of concrete, utilizing the high compressive concrete strength in resisting compression zone stresses, and exploiting the concrete stiffness to increase the overall stiffness of the member. They also added a thin layer of externally bonded carbon fiber reinforced polymer (CFRP) to the section's tension zone. Since CFRP failure strain is less than that of GFRP, the CFRP layer will be the first element to fail, giving some warning of an imminent collapse. Mirmiran s Innovative Combinations of FRP and Traditional Materials [3] provides a good review on the combination of FRP with traditional materials in structural members. This paper looks at a new type of hybrid composite beams (HCBs) that were recently used to construct three bridges (B0439, B0410, and B0478) in Missouri, USA. The underlying concept of the HCB was conceived by John Hillman in 1996, and the first commercial use was in June Hillman suggested that if a concrete arch was tied at the ends and encapsulated in a FRP shell, the embodiment would be a lightweight, strong, and corrosion resistant structural member. This unique configuration, combining conventional materials in conjunction with FRP components, creates a structural member that exploits the inherent benefits of each material in such a manner as to optimize the overall performance of the beam. Since FRP materials are generally too expensive and too flexible when arranged in a homogeneous form, the cost efficiency and stiffness are provided by an efficient use of the steel in purely axial tension, and the concrete in purely axial compression [4]. The Missouri Department of Transportation (MoDOT) is currently exploring the use of the HCB for infrastructure applications. MoDOT entrusted an exploratory evaluation program to Missouri University of Science and Technology (Missouri S&T) and University of Missouri-Columbia (UMC). The program aims to monitor and analyze the behavior of the recently constructed HCB bridges. These bridges incorporate the use of HCBs with traditional reinforced concrete deck systems. This paper describes a field load test which is conducted on the first bridge and concentrates on two finite element models (FEMs) of this bridge using two commercial finite element software packages, ANSYS V13.0 and SAP2000 V14.2. The finite element modeling of this bridge is unconventional due to the unique configuration and the hybrid nature of this new type of beams. Therefore the objective of this paper is to examine the accuracy of the finite element analysis (FEA) of the HCB. The FEMs predictions are verified by comparing them with measured field data. Then the FEMs are used to analyze the internal structural behavior of the HCBs. BRIDGE B0439 DESCRIPTION Three innovative bridges were recently constructed in Dade, Douglas, and Reynolds Counties, Missouri, USA. The first bridge, B0439, is located on MO 76 over Beaver Creek in Douglas County. Its construction was completed in November 2011, and the bridge was opened to traffic shortly thereafter. B0439 is a three-span bridge. The first and last clear spans are 59-ft 2-in. long (18.03 m), while the middle clear span is 60-ft 0-in. (18.29 m). The overall width of the bridge is 30-ft 8-in. (9.35 m). Typical cross sections of B0439 and of one of its HCBs are shown in Figure 1. The bridge consists of fifteen HCBs, five simply supported HCBs in each span. Each HCB is 33-in. x in. (83.8 cm x 61.9 cm). Each HCB consists of a self-consolidating concrete (SCC) arch with 5- in. (12.7 cm) depth and 22-in. (55.9 cm) width. The concrete arch is reinforced with two 0.5-in. diameter, 270 ksi (1862 MPa), seven wire galvanized steel strands and is tied with thirty-six 0.5-in. diameter, 270 ksi (1862 MPa), seven wire galvanized steel strands arranged in one layer. The concrete compression arch, and tension steel reinforcement are encapsulated in a GFRP shell. The concrete arch is connected to the upper GFRP flange with SCC web of varying width ranging from 3- in. (7.6 cm) to 4-in. (10.2 cm). The concrete arch is also reinforced with #5 galvanized shear connectors. The voided spaces between the arch and the FRP shell are filled with Polyisocyanurate Foam. To further expedite the superstructure construction, this bridge incorporated the use of precast stay-in-place deck forms spanning between the beams, which are spaced at 6-ft 4-in. (1.93 m) to accommodate the 30-ft 8-in. (9.35 m) out-to-out dimension of the deck.

3 Figure 1. (a) Cross Section of Bridge B0439; (b) Cross Section of a Single Hybrid-Composite Beam LOAD TESTING OF BRIDGE B0439 A research team from Missouri S&T conducted a load test of Bridge B0439 on March 26, Two MoDOT trucks were used to perform the load test on the first span of the bridge. A Leica total station and 19 prisms were used to measure deflection along each girder. An additional 3 prisms served as control points to make sure the total station did not move during the testing. One of these points is located along the top of the barrier over the abutment and will serve as a reference for comparison with future load tests. The trucks were used to perform three stops simulating three different load cases. The truck stops are shown in Figure 2. The front axle load (P1) of the first truck (T-6732) is 15.2 kip (6.9 metric tons) and the rear axle load (P2) is 35 kips (15.9 metric tons); the front load (P1) of the second truck (T-7627) equals 15.5 kips (7 metric tons) and the rear load (P2) equals 32.2 kips (14.6 metric tons). The truck configuration and distribution of the loads are also displayed in Figure 2. 1st Stop Line 2nd Stop Line 3rd Stop Line 1st Stop Line 2nd Stop Line 3rd Stop Line Girder 1 Girder 1 T-6732 T-7627 Girder 2 Girder 2 Girder 3 Girder 4 T-6732 T-7627 Girder 3 Girder 4 Interior Bent STOP 1 Truck Center of Gravity Abutment Girder 5 STOP 2 15" 15" Girder 5 Girder 1 12" T-6732 Girder 2 Girder 3 82" P1 0.4 P2 0.6 P2 75" T-7627 Girder 4 25" 1' 28'-7" 27'-7" 2' STOP 3 Girder 5 193" 55" Figure 2. Trucks Stop Locations and Trucks Dimensions FINITE ELEMENT MODELING OF BRIDGE B0439 Two commercial finite element software packages, ANSYS V13.0 and SAP2000 V14.2, were used to model bridge B0439. The longitudinal direction of the bridge is modeled on the X-axis in both FEMs, while the vertical (gravity) direction is presented by Y-axis in the ANSYS model and by the Z-axis in

4 the SAP model. The transverse or lateral direction is presented by the Z-axis in ANSYS and the Y-axis in SAP2000. The maximum deflection measured during the load test of B0439 was found to be in. (0.13 cm) and occurred at midspan due to Stop 3. This very small value indicated that there is no need to perform nonlinear geometric analysis and gave an indication that all the materials may behave within their linear elastic ranges. Consequently, the first trial FEA, in which all the materials are modeled as linear elastic, is performed using the two software packages. The results obtained from the two FEMs prove that the linear behavior assumption is valid. Material Properties FRP Composites The FRP composites are anisotropic materials; that is, their properties are not the same in all directions. However, the unidirectional lamina that has fibers oriented only in one direction possesses three orthogonal planes of material symmetry. It is called an especially orthotropic material when the x-direction is the same as the fiber direction [5]. The relation between the stress tensor and the strain tensor of an orthotropic material can be represented by Equation 1. (1) Where S is the compliance matrix, and the S ij coefficients as defined by Gibson [5]. Most of the unidirectional laminates for structural applications have fibers with circular cross-section and hence can be assumed as orthotropic transversely isotropic, because the properties of these composites are nearly the same in any direction perpendicular to the fibers. For orthotropic transversely isotropic materials, the strain tensor can be related to the stress tensor via the compliance matrix as shown in Equation 2 [6] (2) Where ε is the strain in the direction i, γ is the shear strain in the plane ij, σ is the normal stress in the direction i, τ is the shear stress in the plane ij, E i is the young s modulus in the direction i, G ij is the shear modulus in the plane ij, ϑ is the Poisson s ratio (the ratio of strain in the j direction to strain in the i direction when the applied stress in the i direction), i,j= x, y, z and i j. In Equation 2, the axis of isotropy is along the x-axis (direction of fibers in case of unidirectional laminates). From the relationship in Equation 2, it can be concluded that E y =E z, xy, xz yz, G zy xy = G xz, Gyz E / y 21yz, and thus reducing the number of independent elastic constants from nine in case of orthotropic materials, to five in the case of orthotropic transversely isotropic materials. The multidirectional fabrics begin to behave somewhat quasi-isotropic, allowing these composites to be assumed to have isotropic properties and simplifying preliminary designs. The standard laminate composition of the HCB FRP shell is typically a quadweave glass reinforcing fabric infused with a vinylester resin matrix. Experimental tests were performed by the manufacturer at the macroscopic level on the laminate that is used in the shell webs to identify the independent elastic constants of the

5 shell. These tests identified the tensile and compressive in-plane moduli of elasticity (E x +, E x -,E y +,E y - ) and the in-plane shear modulus (G xy ). The test results showed that the tensile elastic modulus in X- direction differs significantly from the tensile elastic modulus in Y-direction. To take into account this difference the FRP shell is assumed to be orthotropic transversely isotropic. The allowable results enable the calculation of only three elastic constants; in order to calculate the remaining two constants, is assumed to be 0.26 and is assumed to be 0.30 [7]. A summary for material properties used for modeling the FRP shell is listed in Table 1. Tensile properties Compressive properties Density Ib/ft 3 (kg/m 3 ) Table 1. Material Properties Used for Modeling the FRP Shell Elastic Modulus Shear Modulus Major Poisson s ratio ksi (MPa) ksi (MPa) Concrete Concrete is a quasi-brittle material and has different behavior in compression and tension. In compression the stress-strain relation is linearly elastic up to approximately 30% of the maximum compressive strength [7]. In tension the stress-strain curve is approximately linearly elastic up to the maximum tensile strength. However the tensile and compressive moduli of elasticity are almost the same in the elastic linear range. In Bridge B0439, self-consolidating concrete (SCC) was used to form the concrete arch of the HCBs. SCC is a new category of high-performance concrete that is used to improve the productivity of casting congested sections and to insure the proper filling of restricted areas with minimum or no consolidation [8]. Data from more than 70 recent studies on the hardened mechanical properties of SCC have been analyzed and correlated by Domone [9] to produce comparisons with the properties of equivalent strength normally vibrated concrete (NVC). Domone concluded that the design rules and practice for NVC developed over many decades can be used for structures cast with SCC. Based on the Domone study, traditional equations that are used with NVC are used in the current study to calculate the concrete properties. The drawings of the three HCB bridges indicated that the concrete arch of a HCB should have a minimum compressive strength of 6 ksi (41.4 MPa), however the field tests showed that the compressive strength of the concrete arches of the HCBs used in the second and third bridges (B0410 and B0478) have average compressive strengths that are much higher than the desired values. Unfortunately, no data is available about the compressive strengths of the HCBs used in the first Bridge B0439. The compressive strength of the HCBs concrete arches are assumed to equal the average of the compressive strength of the HCB concrete arches of the second and third bridges, which is 10 ksi (68.9 MPa). The concrete of the deck is assumed to have an average compressive strength of 6 ksi (41.4 MPa) for the precast prestressed panels and the cast in place concrete. Then the elastic modulus and the maximum tensile strength are calculated by the following ACI procedure [10]: (ASE) (3) 7.5 (ASE) (4) Where is the compressive strength of concrete, is the modulus of rupture of concrete and is the elastic modulus. In equations 3 and 4,, and are in psi. Since the reinforcement bars are modeled explicitly in the FEMs, the density of concrete is assumed to be 140 lb/ft 3 (2243 kg/m 3 ).

6 Steel Reinforcement Two types of reinforcement bars are used in bridge B0439. Typical Grade 60 steel reinforcing bars are used in the deck, while Grade 270 seven wire, galvanized steel strands are used in the HCBs. Both types are assumed to be identical in tension and compression with 29,000 ksi (199,948 MPa) young s modulus, 0.3 Poisson s ratio and 490 lb/ft 3 (7849 kg/m 3 ) density. Polyisocyanurate Foam Polyisocyanurate foam is a 2.0 lb/ft 3 (32 kg/m 3 ), rigid, closed cell foam supplied as blocks with 24-in (61 cm) width. The tensile and compressive elastic moduli and shear moduli are provided by the manufacturer in the longitudinal and perpendicular directions. The Foam is assumed to behave as orthotropic transversely isotropic material. Poisson s ratios are calculated based on the provided moduli, while and are assumed to be A summary for material properties used for modeling the Polyisocyanurate Foam is listed in Table 2. Table 2. Material Properties used for Modeling the Polyisocyanurate Foam Elastic Modulus Shear Modulus Major Poisson s ratio psi (KPa) psi (KPa) 12258, Tensile properties , Compressive properties Element Types and Model Simplifications Hybrid Composite Beam (HCB) Since the GFRP shell of the HCB has a very small thickness relative to its length and width, it is modeled using a shell181 element in ANSYS and a traditional shell element in SAP. Both elements are four-node elements with six degrees of freedom (DOFs) at each node: translations in the x, y, and z directions, and rotations about the x, y, and z-axes. The compressive FRP properties are assigned to FRP webs and upper flange elements while the tensile FRP properties are assigned to the lower flange elements. The FRP webs are modeled with thickness equals in. (0.48 cm), the upper flange is modeled with constant thickness equals 0.39 in (0.99 cm), while the lower flange is modeled with constant thickness equals 0.30-in. (0.77 cm). The concrete web is also modeled using shell181 element and shell element in ANSYS and SAP respectively. The web is modeled with constant thickness equals 4-in. (10.2 cm). The concrete arch is modeled using a solid65 element in ANSYS and a traditional solid element in SAP. Both elements are defined by eight nodes having three translational DOFs at each node. Figure 1-b shows that the clear space between the FRP webs is 24-in. (61 cm), while the concrete arch width is 22-in. (55.9 cm), the space between the concrete arch and the FRP webs is filled with 1-in. (2.54 cm) of the polyisocyanurate foam per each side. Modeling this foam may result in a very poor mesh; consequently the solution may have difficulties in convergence. To avoid the probability of divergence, the concrete arch is modeled with 24-in. (61 cm) width. The polyisocyanurate foam above and below the concrete arch web is modeled using solid185 element in ANSYS and traditional solid element in SAP. Both elements are defined by eight nodes having three translational DOFs at each node. Foam tensile properties are assigned to the elements below the concrete arch, while the compressive properties are assigned to the elements above the concrete arch. For the upper and lower foam axis of isotropy is taken along the gravity direction (Ydirection in ANSYS and Z-direction in SAP).

7 The 36 tension tie strands and the 2 arch reinforcement strands are modeled using beam188 elements in ANSYS and traditional beam elements in SAP. Figure 1-b shows that the 36 strands lie directly on the lower FRP flange. Since the strands are modeled using one-dimensional space element (beam element) in both models, if they are modeled directly on the lower flange, the flange will behave as if it is reinforced with these strands, which is not accurate. In order to maintain the actual behavior of the lower flange, the height of the concrete end block is increased 1-in. (2.54 cm), shifting the flange downward, and the strands modeled in two layers separated from the lower flange and from each other by 0.5-in. (1.27 cm). Figure 4 demonstrates the modification of the end block and the modeling of the tension reinforcement. Modeling the 36 strands via 36 separate beam elements would complicate the meshing, therefore the 36 strands are represented by 6 beam elements, 3 per layer. The total cross-sectional area is equal to the cross-sectional area of the 36 strands. The tension reinforcement is extended in the concrete end block to restrain the deformation of the arch in the longitudinal direction. This is simulated in the FEMs by modeling vertical beam elements in the end block and they were connected to the horizontal tensile reinforcement through rigid joints to allow transferring the force from the arch to the reinforcement. Modeling the tension reinforcement is shown in Figures 3 and 4. The different components of the HCB are assumed to be perfectly bonded to each other. This was achieved in the two FEMs by maintaining the same meshing for all the constituents. Maintaining the same meshing guaranteed that the joints of any component coincide with the joints of other components that are in intimate contact with this component, and consequently achieving the assumption of the perfect bond. The HCBs of bridge B0439 were manufactured with an initial 5.25-in. (13.36 cm) camber to equilibrate the downward deflection of the beams when subjected to the bridge s full dead load. All the HCB components are modeled with the aforementioned camber. Figure 3 displays the finite element modeling of Hybrid-Composite Beam using SAP2000 V14.2. Compression Reinforcement (Concrete Arch) FRP Upper Flange Reinforcement of Concrete Arch Concrete Web FRP Web Concrete End Block 5.25 camber Polyiso Foam Tension Reinforcement FRP Lower Flange Figure 3. Finite Element Modeling of Hybrid-Composite Beam using SAP2000 V14.2 Bridge Deck Due to the HCBs camber, the thickness of the bridge slab varies from 8.5-in. (21.6 cm) to 11-in. (27.9 cm). The slab is modeled with constant thickness 10-in. (25.4 cm) using a solid65 element in ANSYS and a traditional solid element in SAP. Two solid elements are used throughout the slab, one element with a 4-in. (10.2 cm) height and the other with a 6-in. (15.2 cm) height. This allows the modeling of reinforcement bars at 4 in (10.2 cm) from the bottom of the slab as indicated in the bridge drawings. The reinforcement bars in longitudinal and transverse directions are modeled using beam elements in both models. The spacing of bars is controlled by the meshing of deck in each model; however the cross-sectional area of the beam element is modified to maintain the existing total reinforcement area in the deck. The parapet was poured simultaneously with the slab, and its reinforcement extended in the deck. Previous study showed that when a composite action is achieved between the slab and the parapet, the deflection is significantly decreased [11]. Consequently the parapet is included in the FEMs and simulated using a solid65 element and a traditional solid element in the ANSYS and

8 SAP2000 respectively. Perfect bond is assumed between the deck components and between the deck and the HCBs. This is achieved by making the joints of different components coincident at the contact regions. Figure 5 displays modeling of the bridge deck and the HCBs in ANSYS. End Block Concrete Web Bridge Slab (6 thickness) Parapet Parapet RFT Bridge Slab (4 thickness) End Block Modification FRP Lower Shell Concrete Arch Tension Reinforcement RFT in longitudinal and transverse directions HCBs Figure 4. Model of the HCB Tension Reinforcement in ANSYS V13.0 Figure 5. Model of the Bridge Deck and the HCBs in ANSYS V13.0 Load Modeling and Boundary Conditions The load was simulated in the two models as uniform distributed loads applied on the solid elements that represented the deck. The distributed load is calculated based on the tire loads, the tires contact area with the deck and the wearing surface thickness which is assumed to be 1-in. (2.54 cm). The distributed load area is calculated by adding the tire dimension in each direction to two times the wearing thickness. However the load distribution and locations are controlled by the deck meshing and are slightly modified from that are shown in Figure 2. For example, the load of the rear dual tires should applied to an area of 27-in. x 17-in. (68.6 cm x 43.2 cm), but due to meshing it was applied to an area equals 26-in. x 17.5-in. (66 cm x 44.5 cm). Some researchers, such as Kachlakev et. al [7], modeled the tire loads as point loads to accurately model the load locations without the need to change the mesh. However, applying the tire load as a distributed load has apparent advantages over modeling it as a concentrated load. First, using 3D elements is non-consistent with concentrated loads, because this may result in a stress singularity, as the element sizes become smaller and smaller as the stress increases and tends to infinity, hence mesh convergence can t be achieved. Modeling the tire loads as distributed loads avoids the probability of stress concentration and leads to mesh convergence. Second, applying the tire loads as distributed loads simulates the practical situation more accurately than the concentrated loads; consequently it doesn t overestimate the deflection, which may occur in the case of using concentrated loads. To represent the bearing pads, a 1-in. (2.54 cm) thick steel plate is modeled beneath the end block. The plate is modeled using a solid185 element in ANSYS and a traditional solid element in SAP, then the lower joints of the plate are restrained from translation in the X, Y and Z directions. Applying the hinged supports to the steel plate instead of applying them to the beam directly aims to avoid stress concentration and providing better stress distribution over the support area. Due to the continuity of the bridge slab, the edge of the slab between the HCBs is restrained from translation and rotation in all directions over the interior bent, while it is restrained from translation in all direction over the abutment. The bridge incorporated the use of diaphragms at the end of the HCBs. To represent the effect of these diaphragms, the translation of the end block in the lateral direction (Z-direction in ANSYS and Y- direction in SAP) and the vertical displacement of the slab over the diaphragms are restrained.

9 RESULTS DISCUSSION A comparison between a sample of the measured deflections and the predicted ones by ANSYS V13.0 and SAP2000 V14.2 is shown in Figure 6. Only the results of Stops 1 and 3 are shown due to space limitations. Deflection is measured and predicted at the quarter, mid, and three-quarter points of each girder span: 14-ft 9.5-in. (4.5 m), 29-ft 8-in. (9 m), and 44-ft 4.5-in. (13.5 m) respectively from the center line of the interior bent. In general, the results demonstrate a good agreement between the two FEMs predictions and the field measured data and an excellent matching between the two FEMs results. Deflection (in) STOP1 G1 Field SAP ANSYS Deflection (in) STOP1 G3 Field SAP ANSYS 0 14' 9.5" 14' 9.5" 44' 4.5" Distance from interior bent (a) 0 14' 9.5" 14' 9.5" 44' 4.5" Distance from interior bent (b) Deflection (in) STOP1 G5 14' 9.5" 14' 9.5" 44' 4.5" Distance from interior bent (c) Field SAP ANSYS Deflection (in) STOP3 G1 14' 9.5" 14' 9.5" 44' 4.5" Distance from interior bent (d) Field SAP ANSYS Deflection (in) STOP3 G2 Field SAP ANSYS 14' 9.5" 14' 9.5" 44' 4.5" Distance from interior bent (e) Deflection (in) STOP3 G3 Field SAP ANSYS 14' 9.5" 14' 9.5" 44' 4.5" Distance from interior bent (f) Figure 6. Comparison of the Bridge B0439 Deflections Measured at Field and Predicted by ANSYS and SAP2000 The differences between the measured field data and the FEMs predictions can be attributed to the unknown material properties, modeling simplifications, experimental errors and unknown thermal effects. A previous study by Myers et al. [11] summarized the effect of the experimental errors in the bridges field test as follow: a transverse shift in the truck stop location by 12-in. (30.48 cm) changes the deflection by 10%, a longitudinal shift in the truck stop location by 12-in. (30.48 cm) changes the deflection by 5%, 5 kips (2.27 metric tons) error in truck weight reporting alters the deflection by 7%, and the sensitivity of surveying equipment is ± in. (0.013 cm), which equals 10% of the maximum deflection occurred due to Stop 3 and higher percentages for all other deflection values. (e)

10 Some of these errors were incarnated in Stop 2, the results of which are not shown due to space limitations. As shown in Figure 2, the two trucks loads are symmetric about the longitudinal center line of the bridge (centerline of G3), consequently deflections are expected to be the same along girders 1 and 5, and girders 2 and 4. However the experimental data showed that the deflections through G2 are larger than those through G4 by an average value of 28%, and the average difference between G1 and G5 measurements is 12%, which indicates that the two trucks shifted transversely toward G1 and G2. In most cases the two models overestimated the deflection at the girder midspans in the three stops. One of the potential contributing factors is the simplification in modeling the arched shape of the compression reinforcement in the HCBs. Modeling the concrete arch shape exactly would result in a very poor mesh of the upper polyisocyanurate foam elements in the region close to the girder midspan and of the FRP webs elements that are in contact with those elements. To avoid a poor mesh, the middle part of the arch is modeled as straight surface as it is shown in Figure 7. Another factor that may contribute to this difference is the thermal effects. The bottom surfaces of the HCBs are closer to the river and are not exposed to the sun which may reduce their temperature. On the other hand, the upper surfaces can absorb heat from the deck which is exposed to the sun continuously during the daytime. This thermal gradient causes upward deflection, consequently reducing the total deflection at midspan. According to Radolli and Green [12] the stresses induced through the depth of the structure due to diurnal temperature cycles can exceed the live loading in some cases. Since the field test was performed in March, it is not expected that the thermal effects resulted in such high stresses. But it may result in significant experimental stresses that deviate from the analytical predictions, especially because the thermal behavior of the HCB is still unknown. Figure 6-c illustrates that the FEMs underestimated significantly the deflections through G5 in Stop 1 in comparison to the measured ones. It is clear from Figure 2 that the truck loads transferred to G5 in Stop 1 tend to be zero. Consequently, this would result in very small deflections beyond the accuracy of the total station. The deflections through G5 span ranged from in. (0.004 cm) to in. (0.011 cm) according to ANSYS results, while the tolerance of the total station as observed by Myers et al. [11] is ± in. (0.013 cm). This means that the total station measurements are not reliable in this case and FEMs can predict the deflections more accurately. (a) (b) Figure 7. Simplification in Modeling the Compression Reinforcement of the HCB (a) ANSYS Model; (b) SAP2000 Model STRUTURAL BEHAVIOR ANALYSIS The stresses obtained by SAP2000 are similar to those obtained by ANSYS, however the ANSYS stresses are found to be always higher than the ones predicted by SAP2000, This can be attributed to the difference in mesh densities between the both models. The bridge deck is modeled in Ansys using very fine mesh, while it is modeled in SAP2000 with relatively coarse mesh in comparison to the Ansys one. Since the Ansys model has finer mesh and consequently is expected to provide more accurate results, all the stress results presented in the following sections are extracted from the ANSYS model due to the Stop 3 load case. The presented stresses of the different HCB components are the major stresses in the longitudinal direction of the beams (X-direction). These stresses are much higher than the stresses in the Y and Z-directions. Results clarified that the maximum compressive stress induced in the concrete arch due to Stop 3 plus the self-weight of the bridge occurred at the connection of the concrete arch with the end block of G3. The stress is about 1600 psi (11.03 MPa), which is less than 30% of the assumed compressive

11 concrete strength of 10 ksi (68.9 MPa). While the maximum tensile stress induced in the concrete arch due to the same load combination took place at the arch midspan of G3. The tensile stress is 350 psi (2.41 MPa) which is less than the modulus of rupture of concrete, 750 psi (51.7 MPa), calculated via Equation 4. This indicates that the classical arch shape of concrete that is used in the HCB optimizes the use of the concrete and preserves its overall stiffness. Figure 8 displays the stresses in X-direction in the concrete arch of G3 obtained by ANSYS and SAP2000 models. The results also illustrated that due to the boundary conditions and the HCB camber, the FRP shell (lower flange, upper flange, and webs) and the tensile reinforcement have tensile stresses at the mid span, while they have compressive stresses at the end. The maximum stress in the tensile reinforcement of the HCB is found to be 1260 psi (8.69 MPa). The maximum compressive stress in the lower FRP shell is 1060 psi (7.31 MPa) and the maximum tensile stress is 1000 psi (6.89 MPa) which are much lower than the effective longitudinal tensile and compressive strengths of the GFRP determined from the experimental tests on the laminate specimens. The stresses in the FRP upper flange and webs are found to be lower than the stresses in the lower flange due to this load case. These values confirmed that all the materials behave within their linear elastic range and indicate that the total load due to the bridge self weight in addition to the two loaded trucks oriented as in Stop 3 are much lower than the capacity of the HCBs. The stresses in the bridge deck are also found to within the elastic range of the concrete and reinforcement bars. The maximum measured deflection 0.05 in (0.127 cm) due to Stop 3 is much lower than the allowable live load deflection provided by the American Association of State Highway and Transportation Officials Load Resistance Factor Design (AASHTO LRFD) of length/800 [13], which is 0.89-in. (2.25 cm). This is compatible with the stress indication that the applied load is much lower than the ultimate capacity of the HCBs. The results also demonstrated that the forces carried by the concrete arch at the mid span and the end of the HCB are several times larger than those carried by the tensile reinforcement and the FRP shell, which are similar in magnitude. However it is expected that when the concrete starts to crack, the tendons and FRP will carry much higher loads. It is also expected that after the cracking of concrete, the lower FRP flange will contribute to the flexural capacity, while the FRP webs will contribute to the shear capacity. This is similar to the contribution of FRP strengthening laminates to the capacity of the reinforced concrete beams. (a) (b) Figure 8. Stress in X-direction in the Concrete Arch of G3 Due to Stop 3 Plus Bridge Self-Weight (a) ANSYS Model; (b) SAP2000 Model The results show that the HCBs underwent lateral and rotational deformations under the vertical truck loads in the three stops. The value of the maximum lateral displacement occurred in FRP webs in G1 and G5 and is found to be about 20% of the maximum vertical deflection in the three load cases, and about 11% of the vertical deflection due to Stop 3 plus the self-weight of the bridge. This indicates that the HCB may have weak lateral and torsional stiffness. Further study needs to be performed to investigate the lateral behavior of HCBs under lateral loads, such as wind and seismic loading, and the torsional behavior of horizontally curved HCBs as well as straight HCBs subjected to torsional loads. Figure 9 displays the lateral and rotational deformations of the HCBs due to self weight of the bridge and Stop 3 loads. A study was performed to identify the contribution of the polyisocyanurate foam to the lateral stiffness of the HCB. It was found that when the foam is removed from the HCBs, the FRP shell suffered local buckling and instability and the FRP webs underwent large deformations. Consequently, the exact contribution of the foam to the lateral and stiffness could not be precisely determined. But the previous study proved that the foam plays an important role for the stability of the FRP shell because it prevents the local buckling and wrinkling of shell and provide lateral stability to the FRP webs. This gives the foam in the HCB a similar role to that of the core in an FRP sandwich

12 structure. Figure 10 shows exaggerated deformation shapes of FRP lower flange of HCBs without polyisocyanurate foam due to different applied loads. Lateral and rotational deformations of the FRP web (a) Figure 9. Exaggerated Deformed Shape of HCBs Due to Self-Weight plus Stop 3 (a) ANSYS Model; (b) SAP2000 Model (b) Wrinkling of the FRP shell Figure 10. Exaggerated Deformed Shapes of FRP Lower Flange Without Polyisocyanurate Foam Due to Different Loads CONCLUSIONS AND RECOMMENDATIONS This paper describes a load test applied to an innovative superstructure bridge system recently constructed in MO, USA. The superstructure incorporates the use of innovative HCBs with a traditional reinforced concrete deck system. A linear FEA of the bridge superstructure is carried out using two finite element packages, ANSYS V13.0 and SAP2000 V14.2. The deflections predicted by the two FEMs are compared with the field test measurements, and the behavior of the innovative HCB is analyzed. The linear FEA of the bridge superstructure is found to be accurate and can be used to analyze the structural behavior of the new HCB. The predicted deflections by the two FEMs have excellent matching with each other and are in good agreement with the experimental measurements. The model simplifications presented in this work are proved to be acceptable and led to good predictions. However, the simplification of the classical arch shape of the compression reinforcement led to an increase in the predicted deflections in comparison with the measured ones, and it is recommended to maintain the arch shape as much as possible during modeling of the HCB. Additionally, thermal effects are a suspected contributor to the difference between the measured and predicted deflections. Further study of the thermal behavior of the HCB is recommended. The maximum measured deflection due to the three different load cases is less than 6% of the permissible live load deflection provided by ASSHTO LRFD [13]. This proves that the new HCB possesses a sufficient flexural and shear rigidity that avoids excessive deflections under service loads. However, the girders suffered some lateral and torsional deformations under the vertical loads, indicating that the HCB may have low lateral and torsional rigidity. Deep study of the HCB behavior under lateral forces and torsional moments is recommended.

13 The predicted stresses in the components of the HCB are within the linear elastic range of the different materials and are much lower than the ultimate strength of each material. These stresses, in addition to the measured deflection, clarify that the applied loads are much lower than the ultimate capacity of the HCBs. The classical arch shape is proved to optimize the use of the concrete and to preserve the overall stiffness of the HCB under service loads. Also, the low stress carried by the FRP shell, due to the combination of bridge self weight and the loads of the stop with maximized moment (Stop 3), maintains the ability for long-term durability of the shell, increasing the lifetime of the HCB as a whole. Finally, the polyisocyanurate foam is proved to contribute to the lateral and torsional stiffness of the HCBs, to prevent local buckling and wrinkling of FRP shell, and to provide lateral stability to the FRP webs. AKNOWLEDGEMENTS The authors would like to acknowledge the Missouri Department of Transportation and the National University Transportation Center (NUTC) at Missouri S&T for sponsoring this research study. The staff support from the Dept. of Civil, Architectural & Environmental Engineering and Center for Infrastructure Engineering Studies (CIES) at Missouri S&T is also greatly appreciated. REFERENCES 1. Deskovic, N., T.C. Triantafillou, and U. Meier, Innovative Design of FRP Combined with Concrete: Short-Term Behavior. Journal of Structural Engineering. 121(7): (1995) 2. Van Erp, G., et al. An Australian Approach to Fibre Composite Bridges presented at the International Composites Conference ACUN4, Composite Systems: Macro Composites, Micro Composites, Nano Composites, UNSW Sydney Mirmiran, A. Innovative Combinations of FRP and Traditional Materials presented at the International Conference on FRP Composites in Civil Engineering, Hong Kong, China, December Hillman, J.R., Product Application of a Hybrid-Composite Beam System, Gibson, R.F., Principles of Composite Material Mechanics: CRC Press Lubarda, V.A. and M.C. Chen, On the Elastic Moduli and Compliances of Transversely Isotropic and Orthotropic Materials. Journal of Mechanics of Materials and Structures. 3(1): (2008) 7. Kachlakev, D., et al., Finite Element Modeling of Reinforced Concrete Structures Strengthened with FRP Laminates. Final Report SPR. 316, Khayat, K., Workability, Testing, and Performance of Self-Consolidating Concrete. ACI Materials Journal. 96: , Domone, P., A Review of the Hardened Mechanical Properties of Self-Compacting Concrete. Cement and Concrete Composites. 29(1):1-12, Committee, A., A.C. Institute, and I.O.F. Standardization. Building Code Requirements for Structural Concrete (ACI ) and Commentary. American Concrete Institute Myers, J.J., Holdener, D., Merkle, W., Hernandez, E., Preservation of Missouri Transportation Infrastructures: Validation of FRP Composites Technology Through Field Testing In-situ Load Testing of P-962, T-530, X-495, X-596 and Y-298, Missouri Dept. of Transportation Report, Missouri University of Science and Technology, Rolla, Missouri, Radolli, M. and Green, R., Thermal Stresses in Concrete Bridge Superstructures under Summer Conditions, Transportation Research Record, Transportation Research Board, No. 547, AASHTO, L., Bridge Design Specifications, 2012, American Association of State Highway and Transportation Officials, Washington, DC, 2012.