Use of UHPC in Critical Shear Span of CFRP Prestressed Bridge Girders 1

Nabil F. Grace1, Ranjit K. Sharma2, Soubhagya K. Rout3, Kenichi Ushijima4, Mena Bebawy5 2

Abstract 3

An experimental program was conducted to access the shear performance of innovative carbon 4

fiber composite cable (CFCC) prestressed beams. Two beams were built using an innovative 5

hybrid technique; wherein the span of the beam subjected to higher shear stresses (near the 6

support regions) was built using ultra high performance concrete (UHPC) without any transverse 7

reinforcement. The remaining span (mid region) was constructed using normal high strength 8

concrete (HSC) and reinforced transversely with steel stirrups. The beam-ends of each beam 9

were tested under shear loading mechanism to failure at different shear-span-to-depth (a/d) ratios 10

of 3, 4, 5 and 6. The outcomes of the tested beam-ends were used to estimate the performance of 11

UHPC in distributing the shear stresses and also the behavior of the joint region consisting of 12

HSC and UHPC. Further, the results were compared with other beams provided with CFCC and 13

steel stirrups to get a conclusion on the shear behavior of these hybrid beams. It was observed 14

that hybrid beam at a/d ratios of 3 and 4, the beam-ends failed in shear due to diagonal tension 15

while at higher a/d ratios of 5 and 6, the beam failed in flexure due to crushing of top flange 16

concrete. In addition, the observed ultimate shear resistances of the beam-ends were significantly 17

higher when compared to the beams provided with transverse reinforcement (steel and CFCC) 18

and also exhibited higher energy absorption at failure. Finite Element (FE) models were 19

developed for the tested beam-ends to validate the experimental findings and provide a better 20

conclusion about the overall shear performance of the tested beams. The results of this 21

investigation suggests that the hybrid beams constructed with UHPC without any stirrups near 22

the support region and conventional concrete near the mid span are able to provide higher shear 23

resistance compared to transversely reinforced beams at both service limit state (SLS) and 24

ultimate limit state (ULS) with an appreciable level of warning at failure. 25

26

CE Database subject headings 27 Ultra High Performance Concrete (UHPC), Carbon Fiber Composite Cable (CFCC), Prestressed, 28

Shear. 29

1 Dean and University Distinguished Professor, College of Engineering, Lawrence Technological University, Southfield, MI,

48075, U.S.A., ngrace@ltu.edu 2 Graduate Research Scholar, Center of Innovative Materials Research (CIMR), Lawrence Technological University, Southfield,

MI, 48075, U.S.A., rsharma@ltu.edu 3 Former Graduate Research Scholar, Center of Innovative Materials Research (CIMR), Lawrence Technological University,

Southfield, MI, 48075, U.S.A., srout@ltu.edu 4 Senior Engineer, Cable Technologies North America, Inc., Novi, MI, U.S.A. ken.ushijima@ctnainc.com 5 Research Scientist, Center of Innovative Materials Research (CIMR), Lawrence Technological University, Southfield, MI,

48075, U.S.A., mbebawy@ltu.edu

2

INTRODUCTION 30

Due to the involvement of large number of variables and complex shear transfer mechanisms, 31

shear behavior of concrete members is not perfectly understood even after decades of 32

experimental research and latest use of highly sophisticated computational tools. Usually, a 33

sudden collapse of concrete structure without any appreciable level of warning occurs if the 34

structure lacks proper and adequate shear reinforcement (Mitchell et al. 2011). The key to avoid 35

such catastrophic brittle shear failure of concrete structures is to overdesign against anticipated 36

shear with a high factor of safety. The traditional practice is to provide transverse reinforcement 37

in the form of stirrups at closer spacing. This overly reinforced section for shear stresses possibly 38

changes the catastrophic shear failure of a member into a more favorable flexural failure 39

characterized with sufficient warning in terms of large noticeable deflection and cracking prior to 40

collapse. However, there are several issues associated with the use of stirrups; a) shear capacity 41

of section increases directly with the decrease in the stirrup spacing, but various standards & 42

codes has limitation on the minimum spacing of stirrups to avoid congestion of reinforcement 43

cage leading to improper concrete placement and consolidation; b) limitation on the maximum 44

spacing of stirrups to avoid wider shear cracks especially in prestressed concrete sections; c) 45

susceptibility of steel stirrups towards corrosion d) CFCC stirrups experience significant 46

reduction in tensile strength due to bend effect (ACI 440 1R-06). 47

The use of steel fibers in the concrete mix design has emerged as an alternative against 48

the use of stirrups in the concrete members. Ultra High Performance Concrete (UHPC) has 49

emerged as a latest fiber reinforced concrete commercially manufactured by Lafarge North 50

America under the name Ductal® and provides innovative solutions to several challenges 51

currently faced by the US highway infrastructure. AFGC-SETRA (2002) defines UHPC as 52

concrete matrix having compressive strength above 21.7 ksi (150MPa) and internally reinforced 53

with fiber to ensure non-brittle behavior, with very low water to cementitious material ratio and 54

with minimal or no coarse aggregates. 55

Taylor et al. (2011) conducted a life cycle cost analysis for bridge girders and 56

recommended that UHPC are expected to provide at least twice the service life and low cost 57

maintenance as expected from the conventional strength concrete, thus compensating the high 58

initial investment on construction. Further, a significant improvement in mechanical properties 59

of UHPC over HSC is mainly due to the presence of steel fibers and their orientation with 60

3

respect to the direction of stress. Kim et al. (2008) conducted several studies using photographic 61

technique and four point bending test to evaluate the effect of placing and flow direction on fiber 62

orientation, dispersion and tensile behavior of UHPC. Their studies showed that placing and 63

direction of flow results in a significant difference of about 50% in the UHPC maximum tensile 64

strength. Favorable properties are generally obtained when the flow of UHPC is oriented parallel 65

to the direction of the principal tensile stresses. 66

Steel fibers are capable of increasing the shear capacity and most often up-to the nominal 67

flexural capacity (Russo et al. 1991) which ultimately leads to more of a ductile shear-flexural 68

failure. The ultimate shear resistance and failure mode depends upon the percentage of steel 69

fibers. According to Imam et al. (1997) there is a major increase in the shear capacity with the 70

increase in the fiber content compared to the increase in its nominal flexural capacity (Mn). 71

Padmarajaiah and Ramaswamy (2001) conducted a rigorous experimental and analytical work 72

on 13 fully/partially steel prestressed high-strength concrete beams to study the influence of fiber 73

content, location of fiber, and the presence/absence of stirrups within the shear span on the shear 74

behavior of the beam. It was reported that beams having fibers located only within the shear span 75

and over the entire cross-section, showed similar load-deformation and ultimate load response to 76

that of beams, which had fibers over the entire span length. The presence of fibers within the 77

shear span altered the brittle shear failure to more of a ductile flexure failure. Thus, it was 78

recommended that the stirrups can be replaced with an equivalent amount of fibers without 79

compromising the overall structural performance of the member. 80

US bridges is presently facing the biggest challenge of its faster rate of deterioration and 81

spalling of concrete, which are primarily caused due to the corrosion of steel reinforcement as 82

the consequence of the cracks which are formed in bridge decks or girders due to increase in 83

traffic load. Thus, FRP prestressed bridges have come up as one of the feasible solution. 84

However, according to Park and Naaman (1997), reported that fiber reinforced polymer (FRP) 85

prestressed beams are susceptible to shear-tendon rupture failure which is a unique mode of 86

failure due to rupture of FRP tendon caused by dowel shear acting on the shear-cracking plane. 87

This was due to the FRPs brittle behavior and low transverse shear resistance. They 88

recommended that addition of steel fibers in the concrete section could possibly depress this 89

unique shear-tendon rupture failure of FRP prestressed beams by improving the shear capacity of 90

the section. 91

4

Further unlike steel, FRP possess a linear stress-strain behavior and do not show any 92

yielding behavior prior to failure. Irrespective of reinforcement ratio and irrespective of location 93

of load/mode of test, FRP reinforced/prestressed beams show a catastrophic brittle failure with 94

low ductility. Grace et al. (2012) conducted a flexural test on both under-reinforced and over-95

reinforced FRP prestressed decked bulb-T HSC beam. In both of these kinds of beams, a 96

catastrophic flexural failure was obtained with ductility ratio below the limit of 75% for ductile 97

flexural failure. Similarly, Grace et al. (2014) carried out extensive study on the modes of shear 98

failure on over-reinforced FRP prestressed decked bulb-T HSC beams transversely reinforced 99

with either steel/CFCC stirrups at 6 in. (152.4 mm) of spacing at varying shear span to depth 100

(a/d) ratio of 3, 4, 5 & 6 and found catastrophic shear failure mode in all beams, irrespective of 101

location of the load. 102

This paper explains an innovative approach for the optimum use UHPC through a 103

concept of hybrid formulation. The critical shear spans where the shear stresses are dominant 104

(near the support regions) are built with UHPC without stirrups, while the remaining middle span 105

of the beam where flexural stresses are dominant are built with HSC with steel stirrups at 4 in. 106

(102 mm) spacing. The main purpose of using UHPC in the critical shear span of beams were: a) 107

to eliminate the use of steel or CFCC stirrups in the critical shear span causing congestion in 108

reinforcement cage; b) to optimize the economy of beam by optimal usage of expensive UHPC 109

in high shear stress regions; c) to enhance the energy dissipation of CFCC prestressed decked 110

bulb T beams. To provide a better comparative assessment on the shear performance of hybrid 111

beams, its results were compared to that of beams reinforced with steel/CFCC stirrups. 112

RESEARCH SIGNIFICANCE 113

This paper aims to contribute a novel cognition to the engineering and design community 114

in understanding the newly developed material UHPC and its optimal usage in precast 115

prestressed beams for Accelerated Bridge Construction (ABC) by reducing on-site construction 116

time due to its novel design and construction concept. Present investigation describes the 117

ingenious techniques employed for the construction of innovative hybrid decked bulb T beam 118

models reinforced and prestressed with CFCC strands and compares its shear performance with 119

that of a traditional CFCC prestressed beams reinforced with either steel/CFCC stirrups. In 120

addition, finite element (FE) models developed for experimental beams will help the engineers to 121

consider a wide range of configurations for prestressed hybrid decked bulb T beams. 122

5

CONSTRUCTION DETAILS 123

Two one-half scale hybrid beams with an effective span length of 41 ft. (12.5 m) were built, 124

instrumented and tested at the Center for Innovative Material Research (CIMR), Lawrence 125

Technological University (LTU). The variable considered in this investigation includes; shear 126

span-to-depth ratio (a/d) ratios of 3, 4, 5 & 6. Table 1 outlines the test variables considered for 127

hybrid beam and the HSC beams tested by Grace et al. (2014). The typical cross-section of 128

hybrid beams at both critical and non-critical shear span is shown in the Figure 1, having a top 129

flange width of 18 in. (457 mm), an overall depth of 16 in. (406 mm), a web thickness of 3 in. 130

(76 mm) and bottom flange width of 12 in. (305 mm). Both of the beams were pre-tensioned 131

with an effective prestressing force of 100 kip (444.822 kN), using four CFCC (Tokyo Rope 132

2013), 1x7 strands having an effective diameter of 0.6 in. (15.2 mm) and cross sectional area of 133

0.18 in.2 (116 mm2). Figure 2 shows the longitudinal view of both hybrid beam and HSC beam 134

tested by Grace et al. (2014) with dimensions along their reinforcement cages. The material 135

properties for the reinforcements used in both of the above beams are summarized in Table 2, 136

while Table 3 outlines the mix design followed for HSC & UHPC per cubic yard of concrete 137

volume. The anchorage system used to prestress the CFCC strands employed an innovative 138

technique and discussed elsewhere (Grace et al. 2012). The new system significantly reduces 139

seating losses and avoids damage to the strands. The concrete mix was designed to achieve an 140

average 28-day compressive strength of 8,000 psi (55 MPa) and 24,000 psi (165 MPa) for HSC 141

and UHPC respectively. Such a superior strength of UHPC is mainly achieved through its finely 142

graded, homogenous granular material composition and inclusion of silica fume & super-143

plasticizer. Dimensionally largest granular material in the composition of UHPC is fine sand 144

ranging between 150 to 600 µm. 145

Based on the research finding on the modes of shear failure on CFCC prestressed decked 146

bulb-T HSC beams conducted by Grace et al (2013) as discussed earlier, length of the critical 147

shear span was decided with some factor safety to be 8 times the effective depth of the beam. 148

The critical shear span of hybrid beam experiencing higher shear stresses were constructed with 149

UHPC without shear stirrups, while the middle flexural span was constructed with HSC with 150

steel stirrups at 4 in. (102 mm) of spacing. The purpose of providing stirrups in middle flexural 151

span was to restrain extensive propagation of cracks. Due to absence of stirrups within critical 152

6

shear span, end and middle diaphragm reinforcement served the purpose for holding the 153

longitudinal reinforcement in position across the depth. 154

A very simple and innovative technique was developed for the formation of a monolithic 155

joint between UHPC and HSC. Figure 3 shows the various stages of constructions for the hybrid 156

beams. A trap door was built underneath the deck at the location of joint. HSC was placed first in 157

the middle flexural span, while the open trap doors restricted overflow of HSC into the critical 158

shear span by removing the excessive concrete. Once the HSC was placed, trap doors were 159

closed and UHPC was placed in the critical shear span region forming the monolithic joint to a 160

distance of 8·d from the beam-ends. As reviewed in the previous section, favorable mechanical 161

properties of UHPC are obtained when the fibers are oriented parallel to the direction of 162

principal tensile stress. Hence, the placement of UHPC in the critical shear span was performed 163

starting from the end of the beam moving inwards towards the joint. Owing to self-consolidating 164

property of UHPC, use of vibrators was restricted, which helped the fibers to align normal to the 165

direction of flow. 166

The curing of constructed beams were performed by covering the middle HSC with wet 167

burlap and ends UHPC with plastic sheets to prevent any loss of moisture which causes cracks 168

due to shrinkage of concrete. The calibrated load cells were used to evaluate prestress loss 169

continuously for seven days from the day of concrete placement. At the end of seven days as 170

concrete attained it’s enough designed compressive strength, prestressing force was released to 171

the beam by gradual warming of steel prestressed strands attached to CFCC prestressed strands 172

through coupler with an acetylene torch. The concrete compressive strength for all beams was 173

averaged around 5,800 psi (40 MPa) and 17,700 psi (122 MPa) for HSC and UHPC respectively 174

at the time of prestress release. Figure 4 shows the average compressive strength of HSC and 175

UHPC with curing age. 176

INSTRUMENTATION AND TEST SETUP 177

The beams were simply supported over a set of elastomeric neoprene bearing pads as 178

shown in Figure 5 and were subjected to a concentrated vertical shear load, applied by a MTS 179

hydraulic actuator having 220 kip (1,000 kN) capacity. Shear span considered for the beams was 180

45, 60, 75 and 91 in. (1.1, 1.5, 1.9 and 2.3 m) from the center of support equivalent to a/d ratios 181

of 3, 4, 5 and 6 respectively. Linear strain gauges were installed on the surfaces of the top and 182

bottom flanges of the beams near the loading point to measure the compressive and tensile strain 183

7

of the concrete respectively. However, due to the differences in material properties between 184

UHPC and HSC, strain gauges were also installed on the concrete surfaces on either side of the 185

joint, to measure any difference in the concrete strain development. Linear motion transducers 186

(LMT) were installed on the beams under the load to measure the vertical deformation due to 187

shear loading. Linear Variable Displacement Transducer (LVDT’s) were installed on the web 188

within the shear span in sets of three and arranged in a rectangular rosette fashion at 00, 450, and 189

900 directions to monitor and record the progress of cracks and crack width. All various sensors 190

were calibrated and connected to a fine-tuned digital data acquisition system using Mars Lab 191

computer interface. The beams were subject to several loading and unloading cycles prior to 192

ultimate failure to separate the elastic and inelastic energies required to determine the shear 193

ductility indices. 194

RESULTS AND DISCUSSION 195

The experimental investigation evaluated the effect of variation of shear span-to-depth 196

(a/d) ratio on cracking and ultimate shear resistance, crack width and the pattern of crack 197

development, maximum concrete compressive and tensile strains development, ductility indices 198

and modes of shear failure. Following section consists of four parts. First part compares the 199

experimental result of all hybrid beam, second part compares the experimental results of hybrid 200

beam with HSC beam conducted by Grace et al (2014) under similar load scenario, while the 201

third and fourth part compares the experimental results of hybrid beams with numerical and 202

analytical results respectively. 203

1. Hybrid Beam Results 204

Effect on deflection 205

Figure 6 shows the shear force versus deflection response of hybrid beams at various a/d ratios. 206

It was observed that the deflection of the hybrid beam under different loading point increases 207

with increase in shear span-to-depth (a/d) ratio. The maximum deflection observed for hybrid 208

beams at varying a/d ratios of 3, 4, 5 and 6 was 3.4 (86.4), 8.3 (210.8), 9.9 (251.5) and 7.7 209

(195.6) in (mm) respectively. However, the maximum deflection observed for hybrid beam at a/d 210

of 6 was lower than that of a/d of 4 and 5. Since, at a/d of 6, the point of loading was fairly close 211

to the concrete joint which eventually increased the rate of development of compressive strain on 212

HSC side and attained failure strain quickly without showing appreciable deflection 213

8

Effect on concrete cracking and ultimate shear resistance 214

Concrete cracking force and ultimate shear resistance for hybrid beam was found to 215

decrease with increase in a/d ratios as observed in Figure 14 Thus, it shows that the shear-216

moment interaction (M/Vd = a/d) plays a very vital role in determining the concrete cracking and 217

ultimate shear resistance of prestressed UHPC beam. The concrete cracking force observed for 218

the hybrid beams at a/d of 3, 4 5 and 6 was 34.2 (152.1), 21.9 (97.4), 16.1 (71.6) and 12.2 (54.3) 219

kip (kN) respectively, whereas the observed ultimate shear capacity for the hybrid beams at a/d 220

of 3, 4, 5 and 6 was 118.8 (528.5), 100.6 (447.5), 80.9 (359.9), 62.3 (277.1) kip (kN) 221

respectively. 222

Effect on concrete strains 223

Figure 13 shows the response of concrete compressive strain developed at various 224

location along the span in hybrid beams when loaded at varying a/d ratios. The above figure 225

shows that when the position of load moves away from the support i.e. with increase in a/d ratio, 226

the location of development of maximum compressive strain at top flange moves from the UHPC 227

near load to HSC near the concrete joint. The maximum concrete compressive strain experienced 228

at the top flange of hybrid beams was 2206 µε near the loading point at a/d of 3, while 3053, 229

3519 and 3224 µε was observed near the joint on HSC side at a/d of 4, 5 and 6 respectively. 230

Similarly, the maximum tensile strain of concrete experienced in the bottom flange of the hybrid 231

beam carries an inverse relationship with increase in a/d ratio. Maximum tensile experienced by 232

UHPC under load was 5980, 6280, 4781 and 3024 µε at a/d ratio of 3, 4 5 and 6 were 233

respectively as shown in the Figure 8. 234

Effect on crack width and crack pattern 235

The presence of UHPC in the critical shear span of hybrid beam greatly influenced the 236

patterns of shear cracks at various shear span-to-depth (a/d) ratio (loading point) as illustrated in 237

Figure 9. Flexural cracks were first seen at the bottom flange of the hybrid beam underneath the 238

loading point. Gradually, several distributed micro-cracks starts to initiate diagonally near the 239

web regions within critical shear span. The shear cracks in the web propagated diagonally 240

outwards towards the top and bottom flange with an increase in the shear force. The hybrid 241

beams were characterized by tightly spaced cracks normal to principal tensile stress direction. 242

This showed the ability of UHPC in redistributing the stresses through three dimensional (3D) 243

9

steel fiber reinforcements over multiple cracking before the fibers pullout. With further 244

increment in shear force, the steel fiber starts to pullout as the load carried by individual steel 245

fiber exceeds the ability of the UHPC matrix to grip the fiber and it eventually leads to diagonal-246

tensile failure of the beam. The displacements measured by LVDT’s were used to determine the 247

crack width by using the equation given by Shehata et al. (1999). Figure 10 shows the shear 248

force versus crack width response for hybrid beam. It was observed that crack width increases 249

with the increase in the a/d ratio. 250

Shear and flexural cracks in hybrid beam were observed to cross the diagonal monolithic 251

UHPC-HSC joint from one to adjacent concrete, confirming the satisfactory behavior of 252

monolithic concrete joint in distributing stresses. No parallel cracks or premature failure were 253

noticed along the diagonal seam of concrete joint. This shows that concrete joint in hybrid beam 254

had sufficient bond in binding together both types of concretes till the ultimate failure of the 255

beam as observed in the investigation. Further as illustrated by Figure 11, initiation of diagonal 256

cracks within shear span of hybrid beams was much delayed due to the presence of UHPC in the 257

critical shear span as compared to non-critical shear span (HSC section) in hybrid beams. UHPC 258

has higher tensile capacity and greater ability to bridge across micro-cracks through 3D steel 259

fibers. Thus, presence of UHPC within the critical shear span of hybrid beam greatly influences 260

the initiation and pattern of cracks under shear loading and eventually enhances the concrete 261

cracking under service limit state (SLS) and ultimate capacity under the ultimate limit state 262

(ULS). 263

Effect on ductility indices 264

The traditional way of calculating the ductility indices of concrete beams prestressed with 265

steel strands is not suitable for concrete beams prestressed with FRP tendons because unlike 266

steel, FRP does not have any yield plateau. Naaman and Jeong (1995) proposed the energy-267

based method to evaluate the ductility indices for the FRP prestressed beams. Due to the 268

presence of steel fibers, hybrid beams showed more efficiency in dissipating the shear forces 269

through elastic & inelastic energy across the shear crack. The ductility indices for hybrid beam 270

were within the range of 37-29%. Prior to ultimate failure, hybrid beams showed extensive 271

cracks and loud fiber pull-out signals. 272

10

Shear Mode of Failure 273

At a/d of 3 and 4, hybrid beam failed in diagonal shear while at a/d of 5 and 6, hybrid 274

beam exhibited extensive flexural cracks on HSC side (non-shear span) near the joint region 275

leading to compression flexural failure characterized by crushing of top flange. Figure 12 276

illustrates various modes of failure observed in hybrid beams at various a/d ratio. It was noticed 277

that as the load moved away from the support (higher a/d ratio), the hybrid beam experienced 278

increase in the concentration of flexural cracks on the flexural span (HSC) and decrease in the 279

concentration of diagonal shear cracks within the shear span (UHPC). Thus, hybrid beams 280

showed more of ductile flexural-shear failure. 281

2. Comparison with Grace et al. (2014) results 282

As discussed earlier, to provide a better assessment on the performance of hybrid beam, 283

the results of a hybrid beam were compared with the results obtained by Grace et al. (2014) on 284

HSC beams transversely reinforced with steel/CFCC stirrups and tested under similar load 285

scenarios. Table 4 summarizes the comparison of experimental results between hybrid beams 286

and HSC beam. 287

Effect on deflection 288

The maximum deflection observed for HSC beam at a/d of 3, 4, 5, and 6 was 1.4 (35.6), 289

2.6 (66.0), 3.5 (88.9), 4.8 (121.9) in (mm) and 1.6 (40.6), 3.0 (76.2), 4.1 (104.1), 5.5 (139.7) in. 290

(mm) for HSC beam-end with steel and CFCC stirrup respectively. However, on comparison, it 291

was observed that hybrid beams on average deflected 2.3 times the maximum deflection of HSC 292

beams with stirrups under similar a/d ratios. 293

Effect of varying a/d ratios 294

The influence of varying a/d ratio on relative flexural capacity (Mu/Mfl) of the same 295

section with different type of shear reinforcement can be understood from Figure 13, which 296

eventually determines the mode of beam failure. It was observed that the relative flexural 297

capacities (Mu/Mn) and its rate of growth with an increase in a/d ratio for hybrid beam were 298

higher than as compared from HSC beams. Or in other words, it can be explained as when the 299

position of load moves away from the support (increase in a/d ratio), ultimate moment (Mu) is 300

close or higher to nominal flexural capacity (Mn) of HSC section in hybrid beam (or Mu/Mn ≈ 1 301

or > 1), and the beam eventually fails in flexure without exceeding the shear capacity of UHPC 302

11

in the critical shear span as observed in the experimental program. Thus, presence of UHPC in 303

the critical shear span of hybrid beams helps in increasing the relative flexural capacity (Mu/Mn) 304

of HSC section in flexural span and changes the catastrophic shear failures as observed in all 305

HSC beams under similar loading scenarios into more ductile shear-flexural failure. 306

Effect on concrete cracking and ultimate shear resistance 307

After comparison of hybrid beams with HSC beam, it was observed that the cracking and 308

ultimate shear force followed the inverse relationship with a/d ratio as shown in the Figure 6 309

irrespective of type of shear reinforcement within the critical shear span i.e. either steel fibers in 310

UHPC without stirrups or steel/CFCC stirrups. The concrete cracking force for HSC beams at 311

a/d ratio of 3, 4, 5 and 6 was 27.3 (121.4), 20.1 (89.4), 15.8 (70.3), 12.4 (55.2) kip (kN) and 26.8 312

(119.2), 19.2 (85.4), 15.6 (69.4), 14.2 (63.2) kip (kN) for steel and CFCC stirrups beam-ends 313

respectively. However, it was observed that concrete cracking force for hybrid beams was on 314

average 8% higher than HSC beams. The increase in the shear cracking force for hybrid beams 315

can be related to UHPC higher tensile cracking strength as compared to the low tensile strength 316

of conventional concrete. The ultimate shear capacity for the HSC beams at a/d of 3, 4, 5 and 6 317

was 61.2 (272.2), 53.6 (238.4), 49.7 (221.1), 44.2 (196.6) kip (kN) and 58.6 (260.7), 52.2 318

(232.2), 49.1 (218.4), 46.3 (205.9) kip (kN) for steel and CFCC stirrups beam-end respectively. 319

Thus it can be noticed that the ultimate shear capacity of hybrid beams was 94, 88, 63 and 41 % 320

higher than HSC beam-end with steel stirrup and 103, 93, 65, and 34 % higher than the HSC 321

beam-end with CFCC stirrups under similar load configuration. Figure 14 illustrated the cracking 322

and ultimate shear capacity of hybrid beam in comparison with HSC beams. Due to the presence 323

of steel fibers, UHPC served as a three dimensional (3D) reinforcement and increases capability 324

of UHPC in dissipating shear stresses across cracks. This phenomenon significantly increases the 325

post-cracking tensile strength of UHPC through superior bonding between the concrete matrix 326

and distorted fibers even after the initial cracking, which eventually leads to increased shear 327

capacity for hybrid beam in comparison with HSC beam reinforced with stirrup. 328

Effect on concrete strains 329

The maximum concrete compressive strain observed at top flange near loading point for 330

HSC beam-ends with steel stirrups at a/d ratios of 3, 4 5 and 6 was 1642, 2038, 2639 and 2649 331

µε respectively, while it was 1282, 1767, 2624 and 2732 µε respectively for HSC beam-end with 332

12

CFCC stirrups. However, it was observed that the maximum compressive strain experienced by 333

the hybrid beam was on average 1.42 times than that of HSC beams. It was be due to higher 334

compressive strength properties of UHPC over HSC. Again, the maximum tensile strain 335

observed by hybrid beam was on average 13.75 times the typical average tensile strain of 350 µε 336

in HSC beams irrespective of a/d ratio. Apart from superior material characteristics of UHPC in 337

terms of compressive strength, the above experimental results also prove the effectiveness of 338

steel fibers in increasing the tensile strength of UHPC. 339

Effect on crack width and crack pattern 340

The presence of UHPC in the critical shear span of hybrid beam greatly influenced the 341

patterns of shear cracks as compared to HSC beams. Analogous to hybrid beam, HSC beam was 342

also initially marked by flexural cracks in the bottom flange underneath the loading point. But 343

unlike hybrid beam, HSC beam was characterized by an irregular diagonal cracks in the shear 344

span with sparsely spaced wider cracks. Having the same Shehata equation, the crack width 345

observed for hybrid beams was on an average of 40-50% less than those of HSC beam with 346

stirrups. However, irrespective of types of shear reinforcement type i.e. either steel fibers in 347

hybrid beams or steel/CFCC stirrups in HSC beams, crack width increases with the increase in 348

the a/d ratio. 349

Effect on ductility indices 350

Due to the presence of steel fibers, hybrid beams showed more efficiency in dissipating 351

the shear forces through elastic & inelastic energy across the shear crack as compared with HSC 352

beams. Hybrid beams on average showed 4 to 5 times more efficient in absorbing the elastic and 353

inelastic energies than that of HSC beams under similar a/d ratios as outlined by Table 5 and 354

illustrated by Figure 15. However, there was not much significant difference in the calculated 355

ductility indices of hybrid beams based on the energy method in comparison with HSC beams. 356

As discussed earlier, hybrid beam showed higher ultimate deflection than HSC beam. Thus, 357

hybrid beams showed sufficient warning prior to failure in terms of higher deflection, extensive 358

cracks and loud fiber pull-out signals prior to ultimate failure unlike HSC beam with either 359

steel/CFCC stirrups which failed with little or no warning. 360

361

13

Shear Mode of Failure 362

All HSC beam-end reinforced with steel stirrups failed in shear tension with yielding of 363

steel stirrups while all HSC beam-end reinforced with CFCC stirrups failed in shear compression 364

followed by crushing of either web or top flange concretes. The linear stress-strain relationship 365

of CFCC stirrups causes a linear increase in the strain of stirrups until the concrete begins to 366

crush. Thus, all HSC beams failed in shear irrespective of a/d ratio, however hybrid beam 367

changes its modes of failure from diagonal shear to compression flexure with increase in a/d 368

ratio. Thus, hybrid beams showed more of ductile flexural-shear failure rather than a typical 369

catastrophic shear failure as observed in all HSC beams. 370

3. Comparison with Finite Element Analysis Results 371

The finite element models were generated in an attempt to reproduce the structural 372

response of hybrid beams and to validate the observed experimental results. The Concrete 373

Damage Plasticity (CDP) model was tailored to simulate hybrid beams within a commercially 374

available Finite Element Analysis (FEA) package ABAQUS. The CDP model simulates isotropic 375

damage elasticity combined with isotropic tensile and compressive plasticity to represent the 376

inelastic behaviors of concrete materials. Table 6 shows the detailed information on material 377

properties and meshed finite elements used separately for reinforcement, HSC and UHPC in the 378

hybrid beams. Table 7 presents the comparison between the results obtained from the numerical 379

models and the experimental investigation. 380

Diagonal shear failure predicted by numerical model matches with the observed experimental 381

failure as observed in Figure 16. It was noticed that the behavior of the FEA model matched 382

fairly close to that of the tested beams behavior as shown by the Figure 17 with an accuracy level 383

exceeding above 90%. Therefore, these FEA models can be efficiently utilized in further 384

exploration of the wide ranges of configurations for prestressed hybrid decked bulb T beams by 385

changing parameters which influence shear capacities. 386

4. Comparison with Analytical Results 387

Apart from design guidelines provided by the Federal Highway Administration (FHWA), 388

United States has no unified and accepted design method for using Ultra High Performance 389

Concrete (UHPC). Thus, it is very essential to compare the experimental results with the 390

predicted values given by various empirical formulas available internationally for UHPC such as 391

14

JSCE (2006) from Japan and AFGC-SETRA (2000) from France. Both design guidelines 392

divides the shear resistance of UHPC section at Ultimate Limit State (ULS) into three 393

components; i) composite contribution of UHPC matrix and the fibers (Vc) and the shear 394

resistance provided by the average fiber tensile resistance before fiber pullout, acting along the 395

diagonal cracks (Vf), ii) the contribution from prestressing strands (Vp), iii) the contribution from 396

shear reinforcement or stirrups (Vs). The composite shear resistance of UHPC matrix and the 397

fibers according to AFGC-SETRA (2000) is given below: 398

zb'f24.0

V 0cc γ= (1) 399

While according to JSCE (2006); 400

zbfV cc 0'18.0

γ= (2) 401

Where, 402

'f c = 28 days concrete compressive strength, 403

b0 = web width, 404

z = lever arm between the centroids of the compression block and the prestressing strands at 405

ultimate moment, 406

γ = factor of safety which was considered equal to 1 for the purpose of comparison with 407

experimental results. 408

The average fiber tensile resistance before fiber pullout according to both design guidelines is 409

given as: 410

βγσ

=u

0pf

tan

zbV (3) 411

Where, 412

σp = residual tensile stress carried across the shear crack from the time of cracking until a certain 413

limiting tensile strain. 414

βutan = tangent of the compression strut angle in the shear span measured from the horizontal 415

and it has the lower bound value of 30°. 416

However, according to Graybeal (2006), the residual tensile stress values are determined 417

experimentally from the tension tests. Based on the typical values found in various literature 418

(Graybeal 2006, Gowripalan and Gilbert 2000, and AFGC-SETRA 2002), a conservative 419

15

values of 1000 psi (6.9 MPa) was used for residual tensile strength in predicting the shear 420

capacity. In this experimental program, Vp and Vs were set equals to zero as the prestressing 421

strands were straight without any draping and the absence of transverse reinforcement. It was 422

observed that shear capacity predicted by both existing codes were conservative when compared 423

with experimental results as illustrated by Table 8. Further, it should be noticed that none of the 424

above design codes take into account the variation of shear capacity with respect to a/d ratio and 425

thus it was critical to compare the experimental results with analytical results. 426

SUMMARY AND CONCLUSIONS 427

This paper brings an experimental investigation addressing the use of UHPC without shear 428

stirrups in the critical shear span of CFCC prestressed decked bulb T beams. The experimental 429

program included construction, instrumentation and testing of two hybrid beams and four control 430

beams with CFCC and steel stirrups at its beam-end under shear load setup at a/d ratios of 3, 4 5 431

and 6. The experimental results were compared and several conclusions were drawn as follows: 432

1. Irrespective of type of shear reinforcement, cracking and ultimate shear resistance were 433

found to be indirectly proportional to a/d ratios. Thus it is very important for take into 434

consideration the effect of shear span-to-depth ratio or shear-moment interaction while 435

predicting the shear capacity of a UHPC section which is currently ignored in shear design 436

guidelines. 437

2. Hybrid beams with UHPC in the critical span without stirrups exhibited remarkably superior 438

shear performance with an average increase of 8% in the concrete cracking shear force and 439

an average increase of 72% in the ultimate shear resistance as compared to similarly 440

prestressed HSC beam-ends reinforced with either steel/CFCC stirrups under similar loading 441

configuration. 442

3. Hybrid beams showed 4-5 times more efficient in absorbing the elastic and the inelastic 443

energy as compared with HSC beams. Further, hybrid beams demonstrated sufficient 444

warning prior to ultimate shear failure in terms of excessive deflection, extensive multiple 445

cracks and loud fiber pull-out signals unlike HSC beams with either steel/CFCC stirrups 446

which showed catastrophic shear failure. 447

4. Monolithic concrete joint between HSC and UHPC did not acknowledge any parallel cracks 448

or premature failure along the diagonal seam of joint. Shear and flexural cracks were noted to 449

16

cross the diagonal concrete joint from one to adjacent concrete conforming the satisfactory 450

bond behavior of monolithic concrete joint between UHPC-HSC in distributing stresses. 451

5. The pattern of crack initiation and its progress observed in hybrid beams was remarkably 452

different from those of HSC beams. Hybrid beams experienced 40-50% reduction in the 453

shear crack width. 454

6. UHPC in the critical shear span of hybrid beams was effectual in changing the mode of 455

catastrophic shear failure to more ductile shear/flexural failure with the increase in a/d ratio. 456

7. Critical shear span with shear stirrups in CFCC prestressed decked bulb T beams can be 457

effectively replaced by UHPC without any requirement of stirrups with an additional 458

advantage of increased shear capacity at both SLS and ULS. 459

8. The developed FEA models of the tested beams accurately predicted the shear cracking 460

force, the ultimate shear capacity, ultimate deflection, maximum compressive and tensile 461

strains in the concrete and finally the modes of failures. The average percentage error in 462

predicting these experimental values were within 10%. 463

9. Both Japanese (JSCE) and the French Code (AFGC) predicted approximately the same shear 464

capacity, neglecting the major influential shear-span-to-depth (a/d) factor in their equation 465

and were also found to be conservative when compared with experimental ultimate shear 466

capacity. 467

REFERENCES

1. ACI Committee 318 (2011). “Building Code Requirements for Structural Concrete and

Commentary.” American Concrete Institute, Farmington Hills, MI, pp. 503

2. AFGC-SETRA (2002). “Civil Interim Recommendation for Ultra High Performance Fiber-

Reinforced Concrete.” Association with Franćaise de Génie Civil (AFGC).

3. El-Sayed, A. K., El-Salakawy, E., and Benmokrane, B. (2007). “Mechanical and structural

characterization of new carbon FRP stirrups for concrete members.” Journal of Composites

for construction, 11(4), pp. 352–362.

4. Gowripalan, N., Gilbert, R.I. (2000). “Design guidelines for RPC prestressed concrete

beams.” School of Civil and Environmental Engineering, University of New South Wales,

Sydney, Australia, pp.XX-XX

17

5. Grace N. F., Enomoto T., Baah P., Bebawy M. (2012). “Flexural Behavior of CFRP Precast

Prestressed Decked Bulb T-Beams.” Journal of Composites for Construction, ASCE, 16 (3),

29 pp. 225-234.

6. Graybeal, B.A. (2006). “Material Property Characterization of Ultra-High Performance

Concrete.” Federal Highway Administration, U.S. Department of Transportation, FHWA-

HRT-06-103, pp.XX-XX.

7. Higgins, C., Farrow III, W.C., Potisuk, T., Miller, T.H., Yim, S.C., Holcomb, G.R., Cramer,

S.D., Covino, B. S., Bullard, S.J., Ziomek-Moroz, M., and Matthes, S.A. (2003). “Shear

Capacity Assessment of Corrosion damaged Reinforced Concrete Beams.” Department of

Civil Engineering, Oregon State University, U. S. Department of Energy, Albany Research

Center, pp.XX-XX.

8. Homeland Security of Science and Technology. (2010). “Ultra High Performance Concrete

(UHPC) Pathway to Commercialization.” pp. XX-XX

9. Imam, M., Vandewalle, L., Mortelmans, F., and Gemert, D.V. (1997). “Shear domain of

Fiber-Reinforced High-Strength Concrete Beams.” Engineering Structures, 19(9), pp. 738-

747.

10. Japan Society of Civil Engineers. (2006). “Recommendation for Design and Construction of

Ultra High Strength Fiber Reinforced Concrete Structures (Draft).” JSCE Guidelines for

Concrete, No. 9. September, pp.XX-XX.

11. Kim, S., S. Kang, J. Park, and G. Ryu. (2008). “Effect of Filling Method on Fiber Orientation

& Dispersion and Mechanical Properties of UHPC.” Proceedings, Second International

Symposium on Ultra High Performance Concrete, Kassel, Germany, pp. 185-192.

12. Mitchell, D., Marchand, J., Croteau, P., and Cook, W. (2011). “Concorde Overpass Collapse:

Structural Aspects.” Journal of Performance of Constructed Facilities., 25(6), pp. 545–553.

13. Naaman, A. E., and Jeong, S. M. (1995). ‘‘Structural ductility of concrete beams prestressed

with FRP tendons.” Non-metallic (FRP) Reinforcement for concretes structures, Taerwe, L.

ed., E. & FN Spoon, London, pp. 379–401.

14. Nabipay, P., and Svecova, D. (2012). “Shear Resistance of Concrete T-Beams Prestressed

With CFRP Cables.” CICE 2012 proceeding, International Institute for FRP Construction.

18

15. Padmarajaiah, S.K., and Ramaswamy, A. (2001). “Behavior of Fiber-Reinforced Prestressed

and Reinforced High-Strength Concrete Beams Subjected to Shear.” ACI Structural Journal,

98(5), pp. 752-761.

16. Park, S.Y., and Naaman, A.E. (1999). “Shear Behavior of Concrete Beams Prestressed With

FRP Tendons.” Precast/Prestressed Concrete Institute (PCI), Jan-Feb (7), pp. 74-85.

17. Rout, S.K. (2013). “Shear Performance of Prestressed Concrete Decked Bulb T Beams

Reinforced with CFCC Stirrups.” MS thesis, Lawrence Technological University, MI, USA.

18. Russo, G., Zingone, G., and Puleri, G. (1991). “Flexure-Shear Interaction Model for

Longitudinally Reinforced Beams.” ACI Structural Journal, 88(1), pp. 60-68.

19. Shehata, E. F. G. (1999). “Fibre Reinforced Polymer (FRP) for Shear Reinforcement in

Concrete Structures.” Ph.D. Dissertation, Department of Civil and Geological Engineering,

University of Manitoba, March, pp. 316.

20. Shehata, E., Morphy, R., and Rizkalla, S. (2000). “Fiber reinforced polymer shear

reinforcement for concrete members: Behavior and design guidelines.” Can. J. Civ. Eng., pp.

859–872.

21. Tadepalli, P., Dhonde, H., Laskar, A., Mo, Y., and Hsu, T. (2010). “Effect of Steel Fibers on

Shear Behavior of Prestressed Concrete Beams.” Earth and Space, pp. 2755-2764.

22. Taylor, C., Montoya, K., Jáuregui, D., Newtson, C., and Weldon, B. (2011). “Feasibility

Analysis of Using UHPC in Prestressed Bridge Girders.” Structures Congress, pp. 203-214.

19

Table 1. Experimental Variables for Tested Beams 468

Beam

Designation

a/d

ratio

Concrete type, 28th day Compressive

Strength, ksi (MPa)

Stirrups, Spacing,

in. (mm)

Prestressing

Force,

Kip

(kN)

Critical Shear Span Non-Critical Shear

Span

Critical

Shear

Span

Non-

Critical

Shear

Span

Concrete

Type,

ksi

(MPa)

Critical

Shear

Span,

in. (m)

Concrete

Type,

ksi

(MPa)

Non-

Critical

Shear

Span,

in. (m)

HB-100-3-0 3

UHPC,

24.2

(166.8)

8d,

118

(3)

HSC,

8.3

(57.2)

17.6d,

256

(6.5)

None,

0

(0)

Steel,

4

(101.6)

104

(463)

HB-100-4-0 4

HB-100-5-0 5

HB-100-6-0 6

SS-100-3-6 3

HSC,

8.3

(57.2)

Steel,

6

(152.4)

95

(422.6)

SS-100-4-6 4

SS-100-5-6 5

SS-100-6-6 6

SC-100-3-6 3

HSC,

8.3

(57.2)

CFCC,

6

(152.4)

SC-100-4-6 4

SC-100-5-6 5

SC-100-6-6 6

a/d ratio = Shear span (a) / effective depth (d), 469

HB-XXX-Y-Z = Hybrid beam-Prestressing force (kip)-a/d ratio-Stirrup spacing,

SS-XXX-Y-Z = HSC beam with steel stirrup-Prestressing force (kip)-a/d ratio-Stirrup spacing, 470

SC-XXX-Y-Z = HSC beam with CFCC stirrup-Prestressing force (kip)-a/d ratio-Stirrup spacing. 471

472

473

474

475

476

477

478

479

480

20

Table 2. Material Properties for Reinforcements 481

Material Property units CFCC Longitudinal Steel Stirrups

Designation - CFRP 1 X 7 # 3

Diameter in. (mm) 0.60 (15.2) 0.375 (9.5)

Effective Cross Sectional Area in.2 (mm2) 0.179 (115.5) 0.11 (71)

Linear Density lb/ft (g/m) 0.15 (223) 0.38 (565)

Guaranteed Breaking Load kip (kN) 60.7 (270) --------

Yield Strength ksi (MPa) --------- 60 (414)

Tensile Strength ksi (MPa) 373 (2,930) 90 (620)

Elastic Modulus ksi (MPa) 22,480 (155,000) 29,000 (200,000)

Elongation % 1.7 4.9

482

Table 3. Mix Design for Concretes per Cubic Yard 483

High Strength Concrete (HSC)

Coarse

Aggrega

te, lb

(kg)

Fine

Aggrega

te, lb

(kg)

Cement

(type 1),

lb (kg)

Slag

Cement

(Grade

100), lb

(kg)

Water

reducing

admixture,

gal, (m3)

High

Range

Water

reducer,

gal, (m3)

Water

gal,

(m3)

Water

Cement

ratio

(%)

Slump,

in.

(mm)

1,710

(775.6)

1,290

(585.1)

534

(242.2)

288

(130.6) 0.9 (2) 4.5 (10)

0.12

(31.8) 0.37

8

(203.2)

Ultra High Performance Concrete (UHPC)

Premix, lb (kg) Water, lb (kg) Premia 150,

lb (kg)

Steel fiber (2%

by Volume)

Flow table, in.

(mm)

3700 (2195) 219.1 (130) 50.6 (30) 262.9 (156) 9 (228.6)

484

485

486

21

Table 4. Summary of Experimental Results for Tested Beams 487

Beam

Designation

Ultimate

Shear Force,

kip (kN)

Cracking

Shear Force,

kip (kN)

Ultimate

Deflection

Under Load, in.

(mm)

Maximum

Concrete Strain

at Top Flange,

µε

Maximum

Concrete Strain

at Bottom

Flange, µε

Modes

of

Failure

HB-100-3-0 118.8 (528.5) 34.2 (152.1) 3.4 (85.9) -2,206 5980 DS

HB-100-4-0 106.6 (447.5) 21.9 (97.4) 8.3 (210.3) -3,053 6280 DS

HB-100-5-0 80.9 (359.9) 16.1 (71.6) 9.9 (250.2) -3,519 4781 CF

HB-100-6-0 62.2 (276.7) 12.2 (54.3) 7.7 (195.58) -3,244 3024 CF

SS-100-3-6 61.2 (272.2) 27.3 (121.5) 1.4 (36.0) -1,642 416 ST

SS-100-4-6 53.6 (238.2) 20.0 (89.0) 2.6 (66.0) -2,038 280 ST

SS-100-5-6 49.7 (220.8) 15.8 (70.3) 3.5 (90.0) -2,639 416 ST

SS-100-6-6 44.2 (196.6) 12.4 (55.2) 4.8 (122.0) -2,649 312 ST

SC-100-3-6 58.6 (260.7) 26.8 (119.2) 1.6 (41.0) -1,282 412 SC-W

SC-100-4-6 52.2 (232.0) 19.2 (85.4) 3.0 (76.0) -1,767 407 SC-W

SC-100-5-6 49.1 (218.0) 15.6 (69.4) 4.1 (104.0) -2,624 334 SC-T

SC-100-6-6 46.3 (206.1) 14.2 (63.2) 5.5 (139.0) -2,732 390 SC-T

DS - Diagonal Shear Failure, CF – Compression Flexural Failure, SC-W – Shear Compression Failure due to 488

Crushing of Web, SC-T – Shear Compression Failure due to Crushing of Top Flange, ST – Shear Tension Failure 489

due to yielding of Stirrup 490

22

Table 5. Ductility Indices 491

Beam

Designation

Inelastic Energy Absorption (Ei) Elastic Energy Absorption (Ee) Ductility

Indices (%)

= ��

������100

kip-in.

(kN-m)

× Less than

Hybrid Beam

kip-in.

(kN-m)

× Less than

Hybrid Beam

HB-100-3-0 104.3 (11.8) - 180.7 (20.4) - 36.6

SS-100-3-6 21.5 (2.4) 4.85 41.7 (4.7) 4.34 34.1

SC-100-3-6 25.6 (2.9) 4.07 56.7 (6.4) 3.19 34.1

HB-100-4-0 168.7 (19.1) - 413.6 (46.7) - 29.0

SS-100-4-6 23.6 (2.7) 7.15 65.3 (7.4) 6.33 26.5

SC-100-4-6 53.5 (6.0) 3.15 49.6 (5.6) 8.34 51.9

HB-100-5-0 159.9 (18.1) - 341.8 (38.6) - 31.9

SS-100-5-6 43.9 (5.0) 3.64 79.7 (9.0) 4.29 35.5

SC-100-5-6 38.6 (4.4) 4.14 126.1 (14.3) 2.71 23.4

HB-100-6-0 90.9 (10.3) - 244.2 (27.6) - 27.2

SS-100-6-6 48.7 (5.5) 5.01 82.4 (9.3) 1.10 37.2

SC-100-6-6 65.0 (7.3) 3.76 108.1 (12.2) 0.84 37.6

23

Table 6: Material Parameters and Elements used in the FEA Model (Abaqus) 492

UHPC HSC CFCC STRANDS STEEL

STIRRUP

Density, lb/ft3 (kg/m3)

160 (2563) 150 (2397) 76.6 (1228) 497.7 (7972)

Concrete Elasticity

Young Modulus, ksi (GPa)

8000 (55.2) 4910 (33.9) 21170 (146.0 29000 (200.0)

Poisson’s Ratio

0.18 0.2 0.3 0.3

Concrete Compression hardening Plasticity

Compressive

Stress, ksi

(MPa)

Plastic

Strain (-)

Compressive

Stress, ksi

(MPa)

Plastic

Strain (-)

Yield

Stress,

ksi

(MPa)

Plastic

Strains

(+)

14.0 (96.5) 0.000000 5.0 (34.5) 0.000000 372.748

(2570) 0.00

16.0 (110.3) 0.0000284 8.3 (57.2) 0.001167 0.0 (0.0) 0.018

20.0 (137.9) 0.0000720 0.0 (0.0) 0.003100

24.0 (165.5) 0.0001410

27.5 (189.6) 0.0004140

Concrete Tension Stiffening

Compressive

Stress, ksi

(MPa)

Plastic

Strain (+)

Compressive

Stress, ksi

(MPa)

Plastic

Strain (+)

2.3 (15.9) 0.00000 0.65 (4.5) 0.0000 Expansion

2.3 (15.9) 0.00836 0.325 (2.2) 0.0015 0.000001

0.0 (0.0) 0.009 0.0 (0.0) 0.005

Parameters of CDP Models

Dilation

Angle 0° 56°

Eccentricity 0.1 0.1

fb0/fc0 1.16 1.16

k 0.67 0.67

Viscosity

Parameter 0 0

Type of Elements

Three Dimensional Eight

Node Linear Brick

Elements (C3D8R)

Three Dimensional Eight

Node Linear Brick

Elements (C3D8R)

Three Dimensional

Two Node Linear

Truss Elements

(T3D2)

Three

Dimensional

Two Node

Linear Truss

Elements

(T3D2)

493

24

Table 7. Comparison of Experimental and Numerical Results 494

Beam

Designation

Ultimate Shear

Force, kip (kN)

Cracking Shear

Force, kip (kN)

Ultimate Deflection

Under Load, in.

(mm)

Maximum Concrete

Strain at Top Flange,

µε Modes of

Failure

Exp. Num. Exp. Num. Exp. Num. Exp. Num.

HB-100-3-0 118.8

(529)

118.2

(526)

34.2

(152)

38.7

(172)

3.4

(86)

3.7

(93.9) -2,206 -2,350 DS

HB-100-4-0 106.6

(474)

99.8

(443)

21.5

(96)

22.4

(100)

8.3

(210) 7.9 (201) -3,053 -2,933 DS

HB-100-5-0 80.9

(360)

79.5

(354)

16.1

(72)

18.1

(81)

9.9

(251) 9.2 (234) -3,519 -3,120 CF

HB-100-6-0 62.2

(277)

63.6

(283)

12.2

(54)

12.4

(55)

7.7

(196) 7.9 (201) -3,244 -3,015 CF

Exp. – Experimental, Num. – Numerical, DS - Diagonal Shear Failure, CF – Compression Flexural Failure, SC-W – 495

Shear Compression Failure due to Crushing of Web, SC-T – Shear Compression Failure due to Crushing of Top 496

Flange, ST – Shear Tension Failure due to yielding of Stirrup 497

498

499

Table 8. Ultimate Shear Capacity 500

Beam

Designation

Experimental Ultimate

Shear Capacity, kip (kN)

Ultimate Shear

Capacity as per JSCE,

kip (kN)

Ultimate Shear

Capacity as per AFGC,

kip (kN)

HB-100-3-0 118.8 (529)

81.0 (360.4) 83.5 (371.6) HB-100-4-0 106.6 (474)

HB-100-5-0 80.9 (360)

HB-100-6-0 62.2 (277)

501

25

502

503

504

505

506

507

508

509

510

511

512

513

514

(a) Typical Section for HSC Beam and

Hybrid Beam Section in Middle Flexural

Span With Steel Stirrups

(b) Hybrid Beam Section in Critical

Shear Span Without Stirrups

Figure 1. General Cross Section Detailing for tested Beams

26

515

516

517

518

519

520

521

Figure 2. Detailed Dimension for Beams along with Reinforcement Cage

Critical Shear Span

(UHPC) (8 de)

Critical Shear Span

(UHPC) (8 de) Middle Flexural Span

(c) Reinforcement Cage for Hybrid Beam

(d) Hybrid Beam

(a) Reinforcement Cage for HSC Beam Tested by Grace et al. (2014)

(b) HSC Beam Tested by Grace et al. (2014)

27

522

523

524

525

526

A) Hybrid beam reinforcement cage showing shear zone

(without stirrup) and flexural zone (with Stirrup) B) Reinforcement cage sitting on Deck with trap

door setup to facilitate concrete joint

C) Pouring of UHPC starting from beam ends

after HSC poured at mid span D) Hybrid beam concrete joint after pouring

Figure 3. Various Steps in Construction of Hybrid Beams

28

527

528

529

530

531

532

533

534

535

536

0 10 20 30 40 50 60 70 80

0

34

69

103

138

172

207

0

5

10

15

20

25

30

0 10 20 30 40 50 60 70 80

Time (Days)

Com

pre

ssiv

e S

tren

gth

(M

Pa)

Com

pre

ssiv

e S

tren

gth

(k

si)

Time (Days)

UHPC Compressive Strength

HSC Compressive Strength

HSC Cylinder (6" x 12") UHPC Cylinder (3" x 12")

Figure 4. Average Concretes Compressive Strength with Time

29

537

538

539

540

541

542

543

544

545

546

547

548

LVDT’s (For measuring crack response)

LMT

(For measuring

deflection under load)

Elastomeric Neoprene

Bearing Pad

Figure 5. Shear Test Setup for Hybrid Beams

30

549

550

551

552

553

554

555

556

557

558

0 25.4 50.8 76.2 101.6 127 152.4 177.8 203.2 228.6 254

0

89

178

267

356

445

534

623

0

20

40

60

80

100

120

140

0.0 2.0 4.0 6.0 8.0 10.0

DEFLECTION (MM)

SH

EA

R F

OR

CE

(K

N)

SH

EA

R F

OR

CE

(K

IP)

DEFLECTION (IN.)

HB-100-3-0

HB-100-4-0

HB-100-5-0

HB-100-6-0

Figure 6. Shear Force - Deflection Response for Hybrid Beams

31

559

560

561

562

563

564

565

566

567

568

-4000-3500-3000-2500-2000-1500-1000-5000

0

89

178

267

356

445

534

623

-4000-3500-3000-2500-2000-1500-1000-5000

0

20

40

60

80

100

120

140

TOP FLANGE CONCRETE STRAIN (MIRCO-STRAIN)

SH

EA

R F

OR

CE

(K

N)

TOP FLANGE CONCRETE STRAIN (MICRO-STRAIN)

SH

EA

R F

OR

CE

(K

IP)

HB-100-3-0(UHPC at Load) HB-100-3-0(UHPC at Joint) HB-100-3-0(HSC at Joint)HB-100-4-0 (UHPC at load) HB-100-4-0(UHPC at Joint) HB-100-4-0(HSC at Joint)HB-100-5-0(UHPC at Load) HB-100-5-0(UHPC at Joint) HB-100-5-0(HSC at Joint)HB-100-6-0(UHPC at Load) HB-100-6-0(UHPC at Joint) HB-100-6-0(HSC at Joint)

Figure 7. Shear Force - Compressive Strains at Top Flange at Various Location for

Hybrid Beams

32

569

570

571

572

573

574

575

576

577

0 1000 2000 3000 4000 5000 6000 7000

0

89

178

267

356

445

534

623

0 1000 2000 3000 4000 5000 6000 7000

0

20

40

60

80

100

120

140

BOTTOM FLANGE CONCRETE STRAIN (MICRO-STRAIN)

SH

EA

R F

OR

CE

(K

N)

BOTTOM FLANGE CONCRETE STRAIN

SH

EA

R F

OR

CE

(K

IP)

HB-100-3-0

HB-100-4-0

HB-100-5-0

HB-100-6-0

Figure 8. Shear Force - Bottom Flange Concrete Strain Response for Hybrid Beams

33

578

(B) Crack Pattern at a/d of 4

(C) Crack Pattern at a/d of 5

Figure 9. Crack Patterns Observed in Hybrid Beams at Various loading point (a/d)

(A)Crack Pattern at a/d of 3

(D) Crack Pattern at a/d of 6

34

579

580

581

582

583

584

585

586

587

588

0.000 0.051 0.102 0.153 0.204 0.255 0.306 0.357 0.408

0

89

178

267

356

445

534

623

0

20

40

60

80

100

120

140

0.000 0.002 0.004 0.006 0.008 0.010

CRACK WIDTH (MM)

SH

EA

R F

OR

CE

(K

N)

SH

EA

R F

OR

CE

(K

IP)

CRACK WIDTH (IN.)

HB-100-3-0(UHPC)

HB-100-4-0(UHPC)

HB-100-5-0(UHPC)

HB-100-6-0(UHPC)

Figure 10. Shear Force - Crack Width Response for Hybrid Beams

35

589

590

591

592

593

594

595

596

597

598

0.000 0.051 0.102 0.153 0.204 0.255 0.306 0.357 0.408

0

89

178

267

356

445

534

623

0

20

40

60

80

100

120

0.000 0.005 0.010 0.015 0.020

CRACK WIDTH (MM)

SH

EA

R F

OR

CE

(K

N)

SH

EA

R F

OR

CE

(K

IP)

CRACK WIDTH (IN.)

HB-100-4-0(UHPC)

HB-100-4-0(HSC)

HB-100-6-0(UHPC)

HB-100-6-0(HSC)

Figure 11. Shear Force - Crack Width Response for Hybrid Beam at a/d of 4

36

599

Figure 5(a) Diagonal Shear Failure of HB-100-3-0

Figure 5(b) Diagonal Shear Failure of HB-100-4-0

Figure 5(c) Compression Flexural Failure of HB-100-5-0

Figure 5(d) Compression Flexural Failure of HB-100-6-0

Figure 12 Modes of Failure of Hybrid Beams (HB) at various a/d ratios

37

600

601

602

603

604

605

606

607

608

609

610

2 3 4 5 6 7

0.4

0.6

0.8

1.0

1.2

1.4

0.4

0.6

0.8

1.0

1.2

1.4

2 3 4 5 6 7

a/d

Mu

/Mn

Mu

/Mn

a/d

Hybrid Beam (HSC Section at Joint) Hybrid Beam (UHPC Section under Load)

Steel Stirrup Beam CFCC Stirrup Beam

FLEXURE FAILURE

SHEAR FAILURE

Figure 13. Variation of Mu/Mn of various tested Beams with a/d

38

611

612

613

614

615

616

617

618

619

620

621

34.20

27.20

22.80 21.9019.70

16.30 16.10 14.80 15.5012.20 12.40

14.20

118.80

61.2058.60

100.60

53.60 52.20

80.90

49.70 49.10

62.20

44.2046.30

0

89

178

267

356

445

534

623

0

20

40

60

80

100

120

140

HB-100-3-0 SS-100-3-6 SC-100-3-6 HB-100-4-0 SS-100-4-6 SC-100-4-6 HB-100-5-0 SS-100-5-6 SC-100-5-6 HB-100-6-0 SS-100-6-6 SC-100-6-6

SH

EA

R F

OR

CE

(K

N)

SH

EA

R F

OR

CE

(K

IP)

BEAMS

Cracking Shear Force

Ultimate Shear Force

Figure 14 Cracking Force and Ultimate Shear capacity for all Tested Beams

39

6

22

6

23

6

24

6

25

6

26

6

27

6

28

6

29

6

30

6

31

6

32

HB-100-3-0

HB-100-3-0

HB-100-3-0

SS-100-3-6

SS-100-3-6

SS-100-3-6

SC-100-3-6

SC-100-3-6

SC-100-3-6

HB-100-4-0

HB-100-4-0

HB-100-4-0

SS-100-4-6

SS-100-4-6

SS-100-4-6

SC-100-4-6

SC-100-4-6

SC-100-4-6

HB-100-5-0

HB-100-5-0

HB-100-5-0

SS-100-5-6

SS-100-5-6

SS-100-5-6

SC-100-5-6

SC-100-5-6

SC-100-5-6

HB-100-6-0

HB-100-6-0

27.15

SS-100-6-6

SS-100-6-6

37.55

SC-100-6-6

SC-100-6-6

37.16

0 6 11

17

23

28

34

40

45

51

57

0

10

0

20

0

30

0

40

0

50

0

Ine

lastic E

ne

rgy

Ab

sorb

tion

(Ei)

Ela

stic En

erg

y

Ab

sorp

tion

(Ee

)

Du

ctility R

atio

(%) =

[Ei/(E

i+E

e)]

Energy Absorption (kN-m)

Energy Absoption (kip-in)

Fig

ure 1

5. E

nerg

y A

bso

rptio

n o

f vario

us tested

Beam

s

40

633

634

635

636

637

638

639

640

641

642

643

(B) Observed FEA Diagonal Shear Failure of HB-100-4-0

(A) Observed Experimental Diagonal Shear Failure of HB-100-4-0

Figure 16. Comparison Between Experimental and FEA Model Failure

41

644

645

646

647

648

649

650

651

0 25.4 50.8 76.2 101.6 127 152.4 177.8 203.2 228.6 254

0

89

178

267

356

445

534

623

0

20

40

60

80

100

120

140

0.0 2.0 4.0 6.0 8.0 10.0

DEFLECTION (MM)

SH

EA

R F

OR

CE

(K

N)

SH

EA

R F

OR

CE

(K

IP)

DEFLECTION (IN.)

HB-100-3-0(Exp) HB-100-3-0(Num)

HB-100-4-0(Exp) HB-100-4-0(Num)

HB-100-5-0(Exp) HB-100-5-0(Num)

HB-100-6-0(Exp) HB-100-6-0(Num)

Figure 17. Comparison between FEA and Experimental Force - Deflection Response