Girder/Deck Connection forRapid Removal of Bridge Decks

Maher K. Tadros, Ph.D., P.E.Cheryl Prewett Professor

Civil Engineering DepartmentUniversity of Nebraska-Lincoln

Omaha, Nebraska

Sameh S. Badie, Ph.D., P.E.Assistant ProfessorCivil and Environmental EngineeringDepartmentThe George Washington UniversityWashington, D.C.

Mounir R. Kamel, Ph.D., P.E.Senior Bridge Engineer

Dar El Handasah Consulting EngineersCairo, Egypt

This paper presents a debonded shear key systemfor prestressed concrete bridge girders withcomposite cast-in-place decks. The system utilizesthe mechanical anchorage of concrete shear keyscreated on the top flange of a concrete girder,combined with shear reinforcement crossing theinterface. The system has the advantage offacilitating future deck removal, while protectingthe top flange of the girder from damage, which isparticularly significant for bridges in cold climateswhere deck concrete is subjected to deteriorationdue to freeze-thaw cycles and deicing chemicals.Theoretical background, experimental results,production issues, design criteria and constructionspecifications of this system are discussed. Detailsof the first bridge project using this system inNebraska are also given. Lastly, a numericaldesign example is provided to show the requiredcalculations for the proposed shear key system incomposite girder bridges.

Rapid replacement of bridge decks is becoming increasingly important in high traffic areas and in regions of North America where deicing chemicals are

applied to deck surfaces. Traffic delays during rehabilitation of deteriorated decks cause public annoyance, inconvenience and economic hardship.

Demolition of a bridge deck that is compositely connected with precast concrete I-girders is a major concern indeck replacement projects. This operation is time consuming because concrete has to be removed without significantdamage to the precast girders or the composite-action con-

\ Ii

58 PCI JOURNAL

0.35

Fig. 1. Horizontal shear strength code equations versus test results.

nectors. This is especially relevantwhen using modem precast concrete I-girders with thin top flanges, such asthe NU I-girder, the Florida bulb tee,the New England bulb tee, and theWashington super girder.

Current practice for precast concrete1-girders made composite with theconcrete deck slab involves roughening the top flange surface and extending reinforcing bars, usually calledshear connectors, from the precastgirder into the concrete deck. Theseshear connectors provide compositeaction or horizontal interface shear resistance. They are normally the vertical shear stirrups extended above theflange and bent into an L-shape or inverted U-shape.

Although the roughened. bonded,reinforced interface detail has beenused for a long time, it has the following drawbacks:

1. Deck removal is difficult andtime consuming.

2. The top flange is at risk of beingdamaged during deck removal.

3. The shear connectors are susceptible to corrosion.

In Nebraska,1welded wire reinforcement (WWR) is used as the nonprestressed auxiliary reinforcement inprestressed concrete NU I-girders.

Girder reinforcement is not required tobe epoxy coated while deck reinforcement is required to be epoxy coated.Although some states require epoxycoating to the portion of the girder reinforcement extending into the deck,such a measure has not been taken inNebraska, as it is believed that, unlikethe deck reinforcement, corrosion cellsare unlikely to develop in such discontinuous reinforcement. In addition,epoxy coating of WWR is relativelyexpensive.

In an attempt to prevent possible future damage of the thin top flange ofthe NU I-girders during deck removal,the Nebraska Department of Roads1requires a smooth sealed surface of theoutside 8 in. (200 mm) along bothedges of the top flange. These 8 in.(200 mm) strips are trowel finishedand then sprayed with a debondingagent. Although debonding the edgesof the top flange provides partial protection, it weakens the interface shearresistance.

Project l24l2 titled “Rapid Replacement of Bridge Decks,” sponsored by the National CooperativeHighway Research Program(NCHRP), was recently completedby the University of Nebraska. Theproject identified possible methods

of maintaining composite actionwhile facilitating deck removal andreplacement.

The philosophy used in developingthese methods was that bridges shouldbe built similar to parts of a model carthat are snapped or bolted, rather thanglued or welded together. As a result,the concept of a debonded shear keyinterface system was developed forprecast concrete girders and verifiedwith laboratory experiments. The concept was further developed into a design methodology and used on abridge project for the Nebraska Department of Roads.

INTERFACE SHEARSTRENGTH

This section provides a literature review of research related to interfaceshear strength and current code methods for determining the shear strength.

literature Review

Much research3-9has been conducted during the last three decadesregarding the design for horizontalshear in composite members. In studies conducted by Hanson,3 and Kaar etal.,4 the shear strength at the interface

0.3 .

0.25

0.2

.. A

0.15

AA

Holbeck et al. (1969)

x

x• •Ax •X• AI• •

x

x• 1. <

Mattock (1976)

A

0

00.1

.

Mattock et al. (1976)

+

Wairaven et al.(1988)

.

x

A

x

.

0

—a--— Loov &Patniak (1994)

0.05

Walraven et al. (1988)

Wafraven eta!. (1988)

0

Loov & Patniak (1994)

+ Kamel(t996)

0 0.05 0.! 0.15 0.2 0.25

———AC!3!8(2002)

AASHTO Standard Specifications.

Horizontal shear reinforcement index = (A1,/b, a) (f9/f)

AASHTO LRFD Specifications

May-June 2002 59

between the girder and deck was assumed to vary directly with the amountof reinforcement crossing the interface.Parabolic equations for shear strengthwere introduced later by other researchers, as reported by Kamel.9

Several researchers contributed tothe introduction in 1970 of the theoryof shear friction in the ACI BuildingCode (ACI 318). They included testresults obtained by Kriz and Raths,5and Hofbek et al.6 In 1978, Shaikh’°introduced a shear friction design procedure that became the basis for design in the PCI Design Handbook.11

Recent research conducted by Loovand Patnaik’2 correlated the effect ofconcrete strength and clamping stresscaused by the shear connectors on thestrength of composite connections.They suggested the following equationfor determining the horizontal shearstrength:

V =nh

Ii Ahf+ ‘ f’ 0.25ff’ (psi)bs

(1)

where= area of horizontal shear rein

forcement crossing interface(sq in.)

= width of cross section at contact surface being investigatedfor horizontal shear (interfacewidth) (in.)

f, = specified minimum yieldstrength of reinforcing bars(psi)

f’ = specified compressive strengthof deck concrete at 28 days

s = spacing of horizontal shear reinforcement (in.)

Vh = nominal horizontal shear stress(psi)

Note that, for clarity of presentation, equations are given in U.S. customary units. Corresponding equations in SI units are available in thecited references.

Fig. 1 shows a comparison betweenthe test results available in publishedliterature6-9’2and various predictionmethods. It can be seen that all test results form a band that can be approxi

mated by a series of straight lines orparabolic equation. Also, Fig. 1 showsthat Loov’s equation represents an accurate lower bound of the test results.

Current Code Methods

The ACT 318-02 Code,’3 AASHTOStandard Specifications,14 andAASHTO LRFD Specifications’5uselinear relationships to determine thenominal shear strength. Although theseequations are used to design for horizontal shear, the ACT 318-02 Code13and AASHTO Standard Specifications’4 give resistance in terms of vertical shear force, while the AASHTOLRFD Specifications’5give strength interms of horizontal shear force.

This difference among equations inshear strength prediction may cause confusion among designers. Therefore, theauthors have found that it is convenientto reproduce these equations in terms ofthe nominal horizontal shear stress:

4? Vn (2)

wherec = cohesion stress

= coefficient of frictionFor concrete placed against clean

hardened concrete with the surface intentionally roughened to an amplitudeof 0.25 in.,c= lOOpsi and u= 1.0.

While the ACT 318-02 Code’3 andthe AASHTO Standard Specifications’4 appear to be entirely based onexperimental curved fitting, theAASHTO LRFD Specifications’5use alinear relationship based on shear friction theory. All prediction methodsrepresent conservative lower boundapproximations of test results and aregenerally more conservative than

(3) Loov’s’2equation.Based on the comparison given in

Fig. 1, the research team decided toadopt the LRFD representation of theshear strength for the following reasons:

1. It correlates better with test results and provides a more uniform factor of safety than the ACT 318-02’sCode and the AASHTO StandardSpecifications.’4

AASHTO Standard Specifications13

(a) If v,,, 80 psi, with

intentionally roughened

surface, no shear

reinforcement is required.

(b) If 80 < v,,, 350 psi,

with intentionally roughened

surface, use the minimum (4)reinforcement,

5ObsA,,h = (sq in.)

f(c) If Vflh > 350 psi,

/AhfV,,h = 350+0.41 ——-—50I (psi)

bs j

AASHTO LFRD Specifications14

1AVhfYlVflhC + iil—I (psi)

b,,s )Provided that V,,h O.2f’ 800 psi (5)

and A,,,, (sq in.)

(psi)

whereV,h = factored horizontal shear stress

(psi)

4? = strength reduction factorFurther discussion regarding the cal

culation of Vuh is given later in thispaper.

ACT 318-02 Code’3

(a) If v,,,, 80 psi, with

intentionally roughened

surface, no shear

reinforcement is required.

(b) If 80 < v,,,, 290 psi,

with intentionally roughened

surface, use the minimum

reinforcement,

5ObsA,,,,

=(sq in.)

Jy

(c) If 290 < Vflh 500 psi,

Vflh = 260 + 0.6(-) (psi)b,,s

(d) If V,,, > 500 psi,

= i-si —i (psi)b,,s j

60 PCI JOURNAL

2. It is simpler to use than Loov’sparabolic equation.

3. It correlated very well with theresults of tests specific to the proposeddebonded shear key detail.9

PROPOSED DEBONDEDSHEAR KEY INTERFACE

DETAILThe proposed debonded girder/deck

interface consists of:1. Concrete shear keys formed on

the top flange surface of the girderthat provide mechanical interlock between the top flange surface and thedeck concrete.

2. Sealant applied to the girder topflange surface to break the bond between that surface and the deck concrete.

3. U-shaped shear connector barsembedded in the girder web and extended into the concrete deck slab.These shear connectors are made separate from the girder vertical shear reinforcement and are epoxy coated toprotect them against corrosion.

Fig. 2 shows a three-dimensionalview of an I-girder with the proposedconnection detail. It has been found9that steel forms are the most effectivemethod in forming the shear key. Fig.3 shows the forming details for theNU I-girder used in the full-scale testgirder in the NCHRP project.2

The steel forms are attached to thegirder side forms using the yokes thathold the side forms together. Concreteis poured through the gaps betweenthe shear key blockouts. The shearkeys are projected above the standardgirder top flange surface. This conceptwas later modified for the demonstration bridge to have the shear keys recessed below the top flange surface, asshown in Fig. 2.

The reason for proposing recessedshear keys is to make the system suitable for all types of concrete decks, including precast stay-in-place sub-panels, full-depth precast panels such asthe NUDECK system,16 and cast-in-place deck systems. 1718

Required Shear Key Dimensions

Fig. 4 shows the shear key mechanism used in the analytical model. The

angle of the shear key side surface allows the two interfaces to slide acrosseach other and, hence, engage the connecting bars as indicated by the shearfriction theory. While sliding occurs,the two surfaces separate from eachother, causing tensile and shearstresses in the steel connector.

The system of forces acting on thismodel are: (1) the applied force F, (2)the tensile force in the connector P,(3) the shear force in the connectorP, (4) the bearing force on the side ofthe shear key R, and (5) the frictionforce on the side of the shear key 5R,where ô is the “local” coefficient offriction between the two surfaces.

The two surfaces are assumed tohave a coefficient of friction equal tothat used for formed “smooth” concrete-to-concrete interface. To simplify the analysis, bending stresses inthe connector are neglected, since themoment arm is very small. Also, thedebonded shear key system is designed to be stronger than the connection reinforcement to ensure acceptable ductility in the system.

To determine shear key dimensions,

T20.87

b48.2’ (1225 mm)

tsk

%= 6.9 (175 mm)

NU Girder

Fig. 2. I-girder with proposed shear key system.

May-June 2002 61

VUh

Fig. 4. Shear key mechanism.

two modes of failure are considered:1. Bearing on the side of the shear

keys:

where equals 0.7 for bearing designand f’ is the strength of the beam orthe slab, whichever is smaller.

2. Shearing of the base plane of theshear key:

Using the shear friction theory:

where equals 0.9 for shear design, inAASHTO bridge design specifications.

In Eqs. (6) and (7):Ask = area of shear key at base (sq

in.) = bSkw5kb5k = width of shear key at base

S5k = shear key spacing (in.)tsk = depth of shear key (in.)

Vflh = nominal horizontal shear capacity (ibs)

VUh = factored horizontal shear force(ibs)

6bFor concrete placed monolithically,= 1.4, and c = 150 psi, as given by

Section 5.8.4.2 of the AASHTOLRFD Specifications.15More discussion regarding the calculation of Vuh isgiven later in this paper.

The requirements of Eqs (6) and (7)were used to develop the standardshear key dimensions for the NU I-girders.

Design of Horizontal Shear(7b) Reinforcement

The test program2’9”9has demonstrated that beams designed with theproposed debonded shear key interfacesystem performed comparably withsimilar beams designed with the conventional roughened system, both inflexure and shear. Also, the test resultsshowed that the shear friction theorycould be adequately applied to designthe shear connectors, using a coefficient of friction = 1.0, as follows:

vu11(bvSsk) b(CA5i+ 1Avffy)

VLLh(bVSsk) 0.9(0 + 1 .0Af,,)

A= vh(bVSSk)

vf0.9f

Note that a cohesion stress c is setequal to zero in this application because at the interface no bond is assumed to exist between the top flangeof the girder and the concrete deck.Additional details of the theoreticalanalysis of the debonded shear keysystem and the associated test programmay be found in Kamel9 and Kakish.’9

FACTORED HORIZONTALSHEAR STRESSES FOR

COMPOSITE SECTIONSThis section discusses composite

loads and horizontal shear stress.

Composite loads

An important issue that should beconsidered in the design for horizontalshear is identification of the loads thatshould be considered in calculating thefactored horizontal stress uh at the interface. Neither the AASHTO Standard Specifications14nor the LRFDSpecifications’5give explicit guidancein this regard.

While most designers would use thetotal factored load to calculate Vuh, astrong case can be made for excludingthe self-weight of the girder and theweight of the slab since both are introduced prior to composite action takingeffect. This approach can be explainedas follows:

Fig. 5a shows a beam subjected toits own weight only. When a toppingslab is placed over the beam (see Fig.5b), the beam stresses are increaseddue to the weight of the wet concrete.However, even after the slab concreteis hardened, the stresses in the slab aretheoretically equal to zero. Therefore,no horizontal shear stresses exist at theinterface due to either the beam or theslab weight.

After the slab hardens, the beam andthe slab form a composite member,and any loads applied thereafter willbe resisted by the composite memberproducing interface horizontal shear

Unbondedsurface

y(8)

P

R

Vuh

Vh(bS5k) (O.SSfç’)(t3k)(b5k— tk)

(6a) Wsk = length of shear key at base(in.)

Va,, (7a)

Vh(bS5k) (cA51,+ M&jf)

(in.)

62 PCI JOURNAL

stresses, as shown in Fig. 5c. Theseloads are called composite loads,which consist of superimposed deadloads (barriers and wearing surface)and live loads (truck, lane, and pedestrian loads).

This design approach has been usedby some design engineers and stateagencies such as the Illinois Department of Transportation2°for a long timewithout any reported problems. If allloads are considered in the calculationof v,2, the correct loading may be overestimated by as much as 300 percent.

Horizontal Shear Stress

Calculation of the factored horizontal shear stress at the interface of acomposite member is not simple. Itshould account for the fact that concrete does not behave as a linear elastic uncracked material near the ultimate limit state. If it did, the shearstress would be calculated by the classic elasticity formula:

wherevu

1’Q=

I b1,

= factored vertical shear force ata section

I = moment of inertia of the composite section

Q = first moment of the area abovethe interface

Note that Eq. (9) does not accountfor cracking or stress nonlinearity atfactored load levels. Also, calculations of composite section propertiesby transforming the deck concrete togirder concrete through the use of themodular ratio of the two types of concrete may not accurately reflect thedifference in strength or ultimatestrain of the two interfacing concretetypes.

Loov and Patnaik’2 determined thatEq. (9) may yield adequate results ifthe terms I and Q are determinedbased on a cracked transformed section, where the section would betransformed using the slab-to-beammodular ratio used in flexural designby the allowable stress method. However, this approach uses two limitstates in the calculations (service andultimate limit states). Further, crackedsection analysis is complicated in pre

stressed concrete sections becausethey are subjected to combined axialload and bending.

The ACT 318-02 Code13 and theAASHTO Standard Specifications14provide that the following equationmay be used in calculating vh:

= -‘-- (10)

where d is the effective depth from theextreme compression fiber to the tensile reinforcement.

Using equilibrium of forces, Kamel9showed that Vh can be determined by:

VI = (11)V V

where d, is the distance between thetension and compression resultantstresses in the section, for sectionswhere the compression resultant iswithin the slab depth. The symbol dis the same one already used for verti

cal shear design in the AASHTOLRFD Specifications.15

The formula in the ACI 318 Code’3and the AASHTO Standard Specifications14 may be viewed as a simplifiedform of Eq. (11). However, this approximation somewhat underestimates the required horizontal shearstress at the interface for a given load.Obviously, this underestimation hasnot been a problem in applying bothcodes as the applied loads are vastlyoverestimated.

The authors recommend that Eq.(11) be used in calculating the factored horizontal shear stresses forcomposite members with the newdebonded shear key interface systemas well as with the conventionalroughened interface system. A conservative approximation of Eq. (11) is toset d as a constant throughout member length, equal to (d — a12), where ais the equivalent rectangular compression block depth at the maximum pos

Precastbeam

]resistdbutio

(a) Beam under its self weight

Precastbeam

(b) Beam under its self weight and slab weight

(9)

(c) Composite beam under superimposed loads

Fig. 5. Composite beam behavior.

May-June 2002 63

\

L_‘I’ II

fl p23.62’ (600 mm)

DL// pD .

Fig. 6. Standard details of debonded shear key system.

itive moment section. The effectivedepth d may be a variable if drapedstrands are used. In this situation, therules for vertical shear design shouldbe applicable here.

The AASHTO LRFD Specifications15 give no guidance to bridge designers for computing horizontalshear stresses due to factored load. Inretrospect, the authors believe that theLRFD Specifications15should providesuch guidance.

MAXIMUMREINFORCEMENT UMIT

Various maximum reinforcementlimits are available.9Section 5.8.4.1 ofthe LRFD Specifications’3shows thefollowing limits:

Af 0.2fC’bVSSk

where f’ is the strength of the beam orthe slab, whichever is smaller, and

may not be taken greater than 4000 psi(27.6 MPa). This limit, which was developed in the late 1970s, is very conservative. The authors recommendusing the limit given by Loov and Patnaik’2 in 1994:

Affy0.25fC’bSk (13)

STANDARDIZED SHEAR KEYDIMENSIONS

The authors worked with the Nebraska Department of Roads (NDOR)and the Prestressed Concrete Association of Nebraska to develop standarddimensions for the shear keys thatmeet the design criteria given by Eqs.(6) and (7) for the vast majority of

(12) bridge spans and girder spacings usedin practice. The range considered wasfrom 40 to 150 ft (12.2 to 45.7 m)spans and up to 12 ft (3.6 m) spacing.

Fig. 6 shows the standard dimensionsof the shear keys.

These dimensions were developedin hard metric units for consistencywith the other NU girder dimensions.An 11.81 in. (300 mm) module of theshear key spacing was used. Shearconnectors consisting of two No.5(No. 16) U-shaped bars are providedat a spacing of 11.81 or 23.62 in. (300or 600 mm), depending on the required shear capacity. A 0.75 in. (19mm) shear key depth was used.

It should be noted that the top flangeis covered with the deck concrete andshould not be subjected to the standardminimum cover requirement of corrosion protection of reinforcement in thetop flange.

These dimensions selected for thestate of Nebraska are believed to be adequate for I-girders with similar flangesto those of the NU I-girder. For otherstandard precast concrete product

b= 48.22” (1225 mm) 2- #5 (#16) U epoxy-coated bars @ 23.62” (600 mm)outside diameter = 2.95” (75 mm)

14.96” (380 mm) A

2.95’ (75 mm)

A

Horizontal shear reinforcement detail

Plan View A-ANU Girder

Shear Key Typical Dimensions

- 2#5(#16)U bar

@ 23.62” (600 mm)

LSk =0.75 (19 mm)

Shear connector location Shear connector location

64 PCI JOURNAL

I— I.

Fig. 7. Conventional details of northbound and new details of southbound structures of the Waterloo Northwest Bridge.

shapes, Eqs. (6) and (7) may be used todevelop standard shear key dimensions.

DESIGN CRITERIA AND

SPECI F ICATIONS

Using the shear key dimensionsshown in Fig. 6 and a specified concrete strength of the deck of 4000 psi(27.6 MPa), the following procedure issuggested:

1. Using continuous beam analysissoftware, calculate the unfactored (service load) vertical shear forces pergirder line due to composite deadloads (barriers, overlay, and futurewearing surface) and live loads.

2. Determine the factored verticalshear forces due to composite deadand live loads, V, according to theadopted specifications.

3. Calculate the factored horizontalshear stress due to composite dead andlive loads using Eq. (11).

4. Check that Vuh 137 psi (0.94MPa). This limit is developed basedon Eq. (6) in order to protect the shearkey against bearing failure, using theshear key dimensions given in Fig. 6and the 4000 psi (27.6 MPa) concretestrength specified for the deck. If Vh

exceeds this limit, the specified concrete strength of the deck, 4000 psi(27.6 MPa), should be increasedand/or the dimensions of the shear keyshould be changed. This can be doneby solving simultaneously Eqs. (6)and (7). Since these equations havefour variables (tk, bk, Wsk, and fD twovariables have to be chosen in advanceby the designer and then the remainingtwo variables can be calculated.

5. Determine the required horizontalshear reinforcement, Af, using Eq. (8).

6. Check the maximum reinforcement limits using Eq. (13).

The design steps can be easily programmed into a spreadsheet. A de

tailed example illustrating this procedure is given in Appendix B.

DEMONSTRATION PROJECT

The Nebraska Department of Roads(NDOR) assigned a simple spanbridge in Nebraska to implement thedebonded shear key system.21 The project is Structure No. 5275-1584L&Ron Highway 64, Waterloo Northwest,Douglas County, Nebraska. It consisted of two identical 128 ft (39.0 m)long simple-span northbound andsouthbound structures.

The cross section of each bridgeconsisted of five NU1600 girdersspaced at 8 ft 2’/2 in. (2.5 m) supporting a 7.5 in. (190 mm) thick compositecast-in-place slab with specified compressive strength, of 4000 psi. Thetotal width of each bridge was 40 ft 8in. (12.4 m).

This particular site was selected in

9.84(250 mm)

HI I

48.2 (1225 mm)

4,DI 8 (MDI 16)

5.91” —.

________

(150 mm) -

________________

6(160 mm)

NU1600 Girder

2 #5 (#16) U

5.91”

-Shape Bars

(150 mm)

______________________

(380 mm)14.96

NU1600 Girder

II

GirderNorth-bound structure centerline

5.91” (150mm)

_____________

19 spaces @ 23 6

11 II I1 Ii ji ii II

South-bound structure

May-June 2002 65

order to allow the use of a conventional roughened interface system onthe northbound structure and the newdebonded shear key system on thesouthbound structure. This arrangement made it possible to compare theperformance of the two systems and toobtain accurate estimates of the incremental cost of the new system.

The preliminary design of the horizontal shear reinforcement of the composite action using the conventionalinterface system included 2-D18(Dl 16) at a spacing ranging from 1.96to 11.81 in. (50 to 300 mm). Using thedebonded shear key system resulted in

two No. 5 (No. 16) U-bars at a spacingranging from 11.81 to 23.62 in. (300to 600 mm), as shown in Fig. 7.

The precast concrete girders wereproduced by Concrete Industries, Inc.,Lincoln, Nebraska. The precast concrete producer elected to form theshear keys with 3/4 in. (19 mm) plywood blockouts, due to the limitednumber of girders used and the uncertainty about the final standard detailimposed by the Nebraska Departmentof Roads, which would be decidedafter review of the performance of thedemonstration project.

Production of the girder with the

shear keys went smoothly without anyproblems. The top surface of the exposed shear key was steel troweled toa smooth finish. After the girders wereremoved from the prestressing bed,the top flange was sprayed with asealant to break the bond between thetop flange of the girder and the deckconcrete. Figs. 8 and 9 show the girders with the shear keys.

The precast producer bundled thetwo No. 5 (No. 16) U-bars at the webcenterline to avoid interference withthe draped strands at the girder ends.This arrangement did not allow thedeck concrete to fully surround theshear connectors. Because of the possibility of reduction in resistance, theauthors recommended that a cross barbe required to pass through the U-barsin order to provide a stronger mechanical anchorage between the concretedeck and the connectors. This could bedone by passing the deck slab reinforcement or short pieces of barsthrough the connectors.

To determine the size and length ofthese bars, the anchorage requirementsthat are used for the head of headedstuds for steel girders were used. Theratio of anchor bar horizontal projection area to the cross-sectional area ofthe connectors being anchored was setequal to the ratio of the area of the studhead and the area of the stud stem. Thisresulted in using one 10 in. (254 mm)long, No. 6 (No. 19) epoxy coated barfor each pair of No. 5 (No. 16) U-shaped connectors. Fig. 10 shows a topview of the bridge before adding theshort crossbars and placing the deck.

Construction of the demonstrationproject started in the early summer of1998 and was completed by the end ofsummer of 1999. Deflection measurements of the southbound and northbound bridges were taken on November 30, 1999. A three-axle truck wasused. The truck was positioned suchthat one line of wheels was set directlyover the center girder and that the center axle was set exactly over themidspan point.

The weight of each axle wasrecorded before and after taking themeasurements. Table 1 shows the fielddeflection measurements and compares them with those determined byanalysis. From the deflection measure-

Fig. 8. Top view of debonded shear key system.

66 PCI JOURNAL

ments, the following conclusions canbe drawn:

1. The field deflection measurementof the southbound structure, built withthe debonded shear key system, wasalmost the same as that of the northbound structure, built with the conventional roughened shear key system.This indicated that the new shear keysystem did not increase the flexibilityof the bridge.

2. The field deflection measurements were very small compared tothe theoretical deflection. Thus, theactual structure was stiffer than thatconsidered in analysis. This may havebeen due to a number of factors:

(a) The concrete girder may havehad a higher modulus of elasticity thanthat corresponding to the specifiedstrength.

(b) The number of girders shared incarrying the truck load may have beenmore than that indicated in the designspecifications.

(c) The abutment support conditionsmay have created more girder end restraint than that used in the design.

(d) The New Jersey barriers mayhave increased the overall stiffness ofthe bridge cross section.

3. Three months after completion ofthe bridge, field inspection revealedno visible cracks in the superstructure,including at the girder/deck interface,and no signs of distress in the deck oneither the northbound or the southbound structure.

COST ANALYSISUsing the dimensions of the demon

stration bridge, the expected incremental production cost of a girder dueto the addition of the debonded shearkey system can be summarized as follows:

1. Extra cost due to use of No. 5(No. 16) U-bars (assuming 35 centsper pound for epoxy coated bars) =

$126.00 per girder.2. Reduction in weight of the verti

cal shear reinforcement due to barsbeing terminated at the top flange (assuming 35 cents per pound for non-epoxy coated welded wire reinforcement) = $86.00 per girder.

3. Sprayed debonding agent (assuming 10 cents per sq ft) = $26.00.

4. Troweling the spacing betweenthe shear key forms may be assumedequal to roughening the entire topflange in conventional production.

5. Net incremental cost = 126 — 86+ 26 = $66.00 per girder = ($66.00 x 5girders)/(128 ft x 40.75 ft) = $0.063per sq ft.

It should be mentioned that the actual incremental cost of the southbound structure girders was higherthan the above figure because woodforms were used to produce the shearkeys. This resulted in an additionalcost of $305.00 per girder for materialand labor used in making the forms,extra production labor of $437.00used for installing and removing thewood forms, and $88.00 extra engineering cost.

The authors believe that when thesystem becomes standard practice andsteel forms are used in future projects,the extra cost due to production, installation, and removal of the formswould be eliminated. Also, with repeated usage of this system, the precast producer would not incur an extraengineering cost. It is also important

CONCLUSIONS ANDRECOMMENDATIONS

A debonded shear key interface system has been developed for precast,prestressed concrete girders madecomposite with concrete decks. Thenew system has the advantages of facilitating future deck removal, protecting shear connectors against corrosion, protecting the girder top flangefrom damage during deck removal,and optimizing the design for horizontal shear.

Based on theoretical results, laboratory testing and field implementation,the proposed debonded shear key sys

to note that the long term savings intime and labor associated with futuredeck removal due to the use of thedebonded interface system will compensate for any incremental cost.

Although the debonded shear keyinterface system has been developed,tested, and implemented on I-girders,it may be used for other types of girders such as inverted tees, box beams,U-beams, and segmental box girders.

Table 1. Deflection measurements of northbound and southbound structures.

I Southbound structure Northbound structure

I (Debonded shear (Conventional roughenedDeflection key system) .. shear key system)

L_Field measurements 0.08 in. 0.04 in.Analysis 0.44 in. 0.44 in.

Note: 1 in. = 25.4 mm.

May-June 2002 67

tern has been found to perform well andhas no detrimental effects on compositeaction or bridge stiffness. The incremental cost of using the new shear keysystem is expected to be minimal andshould eventually turn to savings whenthe reduction in time and labor duringdeck removal is accounted for.

The following recommendations areput forth:

1. The AASHTO LRFD shear frictionprocedure is valid for calculation ofinterface shear strength using the proposed details, with a friction coefficient, s = 1.0, and a cohesion stressfactor, c = zero.

2. It is recommended that only superimposed loads applied after theconcrete deck has hardened, rather

than the full loads, be used in designfor composite action.

3. It is recommended that the required horizontal shear stress due tofactored loads be calculated at all sections using the same d used in theLRFD Specifications for vertical sheardesign. This distance between the tension and compression stress resultantsmay be conservatively assumed equalto that corresponding to the compression block depth at the maximum positive moment section.

ACKNOWLEDGMENTThis research program was con

ducted under NCHRP Project 12-41,sponsored by the Transportation Re-

search Board (TRB), and Project SPRPL-1(35) P5 16, sponsored by the Ne

braska Department of Roads (NDOR).The support of Amir Hanna of TRB,Leona Kolbet, Lyman Freemon, GaleBarnhill, Mike Beacham, and SamFallaha, of NDOR Bridge Division,and Larry Fischer, Concrete Industries, Inc., is gratefully acknowledged.

The authors would like to expresstheir appreciation to the Center of Infrastructure Research and the graduatestudents of the University of Nebraska, who assisted in the testing program.

Lastly, the authors wish to expresstheir gratitude to the PCI JOURNALreviewers for their constructive comments.

REFERENCES

1. Bridge Office Policies and Procedures (BOPP) Manual, Nebraska Department of Roads (NDOR), Lincoln, NE, 2001.

2. Tadros, M. K., and Baishya, M. C., “Rapid Replacement ofBridge Decks,” Report 407, National Cooperative HighwayResearch Program, NCHRP, National Research Council,Washington, DC, 1998.

3. Hanson, N. W., “Precast Prestressed Concrete Bridges 2. Horizontal Shear Connections,” Journal PCA Research and Development Laboratories, V.2, No.2, May 1960, pp. 38-58.

4. Kaar, P. H., Kriz, L. B., and Hognestad, E., “Precast Prestressed Concrete Bridges 1. Pilot Tests of Continuous Girders,” Journal PCA Research and Development Laboratories,V. 2, No. 2, May 1960, pp. 21-37.

5. Kriz, L. B., and Raths, C. H., “Connections in Precast Concrete Structures — Strength of Corbels,” PCI JOURNAL, V.l0,No. 1, February 1965, pp. 16-61.

6. Hofbeck, J. A., Ibrahim I. 0., and Mattock, A. H., “ShearTransfer in Reinforced Concrete,” ACI Journal, V. 66, No. 2,February 1969, pp. 119-128.

7. Mattock, A. H., Li, W. K., and Wang, T. C., “Shear Transfer inLightweight Reinforced Concrete,” PCI JOURNAL, V. 21,No. 1, January-February 1976, pp. 20-39.

8. Walraven, J., Fronay, J., and Pruijssers, A., “Influence of Concrete Strength and Load History on the Shear Friction Capacityof Concrete Members,” PCI JOURNAL, V. 32, No. 1, January-February 1987, pp. 66-84.

9. Kamel, M. R., “Innovative Precast Concrete Composite BridgeSystems,” A Dissertation Submitted to the Faculty of the Graduate College at the University of Nebraska in Partial Fulfillment of Requirements for the Degree of Doctor of Philosophy,Lincoln, NE, 1996.

10. Shaikh, A. F., “Proposed Revisions to Shear-Friction Provisions,” PCI JOURNAL, V. 23, No. 2, March-April 1978, pp.12-2 1.

11. PCI Industry Handbook Committee, PCI Design Handbook —

Precast and Prestressed Concrete, Fifth Edition, PrecastlPrestressed Concrete Institute, Chicago, IL, 1995.

12. Loov, E. R., and Patnaik, A. K., “Horizontal Shear Strength ofComposite Concrete Beams With a Rough Interface,” PCIJOURNAL, V. 39, No. 1, January-February 1994, pp. 48-69.

13. ACI Committee 318, “Building Code Requirements for Structural Concrete (318-02) and Commentary (31 8R-02),” American Concrete Institute, Farmington Hills, MI, 2002.

14. AASHTO, Standard Specifications for Highway Bridges, 16thEdition, American Association of State Highway and Transportation Officials, 1996, with 1997, 1998, 1999, and 2000 Interims, Washington, DC.

15. AASHTO, AASHTO LRFD Bridge Design Speccations, Second Edition, American Association of State Highway andTransportation Officials, 1998, with 1999, 2000, and 2001 Interims, Washington, DC.

16. Badie, S. S., Baishya, M. C., and Tadros, M. K., “NUDECKAn Efficient and Economical Precast Bridge Deck System,”

PCI JOURNAL, V. 43, No. 5, September-October 1998, pp.56-74.

17. Issa, M. A., Idnss, A., Kaspar, I. I., and Khayyat, A. Y., “FullDepth Precast and Precast Prestressed Concrete Bridge DeckPanels,” PCI JOURNAL, V. 40, No. 1, January-February1997, pp. 59-80.

18. Yamane, T., Tadros, M. K., Badie, S. S., and Baishya, M. C.,“Full-Depth Precast Prestressed Concrete Bridge Deck System,” PCI JOURNAL, V. 43, No. 3, May-June 1998, pp. 50-66.

19. Kakish, H. F., “Composite Action in Bridge I-Girder Systems,” A Dissertation Submitted to the Faculty of the GraduateCollege at the University of Nebraska in Partial Fulfillment ofRequirements for the Degree of Doctor of Philosophy, Lincoln, NE, 1997.

20. Bridge Manual, Illinois Department of Transportation, Bureauof Bridges and Structures, Springfield, IL, 1999.

21. Badie, S. S., and Tadros, M. K., “I-Girder/Deck Connectionfor Efficient Deck Replacement,” Project No. SPR-PL-1(35)P516, Final Report, Nebraska Department of Roads (NDOR),Lincoln, NE, August 2000.

68 PCI JOURNAL

APPENDIX A — NOTATION

a = equivalent rectangular compression block depth(in.)

= area of shear key at base (sq in.)Ah = area of horizontal shear reinforcement crossing in

terface (sq in.)bk width of shear key at base (in.)b = width of cross section at contact surface being in

vestigated for horizontal shear (interface width)(in.)

c = cohesion stress (psi)d = effective depth from extreme compression fiber to

the tensile reinforcement (in.)d = distance between tension and compression resul

tant stresses in section (in.)= specified compressive strength of deck concrete at

28 days (psi)

f = specified minimum yield strength of reinforcingbars (psi)

I = moment of inertia of composite section (in.4)

Q first moment of area above interface (in.3)s = spacing of horizontal shear reinforcement (in.)SSk = shear key spacing (in.)tyk = depth of shear key (in.)VL+, = unfactored vertical shear due to HS2O AASHTO

truck loading with impact (lbs)= nominal horizontal shear stress (psi)

Vflh = nominal horizontal shear capacity (Ibs)VSID = unfactored vertical shear due to superimposed

dead loads (lbs)V = factored vertical shear force at a section (lbs)Vuh = factored horizontal shear stress (psi)VUh = factored horizontal shear force (ibs)Wk = length of shear key at base (in.)

= strength reduction factor= coefficient of friction

APPENDIX B — DESIGN EXAMPLE

Using the proposed debonded shear key system, designthe reinforcement required at the interface of a NU 1600 I-girder build supporting a 7.5 in. (190 mm) thick compositecast-in-place slab.

Input Data

At Section xIL = 0.1, the unfactored vertical shear forcedue to superimposed dead loads, VSJD = 13,600 lbs (60.3kN), and due to HS2O AASHTO truck load with impact VL+I

= 67,300 lbs (299.2 kN).Depth of tensile reinforcement, d = 60.12 in. (1527 mm),

and depth of equivalent stress block, a = 7.0 in. (177 mm).Use the shear key dimensions given in Fig. 6.Step 1: Unfactored vertical shear due to composite dead

and live loads:V10= 13,600 lbs (60.3 kN)VL+I 67,300 lbs (299.2 kN)

Step 2: Factored horizontal shear stress due to compositedead and live loads:

V, = 1.3 [VSID + (S/3)VL+Jl= 163,500 lbs (727.2 kN)

Step 3: Factored horizontal shear stress due to compositedead and live loads:

d = 60.12 in. (1527 mm)a =7.Oin.(177mm)

= d — a/2 = 56.62 in. (1438 mm)

Step 4: Check that Vh 137 psi (0.95 MPa) OK

Step 5: Required horizontal shear reinforcement,

Af- 0.9f

— 60.0 x 48.2 x 11.81

— 0.9 x 60,000

= 0.63 sq in. / SSk (408 mm2 / SSk)

Use two No. 5 (No. 16) U-bars at 23.62 in. (600 mm).Aprovided = 4 x 0.31 = 1.24 sq in./23.62 in.

Step 6: Check maximum reinforcement limits:AVffy 0.25f’(bS2k)(0.5 x 1.24)(60,000) = 37,200 lbs (165.5 kN)0.25(4000)(48.2 x 11.81)= 569,242 lbs (2532.0 kN)OK

V =h bd

— 163,500

— 48.2 x 56.62= 60.0 psi (0.414 MPa)

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