Shear Stud Cap in Steel Decks

7/28/2019 Shear Stud Cap in Steel Decks

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SHEAR STUD CAPACITY IN PROFILED STEELDECKS

RAED J. ZAKI is theAssistant Structural Engineer

at HERA. He obtained his

BE at the University of

Baghdad, Iraq and his

Masters at the University of

Auckland, New Zealand. He

is currently working on the

application of steel semi-

rigid joints in a damage

avoidance seismic system

and on improved design

provisions for shear studs in composite floor

systems.

JOHN BUTTERWORTHis Deputy Head,

Department of Civil and

EnvironmentalEngineering, University of

Auckland. His research

interests include behaviour

of steel structures,

earthquake engineering,

passive seismic protection,

structural dynamics, thin-

walled metal sections,

composites, structural

stability and large scale dynamic testing of bridges

and buildings.

G. CHARLES CLIFTONheads the Structural

Division of HERA (the

New Zealand Heavy

Engineering Research

Association). His role there

is to promote the proper

and effective use of

structural steel in New

Zealand. His principal

activities are research,

development of design

guidance and technicalpromotion. He runs HERAs fire and seismic

research programme, which is aimed at improving

the safety and cost-effectiveness of structural steel

building performance in the high risk but infrequent

events of severe fire and severs earthquake. He is a

Fellow of the Institution of Professional Engineers

New Zealand and of the New Zealand Society for

Earthquake Engineering.

ABSTRACT

The New Zealand design procedure for composite

beams incorporates provisions for headed, welded

shear stud design shear capacity developed forNorth American standards in the 1970s, modified

with a partial strength reduction factor incorporated

in 1992. Following New Zealand and international

tests which showed shortcomings with these

equations, the topic of composite beam design

using shear studs and trapezoidal decking was re-

evaluated in 2003 and a series of experimental tests

undertaken to determine the adequacy of proposed

new provisions.

This paper covers the following: the changes made,

the previous reports on over-estimating the stud

capacity, the testing of eleven shear stud push-

through specimens and the results of these tests.

Also included are constructional issues such as stud

positioning, the effect of decking edge in primary

beams and the influence of the self weight of the

slab on the transverse bending moment acting on

the beam and its effect on the shear studs. The

results of these tests have allowed us to make

recommendations for shear stud capacity, which

more accurately reflect the failure modes that

occurred, and to incorporate these into revisions to

the composite beam design procedure for beams

supporting concrete slabs on profiled steel decks.

Raed J. Zaki

ohn Butterworth

G. Charles Clifton


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1. INTRODUCTION

The use of composite steel beams with profiled

steel decking and in-situ concrete floors has

increased in recent years in New Zealand and

around the world. This increased interest led to a

critical review of current design procedures in

(NZS 3404:1997) arising from concerns that theseprovisions overestimate the strength of shear studs

in profiled steel decks. A number of studies, (Lin,

et. al., 2001), (Jayas and Hosain, 1988 & 1989),

have shown that many building codes over-estimate

the predicted ultimate shear stud capacity, in

comparison with the observed shear stud capacity.

This has led to updating of standards; for example

the latest (CSA S16-01: 2001) has been revised to

correct this error.

This paper presents an overview of research

recently undertaken (Zaki, et. al., 2003) into the

adequacy of current New Zealand design provisions

for composite beams supporting profiled steel

decks. Following a detailed comparison of the

overall New Zealand composite design procedure

(NZS 3404:1997) with that from the United

Kingdom (BS 5950(1990)), Europe (Eurocode 4:

ENV 1994-1-1:1994), and Canada (CSA

S16:2001), recommendations for changes to the

design provisions and to the equations for

determining the capacity of headed, welded shear

studs have been made. The latter were basedinitially on a review of the international standards

mentioned above and an experimental testing

programme was undertaken.

Figure 2.1. Concepts of composite construction

(Oehlers and Bradford 1995).

That programme has led to further recommended

changes to the shear stud capacity equations. The

end result has been minor changes to the composite

beam design provisions and a revision to the

equations for determining shear stud capacity. The

overall effect will be more accurate determination

of the number of shear studs needed in a composite

beam. Further areas of required research have also

been identified.

This paper firstly reviews, in section 2, the concepts

involved in composite action. This includes the

strength of shear connectors in solid slabs and in

profiled slabs, covering the various failure modesidentified in the literature.

While one of the main drivers behind this research

was concern over the design capacity of shear studs

as predicted by (NZS 3404:1997), it is important to

note that the calculated shear strength of a welded

stud is only one of the factors that determines the

shear stud numbers and positions required in a

composite beam. The first stage of the research

therefore involved a comparative calculation study

of the design requirements for a range of steel

beams designed to New Zealand, British, European

and Canadian practice. Section 3 presents details of

this work. The conclusions from this research are

presented in section 7, followed by the references.

2. CONCEPTS OF COMPOSITE

CONSTRUCTION

2.1 Composite Steel-Concrete Sections:

Composite construction is based upon connecting a

concrete slab to a steel beam so that the twoelements act as one. By placing the steel beam in

tension, the high tensile strength of the beam can be

fully used, and the possibility of buckling of thin

steel members can be reduced. Likewise, by placing

concrete slabs in compression, the disadvantage of

weak concrete tensile strength can be avoided and

the advantage of good compression strength/cost

fully utilised. Also, concrete slabs have good sound

and fire insulation properties. A common form of

construction is to use a steel beam to support the

slab, as shown in figure. 2.1(a) and (b). It has been

found that by bonding the steel beam to the

steel/concrete interface

strain distribution

composite beam

A

(a)

reinforced-concrete slab

A

(c)composite slab

steel decking or profiled sheet

composite beam

steel beam

(b)

A-A

B C B

C

concrete

B

cold formed ribs

weff

x

x

(a) Full composite beam

beff

Profiled, cold - formed steeldecking.

(a) Full composite beam-strain distribution.


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concrete slab by shear connectors, the flexural

strength and stiffness of the beam can be at leastdoubled (Oehlers and Bradford, 1995). This in turn

leads to an increase in the span to depth ratio that

can be used.

Figure 2.1 (a) shows the strain distribution

associated with full interaction between the steelbeam and the concrete slab. In this instance, there is

no interface slip between them. In practise, a

limited degree of interface slip is usually allowed

and incorporated into the design; this is known as

partial composite action.

Another type of composite structure is formed

when profiled steel decking is used as a permanent

formwork for a concrete slab and is designed to act

integrally with the concrete so that the two

components act as one, as shown in figure 2.1 (c).

Figure 2.2 Composite beam with profiled steel

decking.

The most commonly used composite section is the

one that connects the concrete slab with the steel

beam by welding connectors between them to

achieve composite action, as shown in figure 2.2.

The connector is welded through the troughs of the

profiled decking to the flange of the steel beam

(Oehlers and Bradford, 1995). The ribs of the

profiled decking, through which the interaction

Figure 2.3 Shear stud and transfer of connector

force (Oehlers and Bradford 1995)

between the steel decking and the concrete is

achieved, can be placed either parallel, or

transverse to the steel element. New Zealand is one

of the very few countries to have all-weather stud

welding capability, which elevates the economic

viability of headed studs when compared with other

anchor types as it allows these to be welded through

the deck or directly onto the beam even in wetweather. The study reported herein focuses on this

type of composite construction and specifically on

the connector between steel beam and concrete slab

known as the headed, welded shear stud (hereafter

referred to as a shear stud).

2.2 Shear Stud.

The shear stud, as shown in figure 2.3 (a), is the

most commonly used shear connector in composite

construction. The shank and the weld-collar

adjacent to the steel flange are designed to resist the

longitudinal shear force, whereas the head is

designed to resist the tensile force that is normal to

the steel/concrete interface and to engage the

concrete. The shank diameter of headed studs

ranges from 10mm to 22mm. The most commonly

used is the 19mm diameter. The length of headed

studs before welding varies from 55mm to 250mm.

After welding the stud length will have decreased

by 3mm to 11mm, depending on the stud diameter

and the stud application, i.e. welding through

decking burns off some 2-4mm more than cleanbeam welding (Clifton, 1998).

Oehlers and Bradford (1995) noted that shear studs

are simply steel dowels embedded in a concrete

medium. As the concrete slab tries to slip relative to

the beam, the stud imposes a very high

concentrated force, which is transferred from the

dowel action of the stud into the concrete, as

illustrated in figure 2.3(b). The resistance of a stud

to this dowel action is known as the dowel strength,

and this strength is specified in the New Zealand.

Standard (NZS 3404:1997) as the nominal shear

capacity of connectors. The dispersing of this

force into the concrete element can induce tensile

cracking particularly when the connector is also

(c) plan view of slab

showing crack formation

Theight

(a) stud

diameter

(b) dowel action

C

embedment

cracking

weld collar

shank

headsteel dowel

steel element

concrete element

ripping

C

T

shear splitting

steel

concrete

Transversecracking

Splitting

Longitudinal

(a) Ribs Parallel

(b) Ribs Transverse

Composite Slab


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resisting separation at the steel/concrete interface of

the composite beam, which in turn is caused bytransverse cracking, shear and longitudinal splitting

actions. Figure 2.3 (b) shows the tensile cracks that

are also known as embedment cracks, and figure

2.3 (c) shows the dispersal of the dowel action

inducing a longitudinal splitting crack.

These forms of tensile failure of the slab can affect

both the dowel strength and ductility of the shear

connection. With the absence of reinforcing bars

crossing the planes of cracking, the strength of the

dowel reduces immediately after cracking occurs

and the slip capacity is also reduced. On the other

hand, the presence of reinforcement across a crack

plane makes failure more ductile and even allows

increases in stud shear resistance after cracking.

2.3 Mechanism of dowel action.

Figure 2.4(a), from (Oehlers and Bradford, 1995),

demonstrates the transfer of longitudinal shear by

dowel action of shear studs. That figure shows the

resulting mechanism where the stud is forced to the

right when the steel beam moves to the right,

causing the stud to bear on the concrete to the right

of the stud. For the dowel action to work, the

concrete adjacent to the bearing zone has to

withstand compressive stresses up to seven times

the cylinder strength, fc, which is only achieved by

the triaxial restraint imposed on the region by thesteel element, the stud and the surrounding

concrete.

The resultant force in the bearing zone in figure

2.4(a) is F, the centroid of which occurs at an

eccentricity, e, from the steel/concrete interface.

There is horizontal equilibrium between the force F

and the shear force in the steel element and, to

maintain the rotational equilibrium, a moment is

induced at the base of the stud. Therefore, the steel

stud must resist both flexural and shear forces,

Figure 2.4 Dowel mechanism (Oehlers and

Bradford 1995)

which are the cause of high tensile stresses in the

steel failure zone shown in figure 2.4 (c). Due to

these forces the concrete can crush in the bearing

zone and the steel can fracture in the steel failure

zone, therefore the mechanism of dowel failure is

governed by the interaction between the steel and

concrete failure zones. This can be best described

by considering the equivalent dowel mechanism infigure 2.4 (b). As shown in figure 2.4 (b), the dowel

behaviour can be considered to be a steel beam

resting on a concrete medium (Oehlers and

Bradford, 1995), where section a-a at the midspan

of the steel beam in figure 2.4(b) is equivalent to

section a-a at the steel/concrete interface in figure

2.4(a).

Consider the distribution of pressure at the

steel/concrete interface in (b), for a constant applied

force 2F and for different configurations of the

material properties. When the steel modulus Esapproaches infinity, the pressure at the interface can

be considered uniform and hence the resultant force

F in one shear span will act at an eccentricity e=h/2.

The section at midspan has therefore to resist a

shear force F and a moment Fh/2, and the dowel

strength of the stud is determined at this section.

Conversely, when Es approaches zero, the steel

beam can be visualized as a layer of paper such that

the resultant interface force in one shear span is

now almost in line with the applied force, so that e

tends to zero. Therefore, the midspan of the beamnow only resists a shear force F. It can therefore be

seen that as Es increases relative to the concrete

modulus Ec, shear stud capacity reduces because

the flexural component Fe at midspan increases.

On the other hand decreasing Ec is equal to

increasing Es, by increasing e which in turn

increases the flexural component Fe therefore

reduces the shear stud capacity. Therefore, the

shear stud capacity is proportional to the modular

ratio parameter Es/Ec, no matter what the shape or

size of the steel dowel.

(b) equivalent dowel mechanism

mid-span

steel beam Es,fu

(a) dowel mechanism

FFe

a ae

cracktip

concrete medium

Ec,fc

shear stud Es,fu

F

2F

F1h F

(c) dowel mechanism with weld collar

Fe1

concrete medium

Ec,fc

dh

a a

F

e

e

steel failure

zone

e2F2

concrete

bearing zone

dsh

weld collar

h co

h st


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Since the studs are welded into the steel element

and specifically in the flange of the steel beam, aweld collar with a diameter of about 1.3 times the

shank diameter dsh and a height of about 0.3 dsh,as

shown in Figure. 2.4 (c), is formed. This collar

increases the shear stud capacity by increasing the

bearing surface at the stud/concrete interface, which

in turn reduces the stresses in the concrete for agiven force F. It also increases the shear stud

capacity by raising the steel failure zone so that it

occurs at the weld-collar/shank interface. Raising

that zone will subject the zone to a portion F2 of the

total shear force F. The weld collar also reduces the

flexural stresses at the failure zone since the

eccentricity e2 of F2 in (c) is less than the

eccentricity e of F in (a).

2.4. Strength of Shear Stud Connectors.

2.4.1 Introduction.

Since the stud is embedded in concrete, the shear

stud capacity is dependent on the properties of the

concrete slab in addition to the properties of the

steel stud. If the concrete in the bearing zone splits

or crushes, it will cause the base of the stud to

move forward, while the head of the stud remains

in the previous position. This will lead to a larger

slip at the steel beam/concrete slab interface than

that caused only by the deformation of the stud.

This failure may also lead to rotation of the stud,whichwill generate flexural rotation actions and anadditional axial tensile force in the stud. The

magnitude of this axial tensile force increases as the

rotation of the stud increases, tending to pull the

stud out of the concrete, according to (Hyland, et.

al., 2000). The strength of the concrete that resists

stud pullout is called the embedment strength of the

shear stud, while the strength of the concrete that

resists local splitting is known as Vsplit.

If the shear stud doesnt fail in a push-through test

by pulling out of the concrete (i.e. by embedment

failure, in which the stud pulls out of the concrete

along with a cone of surrounding concrete), then it

will fail either through concrete crushing/stud

fracture or through longitudinal splitting. The first

case is governed by the dowel strength of the stud

in uncracked concrete and this is described in

section 2.4.2. The second case is covered in section

2.4.3.

2.4.2 Dowel Strength of Shear Studs in

Uncracked Concrete

Ollgaard, et. al. (1971) pioneered research on the

dowel strength of shear stud connectors, and

identified the important parameters that control the

dowel strength. Using statistical analysis, they

derived an equation to determine the mean dowel

strength of shear stud connectors in push off

specimens, in which the concrete slab had not failed

prematurely through splitting, shear or embedment.

This equation for the dowel strength as dictated by

concrete crushing, Dmax, is given by:

44.0c

3.0cshmax EfA83.1D = (2.1)

where the units are in N and mm, Ash is the cross-

sectional area of the stud shank and Ec is theYoung's modulus of elasticity of the concrete,

which is given by equation 2.2a

c5.1

c f043.0E = (2.2a)

However, New Zealand Standard (NZS 3101:1995)

gives another expression for Ec

( )5.1

cc2300

6900f3320E

+= (2.2b)

where Ec and cf are in N/mm, and is the density

of the concrete in kg/m3. There is not much

difference between these two expressions,

In adapting equation 2.1 to design use, the

exponents were changed to make the equation

dimensionally correct and simpler, although this led

to a loss of accuracy. The altered equation is

cc

2sc

ccscmax Ef8

dEfA50.0D

== (2.3)

Certain types of stud welding procedure produce

very small or no weld collars, and it would beunsafe to use equation 2.3 directly to determine

their strengths. It was therefore suggested that theshear strength given by equation 2.3 should be used

in conjunction with a reduction factor when the

mean height of the weld collar is less than d sh/5(Oehlers and Bradford, 1995). The reduction factor

Rco is given by

sc

coco

d3

h5

3

2R += (2.4)

where hco is the mean height of the weld collar, anddsc is the shank diameter of the stud.

Equation 2.3 was adopted by (NZS 3404:1997) forcalculating the nominal shear capacity of headed

studs, along with a steel profile reduction factordc,as determined by (Slutter and Discoll, 1965). That

factor related to the reduction in stud strength in a

ribbed slab. Including dc, as specified in section13.3.2 of the standard, this gives:


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dcscuccscdcr AfEfA5.0q = (2.5)

where qr is the nominal shear capacity, and Asc is

the cross-section area of the stud shank which is

equivalent to Ash in equation 2.3. dc is the deckingreduction factor, fu is the minimum tensile strength

of the stud, which is equal to 415 MPa for studs

manufactured to AS 1443 grades 1010 to 1020 orASTM A108 grades C1010 to C1020. A strength

reduction factor (sc)is used to obtain the designcapacity of headed studs, which is equal to 0.8 for

studs situated in positive moment regions and 0.6for studs situated in negative moment regions.

The expression fuAsc in the second half of equation

2.5 represents shear stud failure, i.e. due to steel

failure in the zone shown in figure 2.4 (c). As

described below, such steel failure is principally an

ultimate shear failure and as such is over-estimatedby the expression fuAsc. This was fortuitouslyrectified through the incorporation of the partial

strength reduction factor of 0.8, introduced into the1992 edition of NZS 3404. However, in the case of

shear stud failure, that 0.8 should be part of the

equation itself.

Amendment No. 1 (2001) for Clauses 13.3.2.1 and

13.3.2.2 stated changes to be made on equation 2.5

by multiplying the reduction factor (sc) by (1/0.8)and (qr) by (0.8), as shown in equation 2.6. This

brings the factor (0.8) directly into the equation for

determining shear stud nominal capacity, which

provides better agreement with experimental resultsissued in (Hyland, et. al., 2000) and makes the

embedment equations;

dcscuccscdcr Af8.0EfA4.0q = (2.6)

In (Clifton, 2002), another recommendation to

replace equation 13.3.2.1 of (NZS 3404:1997),

shown here in equation 2.6, by:

scuccscdcr Af8.0EfA5.0q = (2.7)

This more correctly represents the dowel strength

of a stud in solid concrete, with the left-hand side ofthe inequality representing concrete crushing and

the right hand side shear failure, of the steel stud atthe top of the weld collar.

For ribbed slabs, it has been noted (Slutter and

Driscoll, 1965) that the shear stud shear capacity is

reduced, with the reduction dependent on the

orientation of the ribs. This reduction factor has

been established experimentally, example equation

2.8 for ribs perpendicular to the steel beam, (Grant,

et. al., 1977), which has been adopted by (NZS3404:1997).

=

rc

r

rc

sc

rc

dch

b1

h

h

n

85.0(2.8)

where nrc is the number of studs in a rib, hsc is the

total height of the stud, hrc is the height of the rib

and br is the mean width of the rib. However, thereduction factor given in draft Eurocode 4 (Johnson

and Anderson, 1993) is more conservative, using

0.7 as a reduction factor instead of 0.85

=

rc

r

rc

sc

rc

dch

b1

h

h

n

7.0(2.9)

Equations 2.8 and 2.9 are presented as representing

the reduced dowel capacity of the shear studs. They

are experimentally determined on the premise that

the concrete does not undergo longitudinal splittingat the base. However, review of the pictures from

the Canadian push-off tests (Jayas and Hosain,

1988) indicate that splitting was the likely limiting

mode of failure for their tests. It is therefore

considered likely by the authors that the Canadianreduction equations represent a splitting-induced

loss of shear stud capacity due to the profiled natureof the concrete slab. The equation for primary beam

application-which is not presented herein- is alsobased on a test configuration that is not

representative of that used in practice.

Given that splitting strength can often govern the

shear stud capacity, the background to this is nowgiven.

2.4.3 Splitting Strength of Concrete.

a. Primary Beam.

Splitting of the concrete happens when the high,concentrated dowel stress formed by the shear stud

is dispersed laterally into the concrete element, as

shown in figure 2.3(c). This dispersion isrepresented by the arrows marked as C and, to

maintain equilibrium transverse tensile stresses are

induced in the concrete, represented as T. When

these transverse tensile stresses exceed the splitting

tensile strength of the concrete, fcb, longitudinalcracks will form in the concrete prism in the

bearing zone along the line of the shear connectors.

The splitting tensile strength of the concrete fcb isobtained from cylinder split tests, as given by the

following equation:

ccb 'f5.0f = (2.10)

A detailed background to the shear splitting

provisions is given in section 11.3 of (Oehlers and

Bradford, 1995). The relevant equation for the


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nominal concrete splitting resistance of a haunch ofconcrete of width bce is given by:

+=

c

a

2

c

a

cbasc

d

cbeqce

split

h

h

h

h1

fhd6.0

k

fhb6.0V (2.12)

where:

bce = concrete effective haunch width,

ha = 1.8dsc, effective stud bearing height. (2.13)

fc = characteristic concrete cylinder compressionstrength (MPa).

dsc = diameter of welded, headed shear stud (mm).

Vsplit = nominal splitting capacity (N).

kd =

2

ce

sc

b

d1

1

, lateral force parameter. (2.14)

hc = effective haunch height for splitting = min

(4.5ha, to).

The following dimensional constraints are requiredfor equation 2.12 to be valid:

bc 3dsc (2.15)hc 3ha = 5.4dsc (2.16)

hsc hc (2.17)

where:

bc = actual width of insitu concrete into which the

shear stud is embedded (mm)

hsc = height of shear stud after installation (mm)

In addition, especially for individual concrete ribs

containing studs, hsc 4.5ha = 8.1dsc is highly

desirable, as this places the head of the stud, whichacts as the anchor for preventing vertical slip of the

stud within the concrete, outside of the region ofhigh compression bearing stress where splitting

initiates. This has benefits in increasing the post-

splitting shear capacity, as given by (Oehlers andBradford, 1995).

b. Secondary Beam.

Splitting failure in the secondary beam

configurations, ribs transverse to the beam, has

been represented by some researchers as concretepullout failure. (Hawkins and Mitchell, 1984)

proposed the following equation:

ccc A'f45.0V = (2.18)

where:

Vc = shear capacity due to concrete pull-out failure(N).

fc = concrete compressive strength (MPa).Ac = area of concrete pull out failure surface (mm

2).

Jayas and Hosain (1988) made further researchbased on this equation for 38 and 76mm decks,where they found out that instead of a fixed value

of 0.45, separate coefficients should be used for the38 and 76mm decks.

However, in our push-off tests on the secondary

beams, the governing failure mode for thesecondary beam studs was a splitting failure mode,

as shown by the classic longitudinal splitting crack

developing in the concrete slab (see section 6.2).

This has been approached through the use of

equation 2.12, obtaining Vsplit experimentally and

then back calculating to obtain the effective haunchwidth, for the secondary beam, bce,sb.

3. COMPARATIVE CALCULATION PRIOR

TO EXPERIMENTAL TESTING

3.1 Introduction and Scope.

When designing a composite beam, many factors

govern the selection of steel beam size and the

number of shear studs required. These factors maymean that a significant change in shear stud design

shear capacity may not make much difference to thenumber of shear studs required for a given beam

span, configuration and loading.

The purpose of the comparative calculations was to

determine the number of shear studs required for arange of typical primary and secondary beams

designed in accordance with the strengthrequirements of new Zealand, Canadian, British

and European standards. The range of parametersused in this study are given in Table 3.1.

Each beam was designed to meet all the strengthand serviceability requirements of current New

Zealand practise (NZS 3404:1997) and (Clifton,2002). The software HiBond Design Wizard, which

covers use of the 55mm high trapezoidal deckingprofile used in the study, based on (NZS

3404:1997), was used to achieve this.

Having undertaken these designs and determined

the number of shear studs required, the number of

shear studs required for the same beam size andapplied loads to the requirements of the current

Canadian Standard, British Standard and European

Standard were then determined.

The results of these comparisons were then

tabulated in order to compare the current NewZealand requirements with those of three major

international composite standards. This was done in

order to give an indication as to the combined effect

of the many parameters influencing the number ofstuds required on the design solution and to see

what influence any change in shear stud design


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shear capacity would have against currentinternational best practice.

Parameter

Live Loads 3.5 kPa & 5 kPa

Beam Span 8 m & 15 m

Propping No

Precambered NoPartial Shear

Connection

As stated in the respective

Standard

Slab Span 2.8 m

Thickness 120 mm

Deck Type 55 mm trapezoidal profile.

Stud Height 100 mm

Stud Diameter 19 mm

Stud fu 415 MPa

Concrete fc 25 MPa

Table 3.1 Parameters used in the Study.

Section 3.2 presents the New Zealand requirements

for determining the ultimate limit state capacity of asteel beam and the number of shear studs required.

Section 3.3 presents an overview of the strength

design requirements of the other Standards used,

highlighting the differences between these and the(NZS 3404:1997) requirements.

Section 3.4 gives some of the comparative results,first in terms of shear stud design capacity and

secondly in terms of the number of studs requiredin each beam size. Full details are in (Zaki, 2003).

Following this work initial recommendation were

proposed for the shear stud design shear capacityand also for the partial shear connection limits, inorder to keep the net changes to as-built details to a

minimum. The initial recommendations arising atthe end of the comparative calculations are given in

section 3.5.

3.2 New Zealand Provision for Calculation of the

Composite Ultimate Limit State Design Moment

Capacity, Mrc

This involves the following general steps:1. Calculation of the factored design loads and

design actions (moment and shear).

2. Calculating the maximum axial compressionand tension capacity of the various composite

section components, known as componentcapacities.

3. Calculating the shear stud capacity based on theorientation of decking and number of studs per

set.

4. Determining the percentage of compositeaction (PSC) to use.

5. Calculation of the plastic neutral axis andcomposite section properties.

6. Calculation of the nominal composite moment

capacity.7. Multiplying by the global strength reduction

factor to give the design composite momentcapacity.

The full details of this procedure are given in

section 2.2 of (Clifton, 2002).

3.3 Design for Strength to Chosen International

Standards.

These international standards all follow the same

general design procedure, with certain differences

based on their specific development. These

differences to the New Zealand requirements are

given in section 3.3.1 to 3.3.3:

One significant difference between all three

international standards and New Zealand practice is

the use of global strength reduction factors in New

Zealand and the partial strength reduction factors inthe three international standards.

In New Zealand, the design shear capacity of the

shear stud (which does incorporate a partial

strength reduction factor, although this is 1.0 forstuds in positive moment regions) is matched to the

nominal component capacity when determining thenumber of shear studs required. The nominal

moment capacity is determined and converted todesign moment capacity through the use of a global

strength reduction factor.

In the three international standards, the nominal

capacity of the shear studs and internal componentsis converted to the design capacity through the use

of a partial strength reduction factor, which ismaterial specific. (Typical values, from Canadian

practice, are 0.8, 0.6, and 0.9 for the shear stud,concrete and steel beam components respectively).

The design capacity of shear stud is matched to the

design component capacity when determining the

number of shear studs required. The design moment

capacity of the composite section is then

determined directly.

This has important implications when determining

the number of shear studs required by eachprocedure, as discussed in sections 3.4 and 3.5.

3.3.1 Canadian Code (CSA S16-01: 2001).

The first difference is the load factors where (1.25)

is used for the dead load and (1.5) is for the live

load (Clause 7.2.4). This is the same as for(AS/NZS 1170.1:2002) and is not a critical

difference. The minimum shear connection ratio,


9/20

(PSCmin), allowed is 40% and for the componentcapacity there is a factor of (0.6) added to thenominal concrete capacity in compression (Nccc)

(Clause 17.9.3) and a factor of (0.9) for the nominaltensile steel capacity (Ntsc) (Clause 17.9.3). Finally,

the shear stud capacity for the secondary beam iscalculated according to the following equations:

uccscr Q'fA35.0q = 76mm decking. (3.1)

uccscr Q'fA61.0q = 38mm decking. (3.2)

where:

qr= shear stud capacity.

sc = factor that depends on the concrete type.(Clause 17.7.1). For normal weight concrete,

positive moment action it is 0.8.

Ac = area of concrete pullout surface (cone).

Qu = Ultimate shear capacity, which is the

following equation.

ccscscu E'fA5.0Q = (3.3)

While the primary beam equation is:

uscscc'cscscr fAEfA5.0q = (3.4)

3.3.2 British Standard (BS 5950:Part3:Section

3.1:1990).

The BS 5950:Part 3 differences with New Zealand

practice are as follows. First the load factors are(1.4) used for the dead load and (1.6) for the live

load. The minimum shear connection ratio (PSCmin)

is taken from Clause 5.5.2 where up to 10m, PSCminis 40%, while for 16m and above it is 100%. For

the spans between 10 to 16m, the followingrelationship between PSC and the span must be

satisfied:

( )4.0PSCbut

10

6LPSC

(3.5)

where:

PSC = Partial Composite Action.L = span in meters.

As for the component capacity, the concrete

compression stress is based on (0.45 fcu) fordetermining the design concrete capacity in

compression (Nccc) (Clause B.2.1). To convert thecharacteristic cube strength of concrete (fcu) to

cylinder strength (fc) a factor of (0.8) is used. Asfor the design tensile steel capacity (Ntsc) a factor of

(1/1.1) is used (Clause B.2.1). Finally, the designshear stud capacity is calculated according to the

following equation for both primary and secondarybeams, from Clause 5.4.3:

kQ8.0Q kp = (3.6)

where:

Qp = Shear Stud Design Capacity.

Qk = Characteristic resistance of the shearconnector, determined directly from experimental

testing and read from a table.0.8 = strength reduction factor.

k = reduction factor = 1.0 for both secondary andprimary beam configurations using 55mm

trapezoidal decking.

3.3.3Eurocode 4 (DD ENV 1994-1-1: 1994).

The Eurocode also has similar differences to the

British Standard. The load factors used are the same

as the British (1.4) for dead load and (1.6) live load,

while the shear connection ratio (PSC) is given byClause 6.1.2(2). Up to 10m, PSCmin is taken as

40%, while for 25m and above its taken as 100%.

For the spans between 10 to 25m the followingrelationship between PSC and span must be

satisfied:

L04.0PSC (3.7)

Partial strength reduction factors of (1/1.5) are used

on the calculation of design concrete capacity incompression (Nccc) (Clause 2.3.3.2(1)) and (1/1.1)

for the nominal tensile steel capacity (Ntsc) (Clause

2.3.3.2(1)). The shear stud capacity is given by the

minimum of the following equations for both

primary and secondary beams, from Clause 6.3.2.1:

v

2sc

u1r 4

d

f8.0kq

= (3.8)

v

cc2sc2r

E'fd29.0kq

= (3.9)

where:dsc = diameter of shear connector.

k= constant

3.3.4 Comparative results.

According to (Lin, et. al., 2001) the New Zealand

Standard clearly overestimated the shear studcapacity, as shown in Figure.3.1, where the

difference between the New Zealand Standard

(NZS 3404:1997), the Canadian Standard (CSAS16-01:2001, added by the author) and test results

from (Lin et. al., 2001) are clearly demonstrated,for 30 MPa concrete.

The Canadians recognised the unconservative

nature of their (CSA S16-01: 1984) shear stud

capacity equation for secondary beam

configurations, which prompted (Jayas and Hosain,


10/20

1988) to undertake research on the shear studcapacity. In (Jayas and Hosain, 1988) they

proposed the changes to the shear stud capacity

equation for secondary beams from that given forprimary beams (equation 3.4) to the new equation

3.1 and 3.2. This was followed by a full size test,details of which are published in (Jayas and Hosain,

1989). These changes were implemented in theCanadian Code (CSA S16-01: 2001), in Clause

17.7.2.3. As can be seen from Figure 3.1, these

equations predicted much lower shear studcapacities than the (NZS 3404:1997) provisions,

leading to the potential requirement for more studs

on a given beam.

The results of the comparative calculations are

presented in (Zaki, 2003) with details of 2 cases

given in Tables 3.1 and 3.2 below. These show the

difference between the various standards for deck

and strength reduction factors, shear stud capacity,number of studs required for the given beam and

shear connection ratio used. The designation

numbers 8/10/3.5 mean that the secondary beam

length was 10 m, primary beam length 8 m and theload acting on the tributary area 3.5 kPa.

Figure 3.1. Shear force-slip history for fc= 30 Mpa

(taken from Figure 5.13-unit 120/N30/19R/12F/02

(Lin, Y., et. al., (2001) with addition by author).

In Table 3.1.a, even though the New Zealand

determination for the primary beam shear studcapacity is higher than the rest, the number of studs

used is less than the number of studs required by

the British and Eurocode, as the PSC% used islower in this instance. Table 3.1.b shows that,

although the New Zealand procedure gives the

greatest shear stud capacity for this secondarybeam; the estimated number of studs is nearly

similar to the other standards. This is largely

determined by practical considerations of studspacing in a secondary beam supporting a profiled

steel deck. Looking at table 3.2.a, the New Zealand

shear stud capacity is still higher than the rest, but

the number of studs is higher, as it took a 100%PSC to make the primary beam work. Table 3.2.b

still has a high capacity, but the PSC is lower than

the British and Eurocode, as they are governed byClause 5.5.2 and clause 6.1.2(2) respectively intheir respective standards, (BS 5950:Part3: Section

3.1:1990) and (DD ENV 1994-1-1: 1994). Thesestate that the PSC is taken as a minimum of 40% till

10m increasing to 100% at 16m and 25mrespectively. Minimum PSC ratio is typically a

governing factor for secondary beam stud numbers.

These comparisons show that research to improve

the New Zealand composite design provisionsshould not just be limited to shear stud capacity but

should address the following:

1. The minimum value of PSC required.

2. Whether factors should be added to the

calculation of Nccc and Ntsc to be consistent

with the approach used by the other standards

that use partial material strength reduction

factors to match design shear stud capacity todesign internal action.

3.4.Initial Recommendation

At the conclusion to the comparative studies, the

recommendations on all these considerations

proposed were as follows:

1. For shear stud nominal shear capacity insecondary beams, use the new Canadian

equations equations 3.1, 3.2 withinterpolation as required and equation 3.3. Use

sc = 1.0 as for New Zealand practice.2. For shear stud nominal shear capacity in

primary beams, use equation 2.7 and determine

the potential for rib splitting from experimentaltesting.

3. For limits on PSC, change to the Eurocode

provision.4. Do not make changes to the calculation of

component capacities and the matching ofshear stud design capacity to component

nominal capacity.

The next stage of this project involved an

experimental testing programme to determine shear

stud nominal and design shear capacity when used

in a 55 mm deep trapezoidal deck concrete slab in a

primary and in a secondary beam configuration.

0

30

60

90

120

150

0 2 4 6 8 10 12 14 16 18 20

Mean slip (mm)

Shearforce/stud

(kN)

123 kN (NZS 3404)

83 kN (80% of max)

104 kN (max)

73.2 (CSA S16-01)

128 kN (Eq 2.6 )

58 kN (splitting)


11/20

Code Deck R.F. Strength R.F. Capacity No of Studs PSC%

NZ 1 1 94.1 42 70

Canada 1 0.8 89.5 28 75

UK 1 0.8 76 49 100

Eurocode 1 0.8 66.2 56 100

Table 3.1.a. Primary Beam 8/10/3.5.

Code Deck R.F. Strength R.F. Capacity No of Studs Practical No Studs* PSC%

NZ 1 1 94.1 24 34 50

Canada 0.54 0.8 68.5 24 34 40

UK 1 0.8 76 23 34 42

Eurocode 1 0.8 66.2 26 34 42

* Practical number of studs is dictated by 1 stud per deck rib.

Table 3.1.b. Secondary Beam 8/10/3.5

Code Deck R.F. Strength R.F. Capacity No of Studs PSC%

NZ 1 1 94.1 59 100

Canada 1 0.8 89.5 29 80

UK 1 0.8 76 39 80

Eurocode 1 0.8 66.2 51 90

Table 3.2.a. Primary Beam 8/15/5.

Code Deck R.F. Strength R.F. Capacity No of Studs Practical No Studs PSC%

NZ 1 1 94.1 42 49 50

Canada 0.54 0.8 68.5 28 49 40

UK 1 0.8 76 61 49 90

Eurocode 1 0.8 66.2 47 49 60

Table 3.2.b. Secondary Beam 8/15/5.

4. SHEAR STUD TEST SET-UP

4.1. Introduction

Because of the complexity of composite action, ashas already been discussed in sections 2.3 and 2.4,

the strength and ductility of shear studs are always

determined experimentally. The details associatedwith the shear stud testing conducted at the

University of Auckland are reported in this section.

The test set-up adopted that previously used by(Butterworth, 2000), with changes made to the

dimensions of the test units, the shape of the corbel,

and the position of the test units on the test rig. Theshear force was applied by blocking one side of the

test unit and adding a push force on the other side.The push force and the reaction force were

positioned collinear with the steel-slab interface,

such that the position of the test unit on the test righad no influence on test results.

The associated details of the test set-up including

test rig, test units, instrumentation, loading

procedure, and corbel design are discussed insections 4.2 to 4.7.

4.2. Shear Stud Testing Rig

The test rig with mounted unit is shown in Figure.

4.1. The frame had three components; the base

frame, head frame and foot frame. The baseframe was bolted to the strong floor of the test hall.

A 1000 kN jack was horizontally bolted to thehead frame. Centre lines of the push force and the

reaction force were regulated by steel bearings,which were positioned collinear with the steel-slab

interface and the horizontal stiffener in the head

and foot frames.

Test units were laid on packers with the slabpositioned on top and the steel section positioned

underneath, representing the real situation.

Thickness of the packers was carefully adjusted tomake the steel-slab interface collinear with the the

centre line of the bearings (line of thrust). The steel

corbel of the test units was positioned close to jack

and the foot frame restrained the concrete corbel.


12/20

Figure 4.1. Test rig with mounted test unit

(Butterworth 2000)

Slide rollers were placed between the packer and

the lower surface of the steel beam, in order for the

slab to freely slide during testing.

4.3 Shear Stud Test Units

4.3.1 General

The push specimens are referred to here as shear

stud test units. The test unit was composed of three

main components steel beam, composite slab and

studs. A typical shear stud test unit is shown in

figure. 4.2, 4.3 and 4.4, for both secondary andprimary beams, with each test unit having small

changes from that shown

The design of shear stud test units adopted in thisresearch closely emulated the design used

previously (Butterworth, 2000), although therewere some changes, these being:

1. As a conventional profiled steel decking wasadopted in this study, the total thickness of the

slab was much less than that previously tested,so studs with a length complying with (NZS

3404:1997) Clause 13.3.2.2.1 (d)-ie hsc hr +

40 mm were required. In this case hsc 95mmwas required.

2. There was no concrete haunch over the steelbeam, which in previous testing had a width

the same as the steel beam top flange.

3. Especially for the secondary beam, stud

spacing was controlled by the decking profile.

4. Because of the reduced slab thickness, the

section of the corbel became more critical, sothe corbel required redesign.

5. Nominal mesh reinforcement complying with

(NZS 3404:1997) was used.

4.3.2 List of Shear Stud Test Units

All test units used in this study are listed in Table

4.1. Key results are presented in section 5, with full

details in (Zaki, 2003).

S is for the secondary beams and P is for theprimary beams. All the secondary beam slabs were

1105mm by 1000mm with 120mm depth while the

primary beams were 1240 mm by 1000 mm with

120mm depth no matter how many studs were

tested.

For the secondary beams, Units S1, S2 and S3 are

identical in the number of studs and concrete

strength, while S4 has the same number of studs but

different concrete strength. S5 has double thenumber of studs at the staggered position, while

S6 has the same number of studs as the first 4 with

a concrete strength of 25 MPa but the studs

positioned in the unfavourable position.

For the primary beams P2, P3 and P5 are identical

while P1 is the same but with different concretestrength and finally, P4 has double the number of

studs at double the spacing compared to P2, P3 andP4. These differences were intended to generate a

good representation of shear stud capacity and test

the influence of key parameters, but not change toomany variables to give meaningful results. Theshear connection devices used in this study wereheaded studs with a shank diameter of 19 mm and a

before-weld length of 106mm, clean beam and 110mm through deck. These give an installed length of

99-102mm in each instance.

4.4 Loading procedure

The test protocol was based on the Eurocode 4

(1994) recommendation:

1. Load the test unit to 40% of the expectednominal capacity (failure load) of the stud

group.2. Cycle the load 25 times between 5% and 40%

of the expected nominal capacity.

3. Continue monotonic loading from 40%, such

that 100% of the expected normal capacity isreached in under 15 minutes.

4. Measure longitudinal interface slip

continuously during loading, at least until theload has dropped to 20% below maximum.

80 mm packers

310UC158-HD bolts to strong floor

610UB101295

"foot" frame bearing

610UB101loading plate

"head" frame

horizontal stiffer

bearingload cell

1000 kN jack

2410


13/20

5. Measure transverse separation between thesteel beam and the composite slab close to thegroup of shear studs.

Summary Details of Test Units

No. Code Slab

Dimension

fc No of

Studs

Comments

1 S1 1105x1000x120

25 3 Favourable

2 S2 1105x1000x

120

25 3 Favourable

3 S3 1105x1000x

120

25 3 Favourable

4 S4 1105x1000x

120

30 3 Favourable

5 S5 1105x1000x

120

25 6 Double

Staggered

6 S6 1105x1000x

120

25 3 Unfavourable

7 P1 1240x1000x120

30 4 Single Row

8 P2 1240x1000x120

25 4 Single Row

9 P3 1240x1000x

120

25 4 Single Row

10 P4 1240x1000x120

25 4 Double Rowat Double the

Spacing

11 P5 1240x1000x120

25 4 Single Row

Table 4.1 Shear Stud Test Units.

Figure 4.2 Unit 2 ready for concrete casting.

Figure 4.3 Unit P2 ready for concrete casting.

Figure 4.4 Test Unit S1 ready for testing.

5. SHEAR STUD TEST RESULTS

Reporting of the final shear force-slip history

adopted the format shown in figure 5.1, while

figure 5.2 is from the test results of unit S4.

Figure 5.1 Typical shear force-slip history

diagram.

Figure 5.2 Unit S4 test results.

The following observations were recorded during

testing of the units. For the secondary beams, thecracks formed transverse to the beam especially at

the first and third trough. The crack adjacent to the

third trough typically was the first crack to form butdid not affect the results. The longitudinal crack,

which is also shown in figure 5.3, typically formedat or near the peak shear stud strength achieved, it

also shows an example of the crack formationtypical for a secondary beam. Delamination of the

slab also occurred, figure 5.4, as well as an

interesting phenomenon named the biting

Phenomenon of the decking into the steel beam,

see figure 5.5. Another interesting aspect of testingwas the wedging of the concrete between the

0

10

20

30

40

50

60

70

80

90

100

110

120

0 2 4 6 8 10 12

Mean Slip (mm)

Shearforce/stud

(kN)

100.21 kN (NZS 3404)

86.6 kN (4.1)

106 kN (max) and L1

84.4 kN (80% of max)L2

L3

Mean slip (mm)

Shearforce/stud

(kN)

Starting from end of first 25 cycles

(qr)n (NZS 3404) (=1.0)

Revised qr

Q max (kN)

80 % of max


14/20

shear stud and the decking, see figure 5.6. Theseobservations were common for all secondary beamtest units.

As for the primary beam, longitudinal cracks

parallel to the beam occurred with curving of theslab due to the splitting of the concrete pushing

against the stud. Figure 5.7 demonstrates thecurving while figure 5.8 shows the typical

longitudinal cracks. Refer to (Zaki, 2003), Chapter

6 for detailed observations of the test units.

Figure 5.3 Secondary beam cracks.

Figure 5.4 Delamination of the slab.

Figure 5.5 Biting Phenomenon.

Figure 5.6 Wedging of the concrete.

Figure 5.7 Curving of the slab in the PrimaryBeam.

Figure. 5.8 Primary Beam cracks.

6. SHEAR STUD TEST RESULTS-

DISCUSSION

6.1. Introduction.

This section discusses the shear stud failure modes

and gives final recommendations for shear studcapacity and composite design provisions.

6.2 Failure Mode.

The main modes of failure described in theliterature are concrete pull out failure (stud

embedment failure), concrete crushing failure,shear stud fracture, and concrete splitting failure.

Of these, the major cause of failure in this series oftests was concrete splitting failure, with a case of

concrete crushing and splitting (Unit P4) and two

cases of a new mode of failure, which we arecalling Rolling Failure. This principally affected


15/20

unit S2, which had Rib Rolling Failure and unitP3, which failed due to a construction defect. Thefollowing section talks about the reason why

splitting and not cone pullout failure was the maincause of failure, even though the concrete cone

shape was present in nearly all the units tested.

6.2.1. Why Cone Pull Out Failure Was Not theMain Failure Mode.

Our initial recommendations and calculations used

to develop the test specimens were based on the

Canadian study by (Jayas and Hosain, 1988 &

1989) for the secondary beam, which in turn basedtheir equations on the reported embedment

failure. For the primary beams, initial

recommendations were based upon the longitudinal

splitting concept from (Oehlers and Bradford,

1995). However, embedment failure was not the

observed failure mode for any of the specimenstested. Therefore a new set of recommendations has

to be made, to better represent the failure mode that

occurred in this series of tests. Embedment failure

is the form of failure that always has a concretecone shape surrounding the shear studs where the

concrete around the cone gets separated from the

slab causing a drop in strength. It is not

accompanied by a longitudinal splitting crack. In

the tests conducted, although the cone shape waspresent in all the secondary beam units after

dissecting the unit, it was not the observed failuremode that limited the stud shear strength. As

previously reported, that was longitudinal splittingin all cases except for units S2 and P3 which failed

prematurely and unit P4 which failed by combined

crushing and splitting.

Further support for this can be seen from a study ofthe parameters that govern embedment strength.

According to (Oehlers and Bradford, 1995), theembedment strength is dependent on the ratio of the

stud height to the stud diameter (see figure 6.1)

while the strength can be derived from thefollowing equation.

( )beammaxhemb

DKD = (6.1)

Where, Demb is the shear strength that allows

variation in shear embedment strength, Kh can be

derived from line A-B-C in figure 6.1 and(Dmax)beam can be derived from equation 6.2.

Equation 6.2 represents the strength of a shear stud

in a solid concrete slab. It is also the original

equation that (NZS 3404:1997) Equation 13.3.2.1 is

derived from, as it incorporates both the concretecrushing and shear stud fracture resistance into one

equation. It is determined from a curve fit of testdata on push-off tests in solid slabs, with allowance

made for the test rig setup on the capacity. Details

Figure 6.1 Shear Embedment Strength Relationship

to find Kh (Oehlers and Bradford, 1995).

are given in section (2.4.6.3) of (Oehlers and

Bradford, 1995):

( )4.0

s

c35.0c

65.0uscbeammax E

EffA

n

1.13.4D

= (6.2)

where n is the total number of studs in the test

sample. Using the values from experimental tests

including the actual measured material properties,

the Shear Embedment Strength is calculated and isshown in figures. 6.2 and 6.3 for the secondary and

the primary beams, respectively.

The values of the shear embedment strength (Demb)means that, if no other form of failure occurs before

Demb is reached, then embedment failure will limit

the available strength. However, with these studs,slab crushing and shear stud fracture would occur

before embedment failure. In fact (Dmax)beam ,which

incorporates both the concrete crushing and shearstud fracture, clearly was also greater than the testresults, demonstrating that the failure mode was notthe concrete crushing or fracture of the studs at the

maximum stud strength. Therefore, according to the

Figure 6.2 Comparison between the test results and

the embedment, solid concrete failure modes for the

secondary beams.

results in Chapter 6 in (Zaki, 2003) and the

comparison shown above in figures 6.2 and 6.3

neither did shear embedment, concrete crushing or

0

10

20

30

40

50

60

70

80

90

100

110

120

130

140

0 1 2 3 4 5 6 7 8 9 10 11 12

Mean Displacement (mm)

Shearforce/stud

(kN)

(Dmax)beam = 123 kN

Demb = 135.3kN


16/20

Figure 6.3 Comparison between the test results and

the embedment, solid concrete failure modes for the

primary beams.

stud fracturing initiate the failure, but longitudinal

splitting did. It should be noted that stud fracture

was observed in units S6 and P4, however this didnot occur until well after the peak load was reached

and so was not the limiting factor in terms of stud

design capacity. These results show that revision to

(NZS 3404:1997) Equation 13.3.2.1 is needed.

6.3 Proposed Revision of NZS 3404 Cl 13.3.2.1

As mentioned in Section 2.4.3, equation 13.3.2.1 in

(NZS 3404:1997) has been amended in 2001 fromthe old format of:

dcscuccscdcr Af)Ef(A5.0q = (6.3)

Into:

dcscuccscdcr Af8.0)Ef(A4.0q = (6.4)

where the first part is for concrete crushing while

the second is for shear stud fracture. According toequation 6.3, from which this simplified equation is

derived, the present equation 6.4 underestimates thetrue value of the concrete crushing. It is

recommended that the equation 6.5, initiallyproposed by (Lin, et. al., 2001) as his equation 2.15,

and referenced in (Clifton, 2002), be used:

dcscuccscdcr Af8.0)Ef(A5.0q = (6.5)

This equation gives a value that is consistent withthe test results and with the original equation putforth by (Oehlers and Bradford, 1995), equation

6.2. Given that this is for a solid slab, dc=1.0 andthis term can be omitted giving equation 6.6 as the

proposed replacement to NZS 3404 Equation13.3.2.1:

scuccscr Af8.0)Ef(A5.0q = (6.6)

Comparing equations 6.4, and 6.6 with 6.2 we getthe following results, using the actual material

properties of the test:

Ec = 24.8 GPa Asc = 283.4 mm2

fu = 516 MPa fc = 24.3 MPa

Eq. 6.4 qr= 88 kN 117 kN

Eq. 6.6 qr= 110 kN 117 kNEquation 6.2 (Dmax)beam = 112 kN

It is clear that equation 6.6 shows closer results toequation 6.2 than equation 6.4. Therefore, it is

recommended that the present equation 13.3.2.1 in

(NZS 3404:1997) should be amended to equation

6.6, for welded studs in a solid concrete slab.

6.4 Making Allowance for Concrete Splitting in

Interior Beams.

The experimental tests showed that concrete

longitudinal splitting failure governed all tests

involving a single row of studs, either in a straightline or staggered. The equation for calculating thesplitting strength, Vsh, of a stud within a concrete

rib of effective width, bce, is given by (Oehlers and

Bradford, 1995). Equation 2.12, which can betransformed into equation 6.8:

+

=

c

2

c

sc

ce

sc

cecscsplit

h

8.1

h

d8.11

b

d1

1

b'fd54.0V

(6.8)

In this equation, bce is the effective width, dsc is the

stud diameter, and hc the effective haunch height forsplitting, taken as the minimum of (5.4dsc,to), whereto is the slab height.

In this instance, the actual splitting shear strength

has been determined by experimental testing, using

the process described on page 28, equation 55.26 of

(Hyland, C., et. al. 2001).

That equation is used to back-calculate and find the

effective width, bce, for the interior secondary andprimary beam configurations.

For the secondary beam, this results in the effective

width using the results from our test. Resulting in

the effective width being 440mm, for the presenttest conditions of the slab height being greater or

equal to 120mm, stud height greater or equal to 4xthe stud diameter and stud diameter of 16 to 22mm.

This effective width happens to be close to four

times the stud height used in the tests. (Lin, et. al.

2001) recorded that the concrete cone shape slope

angle was around 24 to 29, while the slope anglefor this test with the effective width of 440 mm was

0

10

20

3040

50

60

70

80

90

100

110

120

130

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21

Mean Slip (mm)

S

hearforce/stud

(kN)

P2

P4

P5

Demb = 125.4kN

(Dmax)beam = 114 kN


17/20

24.4. This indicated that there may be arelationship between effective rib width and studheight. Therefore, a recommendation future testing

is to find if such a relationship between the stud

height and effective rib width exists and itsinfluence on the shear stud capacity.

Taking the effective width to be 440mm tocalculate the longitudinal splitting capacity, for the

studs in the favourable position, another factor isrecommended to be added to this equation to make

it consistent with the test results. This factor is the

stud-positioning factor (sp)taken as:

Favourable = 1.0

Unfavourable = 0.94

Double staggered = 0.84

This factor is the position of the stud in the

secondary beam, whether it is in the favourable,

unfavourable or double staggered.

Therefore, the recommended splitting equation tobe used for shear stud capacity estimation in the

secondary beam configuration is:

+

=

c

2

c

sc

ce

sc

cecscspsplit

h

8.1

h

d8.11

b

d1

1

b'fd54.0V

(6.9)

The following values and limits are required withthis equation:

bce = 440 mmto 120 mm

hsc 95 mmdsc from 16 mm to 22 mm.

hsc/dsc 4.32

Profile is 55mm trapezoidal deck. Where, bce =effective width to use for interior secondary beam

(ie runs over decking).

For the primary beam, splitting was clearly themain cause of failure for all single rows of studs.

Therefore the recommended equation 6.8 is also

used for the primary beam, back-calculating again

to find the effective rib width. This is also 440mmfor the present test conditions of the slab heightbeing greater or equal to 120mm, stud height of

greater or equal to 4 stud diameters, stud diameterof 16 to 22mm, and the deck fastened onto the

beam flange within 10mm at the flange edge.

Figure 6.4 demonstrates that edge distance. Thatequation is applicable to a single row of studs.

It was noted in unit P1, which had 30 MPa concrete

instead of 25 MPa, that the strength was 98 kN

which was 1kN more than the equivalent specimenusing 25 MPa concrete. The unit S4, which was thesecondary beam with 30 MPa strength, was 106 kN

which was approximately the same as the predictedvalue from equation 6.9. This difference in

strength could be due to the different 30 MPa

Figure 6.4 Edge distance of decking on Primary

Beam.

concrete batches used that were hand mixed, and

might have had different compressive strengthsfrom the cylinder test values. It is an aspect for

which future testing is recommended.

Unit P4, with the double row of studs at doublespacing, clearly had a combination of crushing and

splitting failure, as seen in figure. 6.5. The splitting

crack is clearly seen, but is deflected by the stud,

also the crushing of the concrete was also present.

This is also consistent with the parameterscontrolling splitting, which show that a double stud

row is not as susceptible to splitting is a single studrow.

Therefore, for unit P4 which had double the studs at

double the spacing, equation 6.6 is to be used to

better represent the failure mode and it also givesthe best estimate for the shear stud capacity. Where

equation 6.6 for crushing was 110 kN while our testresult was 104 kN, therefore equation 6.6 is slightly

Figure 6.5 Crushing and splitting failure in unit P4.

overestimating the value, which can be from the

minimal splitting cracks that did occur but were

deflected by the stud. A Double Stud safety

factor of (0.9) is recommended until future tests are

conducted to find the actual strength and whether afactor is needed or not.

10 mm


18/20

6.5 Extending Concrete Splitting Provision to

Spandrel Beams

6.5.1 Secondary Beam

For the spandrel secondary beam, it is

recommended to use equation 6.8 but taking theeffective width as:

ofce bbb += (6.10)

This is likely to be conservative but to an unknown

extent. Because the decking is not continuous over

the beam, its role in confining the concrete withinthe splitting zone of the stud (see figure. 2.4 (a))

will be limited and is ignored in this

recommendation. Also, using the post-splitting

reinforcement shown below is recommended:

Figure 6.6 Secondary Spandrel Beam.

The DH12 @ 300 mm was calculated using thepost splitting provisions of (Oehlers and Bradford,1995). Finally, (NZS 3404:1997), Cl. 13.4.1.3(3)

states that if the decking runs across the spandrel by

greater or equal to 550mm then the spandrel beamis to be treated as an internal secondary beam. Thatis consistent with the experimental tests.

6.5.2. Primary Beam

Similarly with the primary spandrel beam, excepttaking the effective width as follows:

oferiorint,ce

ce b2

b

2

bb ++= (6.11)

Also, use of the post-splitting reinforcement as

shown below is recommended:

Figure 6.7 Primary Spandrel Beam.

The DH12 @ 300 mm was calculated using the

post splitting equation 2.19.

6.6 Suppression of Rib Rolling Fracture

Unit S2 had what is called Rib Rolling Fracture

which is when the mesh depth is greater than theintended designed depth, setting the mesh

reinforcement below the tension zone in the slab,

causing the slab to tilt resulting in premature failure

of the unit. It is called the Rib Rolling Fracture

because the action of the forces in the slab cause the

slab to roll caused by the rolling moment

forces acting in the slab. This is demonstrated in

figure 6.8.

Figure 6.8 Rolling Fracture

To explain this failure mathematically, the concrete

elastic moment and the crack moment were

calculated for both unit S1 and S2 to give a

comparison between the two failure modes.

First, for unit S2 it had the following

characteristics: the transverse crack that caused the

failure L5 occurred at a shear force/stud of 58 kN.

The mesh depth at the crack was 50mm, making the

centre of the reinforcement to the top of the slab

height 56mm. Basing the calculations on that we

get the following:

Total area of mesh and reinforcement = 618mm2

Elastic neutral axis = 34.75mm from the top of the

slab.

Resulting Elastic Concrete Moment before cracking

is thus 2.23kNm

While the post-splitting moment, calculated in

accordance with figure 6.9, using the conventional

reinforced concrete theory, is 3.953 kNm

This indicate that the concrete will develop a crack,which actually happened, but when the result of this

unit is compared to unit S1 which had the following

characteristics, the transverse crack that caused the

failure L4 occurred at a shear force/stud of 92 kN.

The mesh depth at the crack was 40 mm making the

centre of the reinforcement to the top of the slab

height 46 mm. Basing the calculations on that we

get the following:

DH 12 @ 300 mm

bobf

50

DH12

bce,interior/2 bobf/2

DH 12 @ 300 mm

DH12

< 220 mm

a. Mesh Depth at 35 mm; Unit S1

.

ActuatorDirection

35 mm

Mesh> 35mm

b. Mesh Depth at 50 mm; Unit S2

ActuatorDirection

Neutral Axis


19/20

Total area of mesh and reinforcement = 618 mm2

Elastic neutral axis = 33.8 mm from the top of the

slab.

Resulting Elastic Concrete Moment before cracking

then becomes 2.11 kNm

While the Moment of the concrete after cracking is

6.36 kNm

Now this clearly shows that for unit S1 the crack

will also appear, but will require a larger moment to

cause it. The calculation of the Ultimate Moment

Capacity was also conducted for further proof.

Yield stress fy was found to be 448.5 MPa.

T = Asfy = 618x0.4485 = 277.2 kN

04.13b'f85.0

Ta

c

== mm

Unit S2: M=277.2(65-56-(13.04/2))= 0.687 kNm

Unit S1: M=277.2(65-46-(13.04/2))= 3.459 kNm

Figure 6.9 Forces actions after cracking.

This shows that for unit S2 there is not sufficient

moment generated to withstand the forces acting on

the slab, while for unit S1 there was sufficient

moment to withstand the force and generate goodstrength. Figure 6.9 shows the ultimate moment

actions.

This mode of failure can be prevented by ensuring

that the mesh depth is not greater than 35mm fromthe top of the slab.

6.7 Recommended Changes to the Composite

Beam Design Procedure.

6.7.1 Minimum Shear Connection Ratio (ie.

Partial Composite Action)

From a comparison of all the provisions it isrealistic to amend the Partial Composite Action

limit to that given by the Eurocode 4 (Clause6.1.2(2)), to give a more realistic shear connection

ratio (PSC). The New Zealand provisions state that

the PSC is taken as a minimum of 50 %, while theEurocode stipulates a minimum value of PSC br

taken as 40 % up to 10 meters, increasing to 100%at 25 meters.

6.7.2 Correlation factor for Component

Capacity Equations.

This is being presented for completeness although it

is not intended to make changes to the design

procedure.

Table 6.2 shows the influence on shear stud numberdetermination if this is determined by matching

design capacity of shear stud to design capacity of

the critical component action, as is done with the

three international procedures studies. Where thecompression generated within the concrete slab

governs, adopting this approach would result in areduction of 0.83x the exact number of shear studs

required by current New Zealand practice. Where

the tension generated within the steel beam

governs, adopting this approach would increase the

number of shear studs required by current NewZealand practice.

The net result across all application would be a

benefit in terms of reducing the number of shear

studs required, however the complexity of changingdesign process outweighs this benefit. It is therefore

preferable to leave the currant practice of matchingdesign shear stud capacity, incorporating a partial

strength reduction factor = 1.0 in positive momentregions, to the nominal component capacity in

order to determine the number of shear studs

required.

Canada UK Eurocode

NewZealand

Shear

Stud

0.8 1/1.25 1/1.25 1.0

Concrete 0.6 1/1.5 1/1.5 1.25/1.5

= 0.83

Steel 0.9 1/1.1 1/1.1 1.25/1.1= 1.13

Table 6.2. Strength Reduction Factor Comparison.

7. RECOMMENDATIONS FOR FURTHER

RESEARCH.

These are as follows:

(i) To determine the influence of rib endsupport conditions on suppressing the rib

rolling failure shown in the secondary

beam tests, by adding stiffening ribs to the

test specimens to represent continuity of

slab beyond an internal secondary beam.(ii) To determine the influence of shear stud

height/diameter ratio on the load-deflection characteristics and design

capacity reached (i.e. provide experimental

data to replace the dotted line AB in

Figure. 6.1).

Actuator Direction

Crack

a

T

Cjd

Compression

Block

Mesh


20/20

(iii) To determine the influence of mesh heighton spandrel secondary beam capacity.

(iv) To determine the splitting capacity of

spandrel beams and so allow the design

recommendations of section 6.5 to be

updated.

REFERENCES

Bird, G.D. and Klemick, M.P.; HiBond Design

Wizard for Composite Design of the HiBond

Flooring System, Version 1.0; Dimond, Auckland,

New Zealand, 2002.

British Standard (BS 5950 (1990)). Structural use

of steelwork in building, Part 3: Design in

composite construction, Section 3.1: Code of

practice for design of simple and continuous

composite beams, British Standards Institution,London.

Butterworth, J. (2000). Push-off tests on shear

studs used with a deep ribbed deck, Auckland

UniServices Limited, University of Auckland,Report No. 8258.00, (pp. 25).

Canadian Standard (CSA S16-01: 1984). Limit

States Design of Steel Structures, CanadianStandard Association, Ontario, Canada.

Canadian Standard (CSA S16-01: 2001). Limit

States Design of Steel Structures, Canadian

Standard Association, Ontario, Canada.

Clifton, C. (1998). Reminders on good detailingand construction practice", Steel Design andConstruction Bulletin, HERA, No. 42, (pp. 8).

Clifton, C. (2002). Notes prepared for Composite

Steel Design and Construction Seminar, HERAReport R4-113, 2002.

Eurocode 4: ENV 1994-1-1:1994, (1994). Designof Composite Steel and Concrete Structures Parts

1-1 General Rules and Rules for Buildings, BSI

Standards, London, pp. 143-148.

Grant, J., Fisher, J.W. and Slutter, R.G., (1977).

Composite beams with formed metal deck,Engineering Journal of the American Institute of

Steel Construction, Vol. 14, 1st

quarter, (pp. 24-42).

Hawkins, N.M. and Mitchell, D., (1984). Seismic

Response of Composite Shear Connections,Journal of Structural Engineering, ASCE, Vol.

110, No. 9 (pp. 2120-2136).

Hyland, C., Clifton, C., Butterworth, J,. andStickland, S., (2000). Shear studs in composite

construction: concepts in determining the shear studdesign shear capacity, Steel Design andConstruction Bulletin, HERA, No.55, (pp. 18-28).

Hyland, C., Clifton, C., Butterworth, J,. and

Stickland, S., (2001). Composite Down-Stand

Steel Beam Behaviour with a Profiled Deep-Deck

Slab, Proceeding of Australasian StructuralEngineering Conference, 2001.

Jayas, B. S. and Hosain, M. U. (1988). Behaviour

of headed stud in composite beams: push-out tests,Canadian Journal of Civil Engineering, Vol. 15,

No. 2, (pp. 240-253).

Jayas, B. S. and Hosain, M. U. (1989). Behaviourof headed stud in composite beams: full-size tests,Canadian Journal of Civil Engineering, Vol. 16,

No. 2, (pp. 712-724).

Johnson, R. P. and Anderson, D. (1993).Designers Handbook to Eurocode 4 part

1.1:Design of Composite Steel and Concrete

Structures. Thomas Telford, London.

Lin, Y., Butterworth, J., and Ingham, J. (2001).

Capacity of Shear Studs Embedded in Lightweight

polystyrene Concrete Diaphragms, School of

Engineering, Civil and Resource Engineering,

University of Auckland, Report No. 607, December2001.

New Zealand Standard (NZS 3101: Part 2:1995).Concrete Structures Standard Part1 The Design

of Concrete Structures, Standards New Zealand,

Wellington, pp. 209-212.

New Zealand Standard (NZS 3404: Part 1:1997).

Chapter 13, Design of Composite Members and

Structures, Standards New Zealand, Wellington.

Oehlers, D. J. and Bradford, M. A. (1995).Composite Steel and Concrete Structural Members,Elsevier Science Ltd, Oxford, U.K.

Ollgaard, J. G., Slutter, R. G., and Fisher, J.W.(1971). Shear strength of stud connections in

lightweight and normal-density concrete,Engineering Journal, American Institute of Steel

Construction, Vol. 8, (pp. 55-640.

Slutter, R. G. and Driscoll, G. C. (1965). Flexural

strengths of steel concrete composite beams,

Proc. ASCE, J. structural Div., Vol. 91, No. ST2,

April 1965, pp. 55-64.

Zaki, R.J. (2003) Shear Stud Capacity in Profiled

Steel Decks, MEngSt, Project Report, University

of Auckland, September 2003.

Documents

Shear Stud Cap in Steel Decks