paper: hicks
Strength and ductility of headed stud connectors
welded in modern profiled steel sheeting
Synopsis
Push tests in Australia have suggested that studs welded in
modern trapezoidal decks possess lower resistances and slip
capacities than expected. This paper presents the results from
a comprehensive research programme consisting of full-scale
composite beam and companion push tests. The beams
exhibited excellent ductility, and the slip capacity at the shear
connection far surpassed the levels considered obtainable by
current British and European Standards. Comparisons of
load–slip performance show that any brittleness exhibited in
push tests is as a result of a deficiency in the standard push
specimen rather than the shear connection. For single studs,
the beam tests demonstrate that the current UK design
practice of using the shoulder height of the deck in BS 5950
to be the most appropriate. Conversely, the performance of
pairs of studs in the current research appears to be adversely
affected by longitudinal connector spacing. Interim design
guidance is proposed for studs in this arrangement.
equations that are multiplied by the design resistance of studs
embedded in solid concrete slabs. The reduction factor formulae in BS 5950-3.1 and Eurocode 4 have been empirically
derived from push tests, which have been revised and modified
as new deck profiles have been developed. It should also be
noted, however, that no advice is given in either of these
Standards on how to adapt the standard specimen in Fig 1(a)
when decking is present; this has led to a wide variation in specimen geometry for push tests with decking which, in part, may
be a reason for the high scatter in results observed in previous
calibration studies7,8. The reduction factor formula given by BS
5950-3.1 is as follows:
k = 0.85/√nr(b0/hp){(hsc/hp) – 1}
...(1)
but k ≤ 1.0 for nr = 1 and k ≤ 0.8 for nr = 2
where nr is the number of studs per rib, b0 is the breadth of
the concrete rib (which should be reduced to 2e when the studs
are welded in the unfavourable position), hp is the height of the
steel deck rib and hsc is the height of the stud (see Fig 2).
Introduction
Historically, the performance of shear connectors has been
established from small-scale specimens of the type shown in Fig
1(a). By applying a load to the top end of the steel-section, the
load–slip behaviour of the connectors can be determined. This
type of specimen is known as a ‘push test specimen’ in BS 59503.1: 19901 and BS EN 1994-1-1: 20042 (Eurocode 4) and, apart
from slight variations in its geometry, has hardly changed since
its inception in the early 1930s3.
a)
b)
According to Eurocode 4, if three nominally identical tests are
carried out, and the deviation of any individual result from the
mean value does not exceed 10%, the characteristic resistance
of a shear connector PRk is defined as 0.9 times the minimum
failure load per stud (see Fig 1(b)). The ductility of a shear
connector is measured by the slip capacity δu, which is defined
in Eurocode 4 as the slip value at the point where the characteristic resistance of the connector intersects the falling branch
of the load-slip curve (see Fig 1(b)). The characteristic slip capacity δuk is taken as 0.9 times the minimum test value of δu.
Alternatively, the characteristic properties of a shear connector
can be determined by a statistical evaluation of all of the results
according to BS EN 1990: 20024.
The rules for partial shear connection given in BS 5950-3.1
and Eurocode 4 are based on extensive numerical analyses,
which considered composite beams over a wide range of spans
and section types5,6. To enable the designer to assume that all
of the studs are equally loaded at the ultimate limit state (i.e.,
the shear connection is fully plastic), the minimum degree of
shear connection in these Standards is based on an assumed
ductility limit. In Eurocode 4, a connector may be taken as
ductile if the characteristic slip capacity δuk is at least 6mm.
When stud connectors are welded within the ribs of profiled
steel decking, their resistance is reduced when compared to
studs embedded in solid concrete slabs. To account for this
reduction, BS 5950-3.1 and Eurocode 4 provide reduction factor
32|The Structural Engineer – 15 May 2007
9
In 1993 the SCI wrote an advisory desk article to clarify the
definition of deck height to be used when calculating the reduction factor for modern trapezoidal decks that possess a shallow
central stiffening rib on their crests (see Fig 2). The article
stated that, for trapezoidal decks of the form shown in Fig 2,
only the height to the shoulder hp,n needed to be considered
provided that: the stud projects at least 35mm above shoulder
height (i.e. hsc – hp,n ≥ 35mm); and the re-entrant portion is not
more than 15mm in height (i.e. hp,g – hp,n ≤ 15mm) and 55mm
in width. This gave less onerous reduction factors than those
based on the overall deck height hp,g.
There has been much publicity of push tests in Australia that
have suggested considerably lower resistances and slip capacities than expected for stud connectors welded in the troughs
of trapezoidal decks10,11. Proprietary solutions, such as special
reinforcement, which are now required in certain situations by
the Australian Standard12, are proposed to overcome these problems.
An obvious question regarding shear connectors in push
specimens is whether their behaviour is comparable to that
which would occur in a full-scale beam. While many full-scale
beam tests have been undertaken in the past, there was insufficient instrumentation to determine the load–slip performance of the shear connectors in situ. As a consequence of this,
although the measured end slips were in excess of the required
ductility limits, it was impossible to know whether the stud
performance was much better in beam specimens or, as implied
by Australian push test results, whether the outer studs were
on the falling branch of the load–slip curve. Moreover, the
degree of shear connection provided in these tests was not low
enough to provoke a brittle failure had the studs possessed
insufficient slip capacity. In an attempt to answer these questions, full-scale composite beam and companion push specimens were constructed and tested to failure at Cambridge
Stephen Hicks
BEng(Hons),
PhD(Cantab)
Manager, Building
Engineering, The Steel
Construction Institute,
Silwood Park, Ascot,
Berks
Received: 08/06
Modified: 01/07
Accepted: 02/07
Keywords: Beams,
Composite construction,
Shear connectors, Studs,
Decks, BS 5950,
Eurocode 4, Testing,
Behaviour
© Stephen Hicks
Your comments on this
paper are welcome and
will be published online
as Correspondence.
Please send your
contribution to the Editor
within 3 months
Fig 1. (left)
Eurocode 4 (a)
standard push test
specimen and (b)
determination of
characteristic
resistance and slip
capacity from
load–slip curve
Fig 2. (below)
Dimensions of
profiled steel
decking and studs in
the (a) central (b)
favourable and
(c) unfavourable
position
paper: hicks
strengths fym presented in Table 1 were evaluated. As well as
being used in the predictions of moment resistance, the stressstrain behaviour of the material was used in the back-analysis
of the strain gauge readings.
The target mean compressive cube strength of the concrete
was 25N/mm², which was achieved by specifying a normal
weight C16/20 mix; this strength was chosen in an attempt to
provide the lowest degree of shear connection that is permitted
by the current Standards to obtain evidence of slip capacity. A
summary of the measured properties on the day of the beam
test is presented Table 2.
The shear connectors consisted of 19 × 100mm long headed
studs (95mm length-as-welded). To determine the mechanical
properties of the stud material, three coupons were machined
from finished studs. From tensile tests on these coupons, the
mean ultimate tensile strength of the stud material was found
to be fum = 513.5N/mm² with an elongation of between 15 to 16%.
University Engineering Department. The remainder of this
paper describes this research and its implications on design.
Fig 3. 10m span
composite beam
specimen
Composite beam specimen 1
The composite beam was simply-supported over a span of 10m,
and consisted of a 305 × 165 × 46kg/m UB (see Fig 3). To provide
the most unfavourable loading to the studs, the beam was
propped at third-points at the wet concrete stage so that the full
self-weight load was applied to the shear connection once the
props were removed. As well as pre-loading the studs, this
construction also ensured that the effects of ponding were
minimised to enable a constant slab thickness to be assumed
in the back-analysis of the test.
To gather as much information on the shear connection as
possible, two stud arrangements were used in the beam specimen. On the left hand side of the beam in Fig 3, seven pairs of
studs were through-deck welded in every other trough in the
‘favourable’ position (see Fig 2(b)); this arrangement was
adopted to enable the lowest degree of shear connection to be
provided, so as to gain information on the in situ load–slip
curves. (N.B. The longitudinal spacing provided was very close
to achieving the maximum spacing requirements given in BS
5950-3.1, Clause 5.4.8.1 of the smaller of 600mm or 4 × overall
slab depth; however, the spacing was compliant with the
requirements given in Eurocode 4, 6.6.5.5(3) of the lesser of 6 ×
the total slab thickness or 800mm). On the right hand side of
the beam in Fig 3, 14 studs were through-deck welded singly
in every trough in the ‘favourable’ position. This welding
arrangement meant that an equal number of studs were
provided in each half of the beam.
The steel used in the fabrication of the specimen was grade
S355. From coupons cut from the UB, the mean measured yield
Table 1: Measured steel properties
Top flange
Web
Bottom flange
fym (N/mm²)
368
378
388
Table 2: Measured concrete properties
Age (days)
fcm,cube
(N/mm²)
fck,cube
(N/mm²)
fcm
(N/mm²)
fck
(N/mm²)
Ecm
(kN/mm²)
42
25.3
16.5‡
20.4
12.4‡
26.1†
‡ characteristic value determined from BS EN 1992-1-1: 200413
† modulus of elasticity derived from beam bending stiffness
Table 3: Cross-sectional properties for 305×165×46 kg/m
UB
D
(mm)
B
(mm)
t
(mm)
T
(mm)
Root
radius
(mm)
Crosssectional
area
(mm²)
309.8
166.6
7.3
10.9
8.9†
5795
† Nominal value
To facilitate the back-analysis, the geometry of the steel section
was measured at each of the 16 instrumented cross-sections.
The average measured properties of the UB are presented in
Table 3.
The slab was 140mm thick, and was conventionally reinforced with one layer of A193 square mesh fabric, which was laid
directly on the deck (i.e. the top of the studs projected 11mm
above the mesh). A total slab width of 2500mm was provided,
which corresponds exactly with the effective width requirements given in current Standards of beam span/4.
To represent UK practice, a typical 60mm deep trapezoidal
deck was fixed perpendicular to the longitudinal axis of the
beam (comprising of a Multideck 60-V2 profile); this provided
an average concrete rib breadth b0 of 150mm, a ‘shoulder height’
hp,n of 60.9mm and an overall height hp,g of 69.9mm (see Fig 2).
Owing to the fact that the rules given in Eurocode 4 reduce the
stud resistance for sheet thicknesses less than or equal to
1.0mm, a 0.9mm gauge deck was used to provide the most
unfavourable case (thinner gauge decks are not currently in use
in the UK).
Instrumentation
Measured material properties
Location
Geometric properties
To enable the internal forces to be evaluated, the steel beam was
instrumented with strain gauges on the top and bottom flange
at cross-sections corresponding to the shear connector positions. As well as determining the build-up of axial force, the
internal bending moments were evaluated to check the reliability of the strain gauge readings. The slip distribution at the
shear connection was established from horizontally mounted
transducers, which monitored the relative displacement
between bars cast in the concrete behind the shear connector
positions and the top flange of the beam (see Fig 4). The internal load–slip curves for the shear connectors were therefore
evaluated by plotting the change in axial force at each crosssection against the corresponding slip.
Owing to the fact that two stud arrangements were provided,
the different load–slip characteristics for the shear connectors
meant that there was no position of symmetry. The steel beam
was therefore instrumented at 16 cross-sections along its entire
length, comprising 48 strain gauges and 30 displacement transducers (see Fig 3). The deflection was monitored by vertical
transducers at third-points and at mid-span of the beam. An
image of the instrumentation on the side with pairs of studs is
shown in Fig 4.
In an attempt to closely simulate the bending moment from
a uniformly distributed load, four equally spaced point-loads
were applied along the centre-line of the beam. The loading was
supplied by 60t jacks with a stroke of 300mm, which were
located at third points along the beam (see Fig 3).
To enable large deflections to be applied to the beam specimen, rollers were provided under the ram of the jacks as well
as to the ends of the spreader beams. The bearing plates were
plastered directly onto the concrete slab and, to eliminate any
question of the shear connection being assisted by horizontal
forces taken through the load system, horizontal sliders were
15 May 2007 – The Structural Engineer|33
paper: hicks
Fig 4. (left)
Instrumentation in
the vicinity of the
support on the side
with pairs of studs
Fig 6. (right)
Concrete pull-out
failure around pairs
of studs revealed
after separating the
slab from the deck
provided at the interface between the bearing plate and the
rollers to the spreader beam.
General behaviour of the beam
The props were left in place until the target compressive
strength of the concrete was achieved. Once the props were
struck, the self-weight loads on the composite cross-section
(which amounted to 74.5kN) resulted in a mid-span deflection
of 66mm. The end-slips indicated that there was a modest
asymmetry in the shear connector behaviour, with measured
values of 0.15 and 0.08mm for the side with pairs of studs and
single studs respectively.
Initially, the beam was subjected to three cycles of load up to
approximately 70% of the predicted ultimate load (which was
based on BS 5950-3.1, using mean measured material properties, and the shoulder height of the deck hp,n in Eq.(1)); this
loading was undertaken to ensure that the beam was elastic in
the initial stages and to identify the load level when first yield
occurred. The applied moment versus mid-span deflection plot
is presented in Fig 5.
As can be seen from Fig 5, for load Cycle 4 and 5 up to a midspan deflection of 300mm (equivalent to span/30), the specimen
exhibited excellent ductility. Significant slip was recorded on the
side with pairs of shear connectors and, at an end-slip of
26.5mm, it was felt that sufficient data had been collected,
which resulted in load Cycle 5 being terminated.
From test observations, it was clear that significant distress
was caused at the shear connector positions with pairs of studs.
By a visual inspection of the locations where the deck had
delaminated from the concrete slab during load Cycle 4, it
appeared that concrete pull-out had occurred, with pyramidal
shear cones of concrete clearly visible around the studs. Pullout failure was later confirmed when the concrete slab was
removed from the deck to reveal the failure surfaces shown in
Fig 6; this failure mode was similar to that experienced in the
companion push tests discussed below and reported elsewhere14. In addition to this, there was some evidence of uplift
at the unstudded concrete ribs adjacent to the positions where
studs were welded in pairs (see Fig 7). In an attempt to gain
further information on single studs, an end stop was welded to
the left end of the beam in Fig 3 to subject these connectors to
greater levels of longitudinal shear force.
Fig 7. Slab uplift of
unstudded rib
adjacent to locations
where studs were
welded in pairs
(Position S7)
As can be seen from load Cycle 6 and 7 in Fig 5, the end stop
was successful in that, as well as significantly increasing the
stiffness, the beam was capable of taking further load. Under
this new set-up, load was applied until an end slip of 4.1mm was
achieved on the side with single studs welded in each trough.
At this point, bearing failure of the concrete in front of the end
stop occurred, which resulted in the test being terminated.
However, as can be seen from Cycle 6 and 7 in Fig 5, although
the test was stopped prematurely, the resistance corresponded
almost exactly with the BS 5950-3.1 predictions for the side
with single studs.
From observations made on the beam, it was clear that the
plateau observed in the moment–deflection plot during load
Cycle 4 and 5 was caused by a plastic hinge forming in the beam
on the side with pairs of studs. A visual inspection of the steel
beam confirmed this in that Lüders’ wedges were observed in
the web of the UB, adjacent to cross-section S3 on the side with
pairs of studs (see Fig 3).
Back-analysis of the beam
Fig 5. (below)
10m span beam
applied moment
versus mid-span
deflection
To enable the internal forces to be evaluated, a special version
of the SCI software ‘BFIRE’ was developed. At each instrumented cross-section, the steel beam was sub-divided into a
number of rectangular elements. The strain at the centroid of
each of these elements was calculated from the strain gauge
readings and compared to the stress–strain curves that had
been established from coupons cut from the UB. From the distribution of stresses on the cross-section, the internal forces and
moments within the steel beam were derived.
Load–slip performance of shear connectors
The load–slip curves for the pairs of studs closest to the support
(7th pair from mid-span) and 646mm from the support (6th pair
from mid-span) are presented in Fig 8. As can be seen from this
plot, the pairs of studs possessed a high initial stiffness until a
peak resistance of approximately 160kN was achieved at slips
of about 1mm. As the slip increases, the resistance reduces to
approximately 66 and 100kN (i.e., 33 and 50kN per stud);
beyond this point, the behaviour is very ductile indeed, and a
plateau is maintained up to a slip of 26mm.
The load–slip curves for the single studs closest to the
support (13th and 14th stud from mid-span) and 646mm from
the support (11th and 12th stud from mid-span) are presented
in Fig 9. As can be seen from this plot, the single studs possessed
34|The Structural Engineer – 15 May 2007
paper: hicks
Fig 8. (above top) Comparison of axial force in beam versus slip for 6th and 7th
pair of studs (Load Cycles 1 to 5)
loading comprised of two point loads applied to the beam.
Owing to the new beam configuration, the left hand side of the
beam had seven single studs welded in the unfavourable position (see Fig 2(c)); whereas, the right hand side of the beam had
seven single studs welded in the favourable position (see Fig
2(b)). To enable the load–slip curves for single studs to be determined in situ, two extra cross-sections were instrumented at
positions N1.5 and N7.5 (see Fig 10).
Because it was expected that the unfavourable studs would
have a lower resistance than the studs welded in the favourable
position, an extra centrally welded stud was provided over the
left hand support. A set of concrete cubes had been set aside
after the previous beam test, which were crushed on the day of
the test to determine the compressive cube strength of the
concrete. Although the age of the concrete was 106-days, it was
found that the concrete possessed exactly the same properties
as those presented in Table 2.
General behaviour of the beam
Fig 9. Comparison of axial force in beam versus slip for 13th and 14th single studs
together with 11th and 12th single studs (Load Cycles 1 to 7)
a much higher resistance than the pairs of stud connectors.
Assuming that the load was equally distributed between the
studs, a maximum resistance of about 113kN was achieved per
stud and was maintained with increasing slip. However, owing
to the fact that bearing failure of the concrete in front of the end
stop occurred, the magnitude of the slips were quite low for the
single studs in comparison with studs welded in pairs (4mm cf.
26mm). As a consequence of this, it was impossible to know
whether the single studs would have been able to maintain
their load for higher levels of slip or, in a similar way as for pairs
of studs, their resistance would have reduced. Furthermore,
the assumption that the load was equally distributed between
the studs may not be entirely correct.
The applied moment versus mid-span deflection plot for the
5.0m span beam specimen is presented in Fig 11. During the
first two load-cycles, much greater end slips were recorded on
the side with unfavourable studs (4.5mm cf. 2.4mm at the
maximum moment applied in load Cycle 1). At an end-slip of
12.9mm, the second load cycle was terminated as it was felt that
sufficient data had been collected for unfavourable studs. As can
be seen from Fig 11, the specimen exhibited excellent ductility
up to this point, which corresponded to a mid-span deflection
of 85mm (equivalent to span/59). Furthermore, the resistance
was 11% higher than the predictions from BS 5950-3.1.
Once the load had been removed, an end stop was welded to
the left end of the beam in Fig 10, in an attempt to subject the
side with single favourable studs to greater levels of longitudinal shear force. As can be seen from load Cycle 3 in Fig 11, under
this new set-up further load and deflection was applied to the
beam specimen. Once an end slip of 24.6mm was achieved on
the side with studs welded in the favourable position, the test
was terminated at a mid-span deflection of 174mm (equivalent
Composite beam specimen 2
Owing to the fact that there was little evidence of a permanent
set within the UB until the plastic hinge position adjacent to
cross-section S3 (see Fig 3), it was decided to gather more data
on single studs by cutting the specimen in half to form a 5.0m
span beam. The tests on this shorter specimen are described
below.
As can be seen from Fig 10, unlike the previous test, the
Fig 10. (below) 5m span composite beam specimen
Fig 11. (above) 5m span beam applied moment versus mid-span deflection
to span/29) owing to concerns over the stability of the test rig
from the large curvatures.
In a similar way as the behaviour exhibited in the early
stages of the test, the applied moment–deflection plot for load
Cycle 3 showed that the specimen possessed excellent ductility.
Moreover, the performance of the studs welded in the
favourable position was much better than that assumed by
current Standards, which resulted in a resistance that was 12%
higher than that predicted by BS 5950-3.1. Failure was caused
by a plastic hinge forming in the beam adjacent to the load position near instrumented position N3 (see Fig 10), which was
confirmed by Lüders wedges being clearly observed in the web
of the UB.
15 May 2007 – The Structural Engineer|35
paper: hicks
normal force to the face of the slab15,16.
Discussion
Characteristic properties of stud connectors from beam
tests
Back-analysis of the beam
Load–slip performance of shear connectors
The load–slip curves for single studs in the favourable, central
and unfavourable position are presented in Fig 12. As can be
seen from these plots, much less slip was needed for the stud
in the unfavourable position to achieve its peak resistance
compared to the favourable and centrally welded studs (2.8mm
cf. 7.9mm and 9.6mm respectively). In terms of the slip requirements given by the current Standards, significant ductility was
achieved for all of the stud arrangements; in particular, for
studs welded in the favourable position (see Fig 2(b)), where a
plateau was maintained up to a slip of approximately 25mm.
Companion push tests
To provide a link between the behaviour of studs in a beam
compared to that achieved in a push test, six push specimens
were constructed using exactly the same batch of concrete that
was used in the beam specimen. These specimens varied
slightly from standard test given in Eurocode 4, in that the steel
section was split to form two structural ‘tees’ (see Fig 13). The
advantage of this set-up was that the casting process was
greatly simplified and the same concrete used in the beam
specimen could be used for both test slabs (thereby reducing the
variation in the results from different concrete mixes).
The load–slip curves from the beam tests compared to the
push tests are shown in Fig 14 and Fig 15; these curves represent the lowest results obtained from each test type for a particular stud arrangement. As can be seen from these plots, with
the exception of initial stiffness, there appears to be no similarity between the performance of studs in these two types of
specimen. It might be argued that the peak resistance achieved
in the push test is comparable with the plateau exhibited by the
studs in the beam specimen. However, the slip measured in the
push test is well below the levels achieved in the beam specimen and, if considered in isolation, would suggest that the
ductility of the shear connectors was lower than required by the
current Standards for partial shear connection design. It is felt
that the reason for the poor performance in push tests is due to
the absence of a curvature and normal force, which exists in real
composite beams from the floor loading. Brittle failures in push
tests have also been observed elsewhere, which has led some
researchers to modify the standard specimen by applying a
36|The Structural Engineer – 15 May 2007
Fig 12. (above)
Comparison of axial
force in beam versus
slip for single studs
in the favourable,
central and
unfavourable
position
Fig 13. ( below left)
Companion push
test specimen
Fig 14. ( below)
Load–slip curve for
single studs in the
favourable position –
Beam test versus
push test
Fig 15. (below right)
Load–slip curve for
pairs of studs in the
favourable position –
Beam test (7th pair)
versus push test
Based on the load–slip curves evaluated from the beam tests for
studs in the favourable, central and unfavourable position, the
characteristic resistances PRk presented in Table 4 have been
taken to be 0.9 times the minimum failure load per stud Pe. To
examine the performance of the current Standards, predictions
based on BS 5950-3.1 and Eurocode 4 (EC 4) are presented for
comparison purposes; in these comparisons, predicted stud
resistances based on the deck shoulder height, PRk,n and overall
deck height, PRk,g are considered. As can be seen from Table 4,
the characteristic resistance of 96kN for single centrally welded
studs compares well with the predictions for PRk,n using BS
5950-3.1. Moreover, the characteristic slip capacity of the stud
in this position out-performs the Eurocode 4 requirements for
ductile connectors by almost a factor of two. However, the
predictions for PRk,n using Eurocode 4 appear to be overly
conservative for single studs.
For single studs in the unfavourable position, the predicted
resistance given by BS 5950-3.1 is conservative in comparison
with that achieved in the beam tests (see Table 4). Furthermore,
the Eurocode 4 requirement that studs should be welded alternately on the two sides of the trough throughout the length of
the span (when it is not possible to place studs centrally within
the trough), seems to be appropriate in that the average of the
favourable and unfavourable resistances corresponds almost
exactly with the resistance of the centrally welded stud.
For pairs of studs, the characteristic resistance is 73kN per
stud. Although this resistance is close to the value of PRk,n calculated with BS 5950-3.1, owing to the shape of the load–slip plot
(see 7th pair in Fig 12), the corresponding slip capacity is only
0.7mm. This is well below the 6mm requirement given in
Eurocode 4 for cases when a connector may be taken as ductile.
In order to achieve a slip capacity of 6mm, the characteristic
resistance has been down-rated to 49kN per stud in Table 4; this
resistance is close to the value of PRk,n calculated according to
paper: hicks
Table 4: Characteristic properties of studs in the Favourable (F), Central (C) and
Unfavourable position (U) evaluated from beam test
nr
Stud
position
Pe
PRk
PRk/PRk,Central
δuk
(mm)
BS 5950
PRk/PRk,n
BS 5950
PRk/PRk,g
EC 4
PRk/PRk,n
EC 4
PRk/PRk,g
1
F
123
111
1.16
9.6
1.27
(87.3)
1.94
(57.2)
2.19
(50.6)
2.88
(38.5)
1
C
107
96
1.00
11.9
1.10
(87.3)
1.68
(57.2)
1.90
(50.6)
2.49
(38.5)
1
U
89
76*
0.79
6.0
1.32
(57.6)
2.36
(32.2)
–
–
2
F
80
49*
0.51
6.0
0.70
(69.8)
1.21
(40.4)
1.18
(41.7)
1.80
(27.2)
†PRk,Central is the characteristic resistance of the centrally welded studs
* down-rated to achieve a slip capacity of 6mm
( ) calculated characteristic resistance in kN
Table 5: Characteristic resistance of studs welded in the
favourable position evaluated from the standard push test
nr
Stud
PRk
δuk BS 5950
position (kN) (mm) PRk/PRk,n
BS 5950
PRk/PRk,g
EC 4
PRk/PRk,n
EC 4
PRk/PRk,g
1
F
76.2
2.6
0.87
(87.3)
1.33
(57.2)
1.51
(50.6)
1.98
(38.5)
2
F
46.1
2.4
0.66
(69.8)
1.14
(40.4)
1.11
(41.7)
1.69
(27.2)
( ) calculated characteristic resistance in kN
Eurocode 4.
Characteristic properties of stud connectors welded in
the favourable position from the companion push tests
Based on the load–slip plots from the companion push tests (see
Fig 14 and Fig 15), the characteristic values presented in Table
5 were calculated according to the requirements given in Annex
B of Eurocode 4.
By comparing Table 5 with Table 4, the single studs in the
beam test out-performed those in the push test both in terms
of resistance and ductility (by 46% and 269% respectively).
However, when pairs of studs are considered, the characteristic resistance of 46.1kN agrees well with the value given in
Table 4.
Design resistance of through-deck welded studs
Based on BS 5950-3.1, the ratio of the stud reduction factors for
pairs of studs to single studs is 0.6/0.85 = 0.71. From the present
investigation, the ratio of the characteristic resistance for pairs
of studs to single studs is 49/111 = 0.44 for the beam tests, and
46.1/76.2 = 0.60 for the companion push tests. From this
comparison it can be concluded that, if results from push tests
can be regarded as basically comparative, the resistance of the
pairs of studs from the beam test is lower than expected.
However, from the observed uplift at the unstudded concrete
ribs adjacent to the positions where studs were welded in pairs
(see Fig 7), it is believed that performance of the studs in this
arrangement was adversely affected by their relatively high
longitudinal spacing.
Although the effect of stud spacing will be examined in future
research, the poor performance of the pairs of studs in the
current investigation suggests that the reduction factor given
in BS 5950-3.1 is too high. Moreover, the current code assumption that the resistance of pairs of studs is proportional to 1/√nr
(where nr is the number of studs per rib), does not appear to be
appropriate. As an interim measure, it is suggested that for nr
= 2 the factor of 0.85 /√2 = 0.6 in Eq. (1) should be reduced to
0.37 (i.e. 0.85 × 0.44 = 0.37) to give the following reduction
factor equation:
k = 0.37(b0 /hp){(hsc /hp) – 1} but k ≤ 0.75 for nr = 2
...(2)
where b0 is the breadth of the concrete rib, hp is the depth of
the profiled steel sheet (provided that the stud projects at least
35mm above the shoulder of the deck, using the ‘shoulder
height’ hp,n as defined in Fig 2) and hsc is the height of the stud.
Eq. (2) gives similar resistances to those calculated from
Eurocode 4, when the shoulder height of the deck is inserted
within the reduction factor formula. From the tests conducted in
this investigation, it is also recommended that pairs of studs
should not be arranged in line (i.e. in the favourable and
unfavourable position within the same trough). A more beneficial
arrangement is to have the studs arranged side-by-side, with the
maximum transverse spacing that may be practically achieved.
Owing to the very good agreement with the characteristic
resistance achieved in the beam tests it is recommended that,
for single studs through-deck welded in trapezoidal decks (i.e.
nr = 1), the current practice of using the deck shoulder height
in Eq. (1) should be maintained.
Minimum degree of shear connection
The minimum degree of shear connection permitted in BS 59503.1 and Eurocode 4 is 40%. However, the definition of the point
on the beam where the degree of shear connection should be
calculated is unclear in both of these Standards. According to
BS 5950-3.1, the critical cross-section that should be considered
in positive moment regions is defined by the point of maximum
moment. This definition was used in the initial design of the test
specimens, in that the critical cross-section was considered to
be at mid-span (i.e. the maximum moment comprised of the self
weight moment plus the moment applied by the jacks).
At maximum moment, the axial forces derived from the strain
gauge readings indicated that the critical cross-sections should
have corresponded to the point load positions closest to mid-span
(i.e. at position N3 and S3 in Fig 3). As a consequence of this, very
low levels of shear connection were actually provided. For
example, for the 10m span beam specimen with studs provided
in pairs, only 24% shear connection was achieved (it is believed
that this low level of shear connection is one of the reasons why
the moment resistance was below the predictions given by BS
5950-3.1, in that the connectors became fully loaded long before
the steel beam was fully yielded). It is therefore recommended
that when a beam is subjected to concentrated loads, the critical cross-section that should be considered in design should
correspond to the point load positions.
Standard push test
From comparisons of the load–slip curves it is clear that, for
the deck type considered in this investigation, any brittleness exhibited in the push test is as a result of a deficiency
in the standard specimen rather than the shear connection.
As a consequence of this, some caution should be shown
towards any new products that are claimed to increase the
ductility of stud connectors since, if these devices were developed using push specimens, the improvements to the
performance could be artificial.
Although the tests presented in this paper have shown
that the performance of studs in typical UK decks are
ductile, there remains the problem that, if new floor decks
are developed, it is difficult to identify cases when the behaviour could be brittle unless beam tests are undertaken.
Furthermore, should the performance of studs in particular
decks be better than that suggested by the current
Standards (which have been developed from push test
results) it is difficult to quantify the improvements in terms
of design rules without a standardised specimen that better
reflects the conditions that exist in a real beam.
In the absence of a new specimen, it is recommended that,
when decking is employed, the push test arrangement shown
in Fig 16 should be adopted. Although this specimen does not
solve the problem of brittleness, this arrangement appears
to eliminate an artificial mode of failure caused by the rotation of the last studded rib at the top of the specimen17,
which has sometimes been described as ‘back-breaking’. It is
also recommended that edge trim from pressed strips of thin
galvanised steel, which are normally used to form the floor
edge in buildings, should be removed from the push specimens prior to testing to reduce the scatter in results (from
earlier comparison tests, it was found that the coefficient of
15 May 2007 – The Structural Engineer|37
paper: hicks
variation for specimens with and without edge trim was
9.98 and 4.4% respectively).
Conclusions
Two full-scale composite beam tests have been undertaken with
a trapezoidal deck positioned perpendicular to the longitudinal
axis of the beam (i.e. a secondary beam), and through-deck
welded shear connectors. Unlike typical UK practice, the specimens were propped during construction to apply the most
unfavourable loading to the shear connectors. This form of
construction, together with other unfavourable combinations of
variables, was adopted to demonstrate the slip capacity that can
be achieved in a beam, together with the level of safety that
exists in current design Standards. Both specimens exhibited
excellent ductility in terms of overall load–deflection performance as well as slip capacity at the shear connection, which far
Fig 16. Standard
push test with
profiled steel
decking
surpassed the levels of slip assumed in the development of the
rules for partial shear connection given in BS 5950-3.1 and
Eurocode 4.
The performance of the beams with single studs welded in the
troughs of trapezoidal decking demonstrated that the current
design practice of using the shoulder height of the deck in
BS 5950-3.1 should be maintained. However, for pairs of studs
in the favourable position, the performance in the beam test was
lower than anticipated. Although it is felt that the performance
of the studs in this arrangement was adversely affected by their
longitudinal spacing, a modified reduction factor formula has
been presented as an interim measure. It is believed that
research, which is currently underway, will enable this proposed
design equation to be relaxed.
The definition of the critical cross-section where the degree
of shear connection should be calculated is unclear in both
BS 5950-3.1 and Eurocode 4. Based on the behaviour of the test
specimens it is concluded that, when a beam is subjected to
concentrated loads, the critical cross-section that is considered
in design should correspond to the point load positions.
From comparisons of the load–slip curves from the beam
tests with push tests, it is clear that any brittleness exhibited
in the push test is as a result of a deficiency in the standard
push specimen rather than the shear connection. It is recommended that a standardised specimen, which better reflects the
conditions that exist in a real beam, should be developed. In the
absence of a new specimen, a standard push test arrangement
for slabs that use decking has been presented.
Acknowledgments
Financial support for this investigation was provided by the
Floor/Deck Group of the Metal Cladding and Roofing
Manufacturers Association (MCRMA), Corus Construction
Services & Development and Corus Strip Products UK. The
author wishes to thank the help and assistance of Dr R. E.
McConnel, M. R. Touhey and the technical staff of
Cambridge University Engineering Department whose
expertise ensured the success of the testing programme.
Thanks also go to Prof. R. P. Johnson of University of
Warwick for his support and advice, together with Dr J. W.
Rackham, Dr W. I. Simms and C. King of The Steel
Construction Institute.
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38|The Structural Engineer – 15 May 2007
