Materials and Structures/Mat6riaux et Constructions, Vol. 32, May 1999, pp 290-297

Behaviour of headed stud shear connectors under low- cycle high amplitude displacements

O. S. Bursi and G. Gramola Department of Mechanical and Structural Engineering, University of Trento, Trento, Italy

Paper received: March 9, 1998; Paper accepted: November 27, 1998

A B S T R A C T R I~ S U M I~

Two series of pull-push specimens embodying headed stud shear connectors were built and tested as part of a general investigation on aseismic design of steel- concrete composite beams. Thereby, both the geometri- cal and mechanical characteristics of specimens are simi- lar to those of steel-concrete composite substructures with full and partial shear connection.

Specimens were endowed with different boundary conditions in order to assess the relevant effects on the response. Moreover, to characterise the specimens from a seismic standpoint, they were subjected to suites of monotonic, variable and constant amplitude displace- ments according to a cumulative damage test program.

Main results are commented upon and assessed in terms of both yielding and maximum shear strength capacity as well as ultimate displacement ductility. Finally, a comparison between experimental and ideal loads derived by relevant codes provides an estimate of their accuracy.

Deux s&ies d'&hantillons de type pull-push compre- nant des goujons ont dtd construites et test&s clans le cadre d'une analyse g&&ale des projets antisismiques de poutres mixtes b&on-acier. II en r&ulte que les caract&'stiques g&md- triques et me'caniques des &hantillons sont semblables h celle des substructures mixtes bdton-acier ayant des connections compl~tes et partielles.

Les &hantillons ont dtd dot& de diff&entes conditions aux extrdmite's, afin d'dvaluer les effets correspondants sur leur r@onse. En outre, pour caract&iser les e'chantillons du point de vue sismique, on les a soumis a des d@lacements mono- tones et h des amplitudes de d@lacement fixes ou variables, selon un programme d'essai d'endommagement cumulatif. Les principaux r&ultats sont comment&, tant en ce qui concerne la contrainte maximale de cisaillement qu' au niveau de la ddformation maximale h la rupture.

Finalement une comparaison entre les charges exp&imen- tales et les charges id&les d&iv&s des normes fournit une esti- mation de la pr&ision des normes.

1. INTRODUCTION

The majority of tests regarding standardised speci- mens of push type under reversed loading were carried out focussing on the performance of stud shear connec- tors subjected to high-cycle fatigue. Thereby, Log S - Log N limit domains for shear connectors or slip versus cycle number relationships were defined for bridge applications [1-3]. Other investigations were conducted in order to estimate the shakedown behaviour of stud shear connectors [4] or to quantify the stiffness variation of shear connectors under cyclic loading with limited reversals [5]. In the aforementioned test series different testing equipments were adopted incorporating speci- mens endowed with shear studs located in two rows or

single studs in conditions of direct shear. Moreover, dif- ferent boundary conditions were reproduced within the set-ups and only Gattesco and Giuriani [3] commented on the adoption of peculiar boundary conditions.

Several loading histories were also applied to the specimens. Indeed, some tests were run under load con- trol encompassing cycles around a positive load [1-2] whilst others displacement sequences included limited reversals in view of bridge applications [3, 5]. Conversely, other investigators performed tests in load control, by imposing loading with the same or the oppo- site sign. In this last condition, the resulting slip was unsymmetrical owing to unequal boundary conditions in the pull and push regime, respectively [4].

From a seismic standpoint, tests on stud shear con-

1359-5997/99 9 RILEM 290

Bursi, Gramola

PROFILED SHEETING /

! 1o, 9 20. 330 ,120,

O ~ g : : : z : IPE 330

J !

9 . ~ . ,~ .

200 200 200

L I / II

~176 g - , II

9 O illl II t

r

REBARS

STUD r 1=100 @ st 37 - 3K

t

Fig. 1 - Geometrical characteristics of the specimens and strain gauge lay-out .

whilst the rele- vant geometri- cal characteris- tic are depicted in Fig. 1. Both the beam stub and the con- crete slab char- acteristics were similar to those of the compan- ion steel-con- crete compos- ite beams [9], according to the Eurocode 4 requirements for specific

nectors were carried out by Hawkins and Mitchell [6] both in a monotonic and low-cycle regime. Hence, specimen performances were compared and those sub- jected to cyclic loading exhibited a reduction of shear strength and ultimate displacement ductility. Test results emphasised the advantage of orienting the profiled steel sheeting parallel to the shear direction as well as the importance of having large stud connector spacing. However, the influence of boundary conditions on specimen performance was not commented upon.

Other investigators [7] subjected push-type speci- mens to reversed cyclic displacements in order to study the effect of a dissipation device, named skirt, on the stud connector performance. They realised that bound- ary conditions are unsymmetrical in a typical push test, and, as a result, the steel-concrete interface slip was adopted as control parameter. To sum up, very few pull- push tests were conducted on stud shear connectors in a low-cycle high amplitude regime. Moreover, in the con- text of reversed displacements, open problems still remain on a proper definition of boundary conditions.

The investigation presented in this paper intends to determine the shear strength capacity and ultimate dis- placement ductility of stud shear connectors. Thereby, pull-push specimens were subjected both to a monoto- nic and to a low-cycle reversed displacement regime according to a cumulative damage test program in order to characterise their seismic performances. Moreover, the influence of boundary conditions of the set-up as well as of loading histories on stud shear connector per- formances was investigated. Finally, a comparison between experimental and ideal shear loads predicted by the relevant codes is provided.

2. EXPERIMENTAL INVESTIGATION

2.1 Specimen design

Eleven elemental push-type specimens divided into two series were fabricated. The nomenclature that iden- tifies the specimens is reported in Column 2 of Table 1

push tests [8]. The studs were placed in two rows, see Fig. 1, to allow stud shear loads within the concrete slabs to be redistributed. As a result, the connectors could fail at their mean strength but also as a group of eight. Indeed, permitting connectors to fail in group under a monotonic or fatigue loading was found to reduce the result scatter [10].

Profiled steel sheeting is widely used as permanent formwork for composite floor slabs in buildings. Hence, headed stud shear connectors are placed in troughs, the span of which is normally either transverse or parallel to the span of the beam. Since Hawkins and Mitchell [6] found that shear studs with large spacing located in lon- gitudinal troughs determine higher shear strength and larger ultimate displacement ductility than studs located in transverse troughs, the orientation and spacing depicted in Fig. i was adopted.

The slab reinforcement is constituted of a mesh of~ 12 and is located as illustrated in Fig. 1. Transverse reinforce- ments were designed against longitudinal splitting. The surface of the steel flanges was waxed to prevent bond and to reduce friction across the steel-flange/concrete slab interfaces. Thereby, the actual interface condition in the companion composite beam specimens is not reproduced correctly by the push-type specimens. However, the shear strength reduction in the push-type specimens is compen-

Tab le 1 - Specimen nomenclature and test parameters

Series

1

Specimen Boundary conditions

PM-Ol A PM-O2 A PC-D1 A PC-D2 A

NPM-01 B NPM-02 B NPC-Ol B NPC-02 C NPC-03 C NPC-O4 C NPC-O5 C

Number of instrumented studs

Displacement test procedure

MONOTONIC MONOTONIC

SPDP-1 SPDP-1

MONOTONIC MONOTONIC

SPDP-1 SPDP-1 SPDP-2 SPDP-3 SPDP-4

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Materials and Structures/Mat~riaux et Constructions, Vol. 32, May 1999

sated by the higher shear capacity expected for this type of test [11].

TRW Nelson studs were endowed with a shank diameter of 15.9 mm and a mean height of 101.7 mm. By using a TRW Nelson welding system a mean welded height of 4.5 mm was obtained.

2.2 Material properties

The properties of concrete and shear studs are col- lected in Column 3 and 4 of Table 2, respectively. Due to the rheological properties of concrete, the first series was characterised by slight different properties among the specimens. Conversely, specimens of the second series were tested in a short period of time, thus main- taining similar material properties.

2.3 Specimen fabrication

Push-type specimens were assembled in a water- proofed plywood form. Due to the use of an IPE 300 beam section, equal to the one adopted in the compan- ion composite substructures [9], each slab was casted horizontally, locating the shear studs in a vertical posi- tion. Thereby, the slabs were casted at different stages owing to concrete hardening. Nonetheless, the variation of properties between the companion slabs was limited.

Moreover, the concrete was compacted with an elec- tric vibrator and a dimensional tolerance of less than 5 percent was achieved in the specimen construction.

2.4 Test set-up

As mentioned in Section 1, different test set-ups embodying consistent boundary condit ions were adopted around the world in order to carry out push and/or pull-push tests. Indeed, only Gattesco and Giuriani [3] used boundary conditions able to reproduce a stress state of direct shear onto the connectors, which turns out to be appropriate for connectors close to ends in simply supported beams.

Table 2 - Specimen material properties

Series

1st series 1st series 2nd series

Displacement test procedure

Monotonic Cyclic

Monotonic & cyclic

Concrete

fcm fctm Ecm (Mea) (MPa) (MPa)

41.1 3.4 34800 36.5 3.4 33700 32.6 3.1 32700

Shear studs

(Mfl~a) fu (MPa)

427 578 427 548 414 528

Other researches decided to use steel plates on the slab edges aiming at simulating the passive resistance offered by the continuation of concrete slabs [4, 7]. It is evident that the aforementioned fastening devices create a confinement effect on shear connectors, thus increas- ing the ultimate displacement ductility and the ensuing energy dissipation. Thereby, different boundary condi- tions were adopted in the two series covering the ones close to the standard push-tests [8] as well as the ones that reproduce the maximum confinement effect in the transversal direction [4, 7].

The testing equipment adopted to exert the monoto- nic or cycling displacement d on each specimen is illus- trated in Fig. 2 schematically. The specimens were located in a reaction frame able to transfer the load to the support- ing steel columns. For brevity, only a part of the rigid counterbeam that belongs to the reaction frame is shown in the same figure. Due to the alternating nature of load- ing, it was necessary to adopt post-tensioning bars in order to transfer the reaction force from the concrete slabs to the counterbeam. In detail, the fastening bars were tensioned to exert a slight compression of about 0.5 N/mm 2. That stress level represents the case in which the shear studs are located near the inflection points of a beam.

To sum up, boundary conditions labelled "A" in Fig. 2, correspond to the conditions suitable for monoto- nic tests suggested by Eurocode 4 [8]. Fastening configu- rations "B" also depicted in Fig. 2 were conceived to sim- ulate the passive resistance offered by the continuation of concrete slabs [4, 7]. However during testing, some hori- zontal movements of concrete slabs was observed owing to the limited value of the post-tensioning pressure. Thereby, boundary conditions with locking devices were conceived, they are labelled "C" in Fig. 2, in order to impede any relative movement of concrete slabs.

FAST[ BAI

ANGLES

LOCKI DEVIC

Fig. 2 - Boundary conditions of the specimens: type "A"; type "B"; type "C'.

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2.5 Displacement test procedures

Specimens of both series were monotonically (push regime) and cyclically (pull-push regime) loaded in a quasi-static fashion, by a suite of sequential-phased dis- placement procedures. Hence, it is intrinsic in the afore- mentioned conventional quasi-static cyclic approach, the uncertainty that relates any cyclic response to the seismic performance. To acquire a comprehensive set of infor- mation from the specimens, two recommended testing procedures were used: i) the so-called Complete Testing Procedure proposed by the European Convention for Constructional Steelwork [12]; ii) the Cumulative Damage Testing Program suggested by the Applied Technology Council [13]. The first type of procedure is adopted to acquire data on the capacity of the specimens, such as the maximum shear strength, ultimate displace- ment ductility, maximum absorbed energy, etc. [12]. It is by far, the most common procedure used in Europe to test components in a cyclic fashion. Conversely, the sec- ond type of procedure permits the establishment of structural capacity parameters that need to be used with a cumulative damage model to predict component perfor- mances under arbitrary loading histories [13].

The test program is collected in Col. 5 of Table 1. In detail, two classical monotonic tests were carried out both in the first and the second test series in accordance with the Complete Testing Program [12]. As a result, a displace-

+ ment elastic limit e. and the corresDondine vield shear + strength P. were determined by pushing the steel beam

versus the counterbeam (see Fig. 2). In addition, the sequential-phased procedure SPDP-1, which is depicted in Fig. 3 and is characterised by sets of equi-amplitude dis-

placements (2 + 2k) e 7, (k = 1,...,n) was applied to the specimens.

The procedure SPDP2 is also illustrated in Fig. 3 and is characterised by a set of equi-amplitude constant dis- placements at 10 eA + in agreement with the Cumulative Damage Testing Program [13]. Moreover, the procedure SPDP-4 that is endowed with an amplitude of 40 e. +, . . . . . ) t well beyond the conventional elasnc hmlt e. + was carried out (see Fig. 4). Finally, the SPDP-3 procedure was con- ceived in order to reproduce a suite of displacements more close to a seismic event. This procedure is depicted in Fig. 4 and derives from the SPDP-1 one in which dis- placements equal to one half and one quarter of the maximum amplitude of each set are included.

It is clear that strain-rate effects may alter the quasi- static load-displacement response of specimens com- pared to the corresponding dynamic response. However, slow-cyclic testing results in a small decrease in strength and in an increase of the deterioration rate [13]. Thereby, results from quasi-static cyclic tests can be con- sidered as conservative for the purpose of performance assessment. Moreover, the characterisation of the long- term behaviour of the specimens would require the analysis of viscous effects. The aforementioned effects are not investigated in this particular research because seismic actions belong to the class of accidental loads.

2.6 Instrumentation

The instrumentation was designed to monitor the behaviour of the specimens during testing, by providing a continuous time record both of displacements (slips)

40 30 20 10

.f0 -10 -20 -30 -40

CYCLE NUMBER CYCLE NUMBER

Fig. 3- Sequential-phased displacement test procedures SPDP-1 and SPDP-2 with variable and con- stant amplitude.

401 301 2Ol 101

"10 I -20 -30 -40

CYCLE NUMBER CYCLE NUMBER

Fig. 4 - Sequential-phased displacement test procedures SPDP-3 and SPDP--4 with variable and con- stant amplitude.

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and reaction force. In detail, the reaction force was measured with a load cell connected to the end of the actuator. The interface slip e between the steel beam and the concrete slabs shown in Fig. 2 was assumed to be the prime parameter of the test control, owing to the different stiffnesses of the test set-up. Such slip e was detected by means of four linear variable differential trans- formers (LVDTs) located as depicted in Fig. 2. Moreover, the slip was measured also with addi- tional four LVDTs that are labelled with circles and illustrated in Fig. 2.

The main response of each specimen was expressed in terms of a monotonic or hysteretic reaction force-slip (P - e) relationship. Deformations of stud connectors were recorded too endowing them with linear strain gauges able to detect both axial and bending deformation compo- nents. The ensuing gauge numbering is illustrated in Fig. 1, whilst the specimens endowed with instrumented shear studs can be inferred from Column 4 of Table 1.

The analog signals from the various measuring devices were conditioned, amplified and digitised by using an A/D converter and then recorded in a com- puter-based data acquisition system.

Series

3. RESULTS

3.1 First Series

For the sake of brevity, only the most significant results are illustrated and commented upon. Moreover, in order to disregard the variability of the material prop- erties, all reaction force-slip relationships (P - e) are plot- ted in a non-dimensional form in which Pcode represents the monotonic shear strength predicted by Eurocode 4 [8]. However, the interface slip has not been dimension- alised, because the relevant conventional elastic limit ey + collected in Column 3 of Table 3 is rather small.

All specimens of this first series failed by stud shear- ing and concrete crushing in the monotonic regime and low-cycle fatigue in the cyclic regime. Thereby, the per- formances of specimens were different as a function of the sequential-phased displacement procedure adopted.

Table 3 - Parameters of the load-slip response Specimen ey+ eu+ eu+ Py+ Pmax + Pcode

(mm) (mm) ey+ (kN) (kN) (kN)

PM-01 0.08 6.1 76.3 339.1 570.4 643.4

PM-02 0.08 3.4 42.5 386.3 565.6 643.4 PC-01 0.25 3.2 12.8 282.2 418.7 643.4 PC-02 0.22 1.3 5.9 377.6 517.9 643.4

NPM-01 0.10 9.9 99.0 315.7 554.3 613.5 NPM-02 0.10 14.7 147.0 330.7 549.2 613.5 NPC-01 0.03 3.2 106.7 226.4 389.7 613.5 NPC-02 0.04 3.7 92.5 277.1 394.2 613.5 NPC-03 0.01 1.0 100.0 50.0 379.2 613.5

NPC-04 0.03 3.2 106.7 261.0 431.3 613.5 NPC-05 0.17 4.0 23.5 315.2 468.1 613.5

Pmax +

Pcode

0.89 0.88 0.65 0.80

0.90 0.90 0.64 0.64 0.62 0.7O 0.76

This observation becomes clear from the responses of PM-01 and PC-01 specimens depicted in Fig. 5a. It is evident both the shear strength and the displacement ductility reduction associated with the reversed cyclic response.

To estimate the aforementioned effects, a conven- tional ultimate displacement eu + was determined in order to define the range in which each specimen can provide a shear strength P > P.+. As a result, the ultimate displace- . . J / . . . ment ductility factor e +/e + in the pushing regime was u y defined. From Columns 4 and 5 of Table 3, in which for brevity, only the value relevant to the pushing regime have been collected, one can quantify the displacement reduction that characterises the specimens subjected to a cyclic regime. Moreover, from Columns 6 and 7 of the same table the maximum shear strength reduction can be assessed. In particular, a severe reduction of shear strength for the specimen PC-01 and a limited reduction for the specimen PC-02 are evident. These differences have to be attributed to the boundary conditions "A" (see Fig. 2) that hamper the reliability of the response. Moreover, boundary conditions "A" determine the maximum shear strength difference between pull and push regimes. This is evident both from Fig. 5a and Fig. 5b in which the skeleton curves of specimens are reported. In detail, an increase of maximum shear strength in the pushing regime of 14.4 and 29.8 percent is found for the PC-01 and PC-02 specimen, respectively.

,.o1- I . . . . . . P ,.o/I ~,~ 0.8t- L PC.Ol ] "" . . . . . . ""

o.61- - 0.41- 0.21-

o ~ -0.2 [ ~. -0.4

-0.6 a) -0.8 -1.0 PULL PUSH

I I I I I I

-6 -4 -2 0 2 4 6 SLIP (mm )

1.0 "~ 0.8 a, ~ 0.6

0.4 0.2

-0.2 ~ -0.4

-0.6 -0.8 -1.0

PM-OI PM-02 PC-O1 PC-02

PULL I I

-6 -4 I I I

-2 0 2 4 SLIP (mm )

b)

PUSH I

6

Fig. 5 - Non- dimensional reac- tion force vs. con- trolled slip of pull-push speci- mens: a) monoto- nic and hysteresis loops; b) skeleton curves.

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'.~ F---OAUO I o.8 --GAUGE ,I 0.6 k ,|,~

o % -0.2 ~ i

-0.4 -0.6 -0.8 a) - 1.0 PUSH

-6000 -3000 0 3000 6000 DEFORMATION ('10 6 )

1.0 0.8

0.6 ~,, 0.4

0.2

-0.2 ~ -0.4

-0.6 -0.8

-1.0

J *

I[--AX'ALST A'N [-- -- FLEXURAL STR.

b)

PULL PUSH , i I i

-6000 -3000 0 3000 6000 DEFORMATION ( '106)

Fig. 6 - Hysteresis loops of non- dimensional reac- tion force vs. stud shear connector strains: a) single strain; b) com- bined strain.

i[ "~ 0.8 ~" 0.6 0.4 o. -0.2

f -0.6 -0.8

-1.0 PULL PUSH I I 1 I I I I I I

-8 -6 -4 -2 0 2 4 6 8 SLIP ( ram )

0.8

0.6 0.4

~ o.

~ -0.2

-0.6 -0.8

-1.0 -8 -6 -4 -2 0 2 4 6 8

suP ( mm )

Fig. 7 - Non- dimensional reac- tion force vs. con- trolled slip of pull-push speci- mens.

The state of deformation in the stud connectors can be assessed by means of the graph depicted in Fig. 6 rele- vant to the PC-01 specimen. In detail, measurements regard gauges 3 and 4 that are located on the connector shank close to the beam flange (see Fig. 1). Strain levels indicate that stud connectors experienced high inelastic deformations. In addition, the strain decomposition in axial and bending components allows both the tensile and bending strains to be quantified. In particular, Fig. 6b highlights that tensile strains of the stud shank remain in the elastic range. Finally, connectors exhibited inelas- tic curvatures at about 1.3 and 1.8 times the shank diam- eter from the stud base, with reference to the push and pull-push specimens, respectively.

3.2 Second Series

As far as the second test series is concerned, for brevity, only some of the specimen responses are pre- sented and commented upon. In detail, the reaction force-slip (P - e) response of the NPC-02 specimen is depicted in Fig. 7a. The observed inelastic behaviour is governed by stud shearing owing to bending and shear whilst the specimen collapse was governed by concrete crushing and stud fracturing. The stiffness and strength degradation in the range of maximum load appears to be limited.

The corresponding cyclic response of the NPC-05 specimen is illustrated in Fig. 7b. An increase of shear strength at large displacements is noted (see also Column 7 of Table 3) owing to the displacement procedure SPDP-4 which is characterised by one cycle only in the variable amplitude range.

Typical failure modes both of concrete and shear connectors in specimens subjected to cyclic loading are illustrated in Figure 8. In detail, the concrete zone close to the shear stud bases within the four connectors appears heavily crushed. Major cracks depart from each connector shank caused by the large steel-concrete inter- face pressure. Conversely, only two large cracks formed outside the connectors owing to the reversal effect of the loading. Clearly, only major cracks formed in the front of connectors for the case of monotonic loading. The aforementioned failure modes repeated themselves among specimens though hysteretic behaviour renders predictions tough. As far as the connector failure mode is concerned, low- fatigue cracking followed by fracture at the weld-collar/shank interface occurred (see Fig. 8b). Indeed, the number of cycles associated with that type of failure could be predicted by means of a cumulative damage model.

Primary load-slip relationships provided by the two specimens tested in a monotonic regime as well as the skeleton curves of the remaining specimens are illus- trated in Fig. 9. One can observe that reversed displace-

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Materials and Structures/Mat6riaux et Constructions, Vol. 32, May 1999

Fig. 8 - Failure zones in concrete slab and at the weld-collar/shank interface.

1.0

"~ 0.8

~.'~ 0.6

0.4

0.2

o

-0.2

-0.4

-0.6

-0.8

-1.0

~ PUSH NPM.O1 - - - NPM-02

: ,~ "..,

SLIP ( tara )

NPC-O1 NPC.02

. . . . NPC.03 NPC.041 NPC.051 PULL

I I I I I I I I

-16 -12 -8 -4 0 4 8 12 16

Fig. 9 - Skeleton curves of SERIES 2 specimens.

i ~ i i i

140

120

100

80

60

40

20 10

Studs ~ 16mm, 1=100 m~ St 37.3k, fu = 528 Mea I

LRFD (1993) /

/ /

, I / Eurocode 4 (1992) /

y ~ NPM-O1, NPM.02

NPC-O1 - NPC.05

I I I I I

20 30 40 50 60 CONCRETE STRENGTH (MPa)

Fig. 10 - Comparison between specimen maximum strength and code prediction.

ment cycles inflict to the specimen responses a reduction both of maximum shear strength and ultimate displace- ment ductility. In detail, the minimum strength ratio Pmax, NPC 03+/Pmax NPM 01 reaches 68.4 percent (see Column 7 of Table 3) whilst the corresponding mini-

mum ductil ity ratio + + (e+le l - ) (e u ley . ) NPC-02 I

NPM-02 lS about 62.9 percent if the tests at equi-constant amplitude, NPC-03-NPC- 05, are disregarded. Indeed, a peculiar displacement test procedure like the SPDP-4 one can determine a large reduction of specimen duc- tility (see Column 5, Specimen NPC-05). It is evident that reduction val- ues both of shear strength and of ultimate displace- ment ductility under cyclic loading have not to be con-

sidered per se. Indeed, they have to be compared to the corresponding demands that characterise composite beams under seismic loading. Thereby, tests on compan- ion composite beams represent a prerequisite in such type of investigation.

Observations can also be made comparing the char- acteristic parameters between the First and the Second series. In particular, the displacement elastic limit e .+ of

9 . } /

PC-01 and PC-02 specimens is much larger than the one that characterises NPC-01 and NPC-02 specimens. The aforementioned observation can be explained through the flexibility of the testing equipment inher- ited by the Boundary conditions "A" with respect to the Boundary conditions "B" and "C" (see Fig. 2). Moreover, the NPC-02 specimen is able to provide a maximum displacement e u + larger than the one offered by the NPC-01 specimen, as the result of Boundary condition "C" with respect to Boundary conditions "B". Such an effect is reflected also by the relevant increase of the yield strength Py+ as well as of the maxi- mum strength P,+, respectively.

Finally, elastic limits values ey + that characterise the specimens of the second series are consistent, thus con- firming the inherent reliability of the test set-up with boundary conditions "C".

4. CODE COMPARISON

Test results on pull-push specimens, i.e. specimens subjected to reversed displacements, show that both stiff- ness and strength of stud connectors reduce at all stages owing to marked yielding and fatigue cracking in the studs as well as propagation and coalescence of microcracks in concrete. At present, design codes do not predict the shear strength of stud connectors in a low-cycle high displace- ment regime and, therefore, it is worthwhile to quantify their accuracy in such conditions.

Let define Pmax + as the ideal shear strength predicted by the code using measured rather than nominal material properties. The corresponding values obtained both with Eurocode 4 [8] and AISC [14] respectively are depicted in Fig. 10. The unsafe prediction provided by

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non-seismic codes is evident as a result of reversed dis- placement effects.

With reference to Eurocode 4 [8] these effects can be quantified by means of Columns 8 and 9 of Table 3, respectively, in which the overall shear strength value corresponding to each specimen, i.e. eight connectors (see Fig. 1), is considered. In the worst case, the code overestimates the maximum cyclic strength of about 38 percent (specimen NPC-03). However, a maximum overestimation of 12 percent (specimen PM-02) is obtained for specimens subjected to monotonic displace- ments too. The aforementioned discrepancy is mainly due to the presence of the profiled steel sheeting which is taken into account only in an approximate manner in the Eurocode 4 [8].

5. CONCLUSIONS

Test results of two series of pull-push specimens embodying headed stud shear connectors were presented in this paper. These results, derived from specimens sub- jected to suites of monotonic, variable and constant equi-amplitude displacements, were re-evaluated and compared in terms of global parameters such as yield shear strength, maximum strength as well as elastic limit and ultimate displacement ductility factors.

To sum up, on the basis of a comparison between the results of the first and the second series, it was found that symmetric boundary conditions improve the perfor- mance of specimens in terms of maximum shear strength and displacement ductility.

Moreover, a comparison between experimental and ideal loads predicted by relevant codes highlights that codes calibrated upon monotonic loading overestimate the actual strength of stud shear connectors and, thereby, appear to be inadequate when reversed displacements govern the stud shear connector response.

6. ACKNOWLEDGEMENTS

This research project is sponsored by grants from the Italian Ministry of the University and Scientific and Technological Research (M.U.R.S.T.) for which the

authors are grateful. However, opinions expressed in this paper are those of the writers, and do not necessarily reflect those of the sponsor. The authors gratefully acknowledge the laboratory assistance.

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[2] Oehler, D. J., 'Deterioration in strength of stud connectors in composite bridge beams',/, of Structural. Eng. 116 (12) (1991) 3417-3431.

[3] Gattesco, N. and Giuriani, E. 'Experimental study on stud shear connectors subjected to cyclic loading', Journal of Constructional SteelResearch 38 (1) (1996) 1-21.

[4] Taplin, G. and Grundy, P., 'Incremental slip of stud shear con- nector under repeated loading', in 'Composite Construction - Conventional and Innovative', Proc. Int. Conf. Innsbruck, September,1997 145-150.

[5] Oehler, D. J. and Coughlan, C. G., 'The shear stiffness of stud shear connections in composite beams', Journal of Constructional Steel Research' 6 (1986) 273-284.

[6] Hawkins, N. M. and Mitchell D., 'Seismic response of composite shear connections',_l, of Structural. Eng. I10 (9) (1984) 2120- 2136.

[7] Astaneh-Asl, A., McMullin, K. M., Fenves, G. L. and Fukuzava, E., 'Innovative semi-rigid steel frames for control of the seismic response of buildings', Report UCB/EERC-93/03 (1993).

[8] European Committee for Standardisation, 'Eurocode 4 - ENV 1994-1-1. Design of composite steel and concrete structures-Part 1-1: General roles and rules for buildings' (1992).

[9] Bursi, O. S. and Ballerini, M., 'Low-cycle behaviour and analysis of steel-concrete composite substructures', in 'Composite Construction - Conventional and Innovative', Proc. Int. Conf. Innsbruck, September, 1997 615-620.

[10] Oehler, DJ . and Johnson, R. P., 'The strength of stud shear connections in composite beams', The Structural Engineer 65B (2) (1987) 44-48.

[11] Oehler, D. J. and Bradford, M. A., 'Composite Steel and Concrete Structural Members - Fundamental Behaviour' 1st edn Elsevier Science (1995).

[12] Technical Committee 13, 'Recommended testing procedures for assessing the behaviour of structural steel elements under cyclic loads' European Convention for Constructional Steelwork, 45 (1986)

[13] Applied Technology Council, 'Guidelines for cyclic testing of components of steel structures', 24 (1992).

[14] American Institute for Steel Construction, 'Load and Resistance Factor Design - Specifications for Structural Steel Buildings', 1 (1993).

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