Structural performance of riveted connections in cold-formed steel wall stud.
The application of cold-formed steel (CFS) in residential framing construction has become increasingly popular all over the world in recent decades, and is now a highly competitive alternative to traditional timber framing. The increase in CFS construction is due to the ongoing development and improvements in the field, the consequential availability of more cost-effective solutions and the broad recognition of the advantages of steel framing.
Some of the widely acknowledged advantages of steel framing over traditional timber framing are: lightness in weight, economy in transport and handling, ease of prefabrication and mass production and recyclable material. However, the optimal design of the CFS framing systems offered to the housing industry is very important in ensuring they can produce a cost-effective solution which is competitive with other traditional structural solutions. The improvement of their framing connections, in particular, could increase the capacity of the framing system without changing the framing materials. Hence, this research proposed to complete the comparative evaluation of the currently-in-use riveted connections of wall studs in Australia and New Zealand.
The most common types of failure which occur at riveted connections are bearing, tear-out, yielding of the gross cross section, fracture of the net section, tilting, pull-out, pull-over and shear failure of the rivets. Hancock et al.(2001) suggested that when the edge and spacing distances of rivets are large enough to avoid tear-out failure, bearing failure of the sheet could occur. Bearing failure often produces stretching of the hole on one side of the rivet, while the sheet material is bunched together on the other side of the rivet. Tear-out failure occurs mostly in connections where the rivet is near the edge of the plate, or the distance between adjacent rivets parallel to the line of force is small. In this condition, the plate tears from a rivet hole to the edge of the plate or to another adjacent rivet hole. To prevent this type of failure, design guidelines specify minimum edge and spacing distances for riveted connections.
When the stresses in the net section of the connected sheet are large enough, failure of the sheet can occur across the rivet hole. This type of failure occurs at the connection because it is often the cross section of the sheet with the lowest net area, and hence it is the weakest section of the sheet. Rivet shear failure occurs when the grade of rivet used does not have shear strength enough to resist loads beyond the load capacity of the sheets. Designers should design in such a way that this type of failure is very unlikely. Tilting usually occurs when two materials of the same thickness are connected, or when the thicker material is against the rivet head. As the two sheets move over each other, the rivets can become tilted, and when the tilting angle becomes large, pull-out failure can occur. Pull-out failure involves the rivet(s) being 'pulled out' of the supporting sheet material under load, as a result of the axial tensile forces in the rivets due to the rotated position of the rivets relative to the direction of load in the connection (Rogers and Hancock 1999). This type of failure is often associated with tilting failure. Pull-over failure involves when the connected sheet 'pulling over' the rivet head. Pullover is enhanced by cyclic loading conditions experienced in high-wind areas and/or during seismic activity, as this can reduce the tensile strength of the sheet. Rivet shear failure usually involves shearing between the rivet stem and rivet head. This type of failure is brittle, and therefore riveted connections are usually designed to prevent to fail on such undesirable failure.
2. Literature review
Yener and White (1984) conducted an experimental study on riveted CFS panels. They compared the behaviours of identically designed riveted and welded panels, and determined shear strength and slip load capacity of the rivets considering different combinations of sheet thickness. They reported that the ultimate capacity of identically designed riveted and welded specimens was the same, and therefore concluded that the riveted connections are completely as satisfactory as are welded connections. Lennon et al. (1999) reported on a comparative investigation into the shear behaviour of some mechanical connections in CFS frames. Five different mechanical fasteners were considered, including rivet. They described that the most prominent failure mode for the 1.0, 1.2 and 1.6 mm steels is the sliding of the rivet head down the rivet shaft following local deformation, and for the 2.0 mm thick steel plate is the shearing of the rivet shaft. Porcaro et al. (2006) investigated the behaviour of riveted connection experimentally and numerically. They conducted experimental programme on elementary riveted joints in aluminium alloy AA6060 in two different tempers. They studied the influence of model parameters such as plate thickness, specimen geometry, material properties of the plates and loading conditions. They reported that residual stresses and the local changes in material properties due to the riveting process are important to verify the model. They also developed a finite element model to be able to describe the correct structural behaviour and thus the failure mechanisms of the riveted connections. He et al. (2008) reviewed the published works relating to self-pierce riveting (SPR) as a high-speed mechanical fastening technique for point joining of sheet materials. They illustrated the mechanics of joint formation and the types of defects that may occur. They also discussed on the main mechanical properties of the joints such as strength, corrosion properties and free vibration properties. In addition, they introduced the prediction of joint distortion when SPR is used to create. Moss and Mahendran (2003) investigated the structural behaviour of riveted connections used in steel framed housing. They tested 0.75 mm G300 steel to 1.15 mm G550 steel to find out the main failure modes of the connections, also to verify the accuracy of the available standards to predict the connection capacities. They observed that rivet tilting and bearing are the prominent failure modes of the connections. Mysak et al. (2014) studied the structural behaviour of blind riveted connections with nuts in the frames, experimentally. They investigated four different types of rivets including: combined aluminium-steel blind rivets, zinc-coated steel blind rivets, stainless steel blind rivets and blind nuts. They focused on the bearing capacity of the connections. They tested connections between elements with significantly different thicknesses, which take place in reality in the roofing of buildings, where trapezoidal sheet can be fastened to purlin by rivets. They reported that local deformations occur under rivet head in the thicker elements; and this phenomenon significantly affects the behaviour and bearing capacity of the connection. Mathieson et al. (2016) tested a pin-jointed truss connector to determine its strength, stiffness, and failure modes. They reported that the connection performed similarly to a bolted connection; and the failure modes were bearing and rivet shear failure, which mostly occur in CFS bolted connections. They also compared the connection strength in bearing and shear to existing formulae employed in AS/NZS4600 (2005) standard; and found close similarities.
The findings of the above review have provided indications of failure behaviours exhibited by test specimens and some of the factors which effect on these behaviours. There have also been criticisms on the strengths and weaknesses of the current design standards. The findings of this literature review will be taken into consideration when the results of the experimentation conducted for this research are evaluated.
3. Design guides and standards
The available standards and design guides, which currently govern CFS connection design, were analysed and compared in order to provide an outline of the various provisions which apply to the design of riveted CFS connections. This will also form the basis for comparison between the capacity recommendations of the design provisions and the connection capacities found through the experimentation. AS/NZS4600, Eurocode 3 (Eurocode3 2001) and the AISI standards (AISI-S100-16 2016) are the standards which set out provisions in the form of equations and rules which designers can apply to specific applications. In contrast, the NASH Handbook (NASH 2009) and the US Army Corps of Engineers Technical Instructions (TI 809-07 1998) provide assistance for the designer in the application of the design rules of AS/NZS4600 and AISI standards, respectively. They provide advice regarding the use of materials in particular applications, and give values for the capacities of materials commonly used in CFS construction. These values are based on experimentation and manufacturer recommendations.
Eurocode 3 specifies the same spacing and edge distance requirements as AISI, and recommends that all capacities be evaluated by testing, except for the bearing resistance capacity, which is specified for as outlined with the AISI bearing capacity provisions. AISI applies when the nominal rivet diameter (df) is between 3.0 and 7.0 mm, and Eurocode 3 applies for df between 2.6 and 6.4 mm. The AISI commentary recommends that at least two rivets should be used in any connection between individual components. It states that, as for screwed connections, this is to provide redundancy for installation inadequacies, and limit lap shear connection and distortion of flat unformed members like straps. It is also recommended that in connections between sheets of different thicknesses, the head of the rivet should be placed against the thinnest sheet where possible.
AS/NZS 4600 is predominantly based on the AISI Specification and as a result many of the provisions in these standards are the same. Eurocode 3 has similar provisions, with the greatest differences being with regard to the strength reductions which are applied to the various design capacities. The AISI Specification includes additional capacity factors for load and resistance factor design; however, only the factors for limit states design have been included in this outline in order to compare the specification with AS/NZS4600 and Eurocode 3, which both exclusively utilise limit states design.
Although the NASH Handbook and US Army Corps of Engineers Technical Instructions are relatively simple to use, and the tables included in them provide quick, reasonable indications of the capacities of certain designs, they are limited to a small range of materials, and often refer the designer back to design standards, or recommend experimental testing to confirm the design requirements.
4. Wall stud to plate connections
One of the commonly used CFS wall stud system in Australia and New Zealand is constructed of 90 mm C-section, which is formed from cold formed G550 steel sheet. A sample of the 90 mm C-section which was tested in this study is shown in Figure 1. The connections between individual wall studs and tracks (top, bottom and nogging tracks) are currently fastened with rivets through the two flanges of the wall members being connected. To increase the ultimate capacity of the framing system, the connections between the members should be improved where possible.
In this research study, a comparative evaluation was carried out regarding two types of riveted connection capacity that are currently used in wall framing system in Australia and New Zealand, i.e., type A and type O (see Figure 2). The principal difference between the two rivet types is that the type O rivets have a head diameter (dh) of 8.5 mm, while the type A rivets have a head diameter of 10.0 mm. The specimens of wall stud to track connections for both types of rivets at two different angles of 90[degrees] and 30[degrees] were fabricated. These angles are commonly used in the wall framing system, especially in K-braced CFS walls and truss framing system. Each different connection has been tested in both tension and compression. Three specimens of each connection were tested in order to gain a reliable indication of the capacity of each connection type. The specimens' name notation is illustrated in Figure 3.
The section structural material properties are presented in Table 1. It is necessary to mention that the CFS properties were obtained from the coupon tests (Standard 2005).
All wall stud to track connection specimens used for this study were assembled and tested in the Structural Laboratory of the School of Civil Engineering, the University of Queensland. It is necessary to mention that this study is part of a major research project (Zeynalian and Ronagh 2013; Zeynalian 2015; Zeynalian, Shelley, and Ronagh 2016; Zeynalian 2017) that is concerned with lateral seismic performance of different currently-in-use CFS frame configurations.
The test setup configuration is shown in Figure 4. At the top of the vertical member, two pieces of steel, which were the width of the stud section apart, were attached to the testing machine so that the vertical stud member could have three holes drilled through it and be bolted to the testing rig. A piece of timber was cut to the inner dimensions of the stud section and bolt holes were drilled through it; so that it could be placed inside the top of the vertical member at the point where it connected to the testing rig. The timber provided support for the stud section and prevented failure at the connection between the specimen and the testing rig.
It is necessary to mention that the testing machines were fitted with additional steel supports to which the angle stud connection could be bolted, as shown in Figure 5. In addition, the bolts were connected to the specimen through thick pieces of steel plate (see Figure 5), which was placed on the inside face of the non-vertical stud member in order to spread the concentrated load throughout the area of the plate and prevent failure of the CFS member at the bolts.
As mentioned earlier, three specimens of each type of wall stud to track connection were tested statically in tension and another three in static compression. This is the recommended minimum number of prototype specimens to be used in capacity testing, as recommended by the AISI. The specimen was loaded at a rate of 10 mm/min until failure occurred. The maximum load as well as load-displacement curve of each specimen was recorded. The outputs are in the Excel format, and can be used for the required post-experimental analyses. Schematic diagrams of these test specimens are presented in Figure 6.
5.1. Experimental results
A load-displacement graph was plotted for each test specimen. An example of the load-displacement graphs produced for each specimen is shown in Figure 7.
The results of the wall stud to track specimen tests including the maximum load capacity of each specimen are summarised in Table 2. The averages of the load capacities found in the three tests for each different type of specimen were calculated and also reported in the results table. The tension and compression capacities of the connections were compared and discussed with regard to the observed failure modes and the capacities of the different connections.
The mode of failure exhibited by the 30[degrees] and 90[degrees] specimens connected with the type O rivets in tension was the shearing failure of the rivet. The shearing occurred just below the head of the rivet, and occurred before any bearing failure was visible at the holes in either of the wall stud members. A specimen which failed in this way is shown in Figure 8.
This type of failure was also observed in the tension testing of type A rivets connected at 90[degrees], as shown in Figure 9. The 30[degrees]Connections with the type A rivets also failed by shearing of the rivet; however, they also exhibited significant bearing failure at the hole in the vertical wall stud member. This failure is shown in Figure 10.
The connections with type A rivets have significantly lower capacities in tension than the connections with type O rivets. For the 30[degrees]Connections, the type A rivets demonstrated 12% less capacity, and for the 90[degrees]Connections they had 23% less capacity. This indicated that the shear strength of the type A rivets is lower than that of the type O rivets.
The load-displacement curves in Figures 11 and 12 for the connections fastened with the type O and type A rivets, respectively, show that not only the type O rivets have a higher capacity, but they also exhibited longer displacements before failure.
In compression, all of the wall stud connections failed by buckling of the wall stud sections at the connection. One of the specimens fastened with type O rivets, which failed in this manner, is shown in Figure 13.
The connections fastened with the type A rivets buckled in both halfway between the connection and the testing rig bolts, and at the connection itself (see Figure 14).
The specimen in Figure 14 shows that there was some bearing at the bolt holes for the bolts which connected the specimen to the testing rig. This was found to occur in all of the compression tests on the wall stud specimens. This bearing was deemed not to have altered the maximum load capacity found for the specimens, but when the load-displacement plots are considered, it should be noted that the maximum displacement includes an error of approximately 2-3 mm.
The 90[degrees]Connections with type O rivets failed, as shown in Figure 15. The vertical member buckled near the connection, but in a different manner to the buckling at the connection for the 30[degrees] Connections. In the 90[degrees]Connections, the end of the vertical member flanges are flush against the inside of the horizontal member, which is most likely the cause of the different type of buckling which occurred in the vertical members of these connections. In contrast, in the 30[degrees]Connections, the vertical member was not flush against the non-vertical member; hence, it buckled by pushing out the flanges at the connection.
The results for the 90[degrees]Connections with the type A rivets in compression showed that the specimen WSA-C-5 had a lower capacity than the other two specimens of the same type by approximately 12%, which is a greater variation in capacity than was experienced by any other type of specimens. This specimen failed by section buckling away from the connection as well as buckling at the connection, similar to that exhibited by the 90[degrees]Connections with type O rivets. The other two specimens of the same type also exhibited buckling failure in these areas; however, rather than buckling inward, the flanges of the vertical member buckled outward, as shown in Figure 16. The inward section buckling in the WSA-C-5 specimen (see Figure 16) may have been the result of an imperfection in the steel sheet, or a weakness caused by slight bending or flexing of the section during transportation.
As for the connections tested in tension, the connections tested in compression demonstrated that the type O rivets had more capacity than the type A rivets. The type A riveted 90[degrees]Connections had 7% less capacity than the same connections with type O rivets, while for the 30[degrees]Connections, the type A rivets had 5% less capacity than the type O rivets. However, the capacity of the type O rivets is significantly lower in tension than in compression. The tension capacity of the 30[degrees]Connections is 36% less than the compression capacity of the same connection, while for the 90[degrees]Connections, the tension capacity is an even more significant 55% less than the compression capacity. While the compression capacity of the connections is governed by the section buckling capacity, rather than the fasteners themselves, the tension capacity of the connections is limited by the shear strength of the rivets. It would be advantageous for the design of the connections if tension capacity of the connections could be increased, so that it was closer the capacity of the connections in compression. This could be achieved by increasing the number of rivets used at each connection or by replacing the type O rivets with a different type of rivet which has a higher shear capacity.
The first possible improvement of increasing the number of rivets has some obvious problems including the increased fabrication time, which would result from increasing the number of rivets used at each connection. Another issue is that the connections between these wall stud members only have a limited area of overlap between the two sections, which would allow for two rivets to be placed at each side of the connection. The edge distance and centre-to-centre spacing requirements would not allow a second rivet of 4.8 mm diameter to fit in the available space on the current flange width of 34 mm. The second possible improvement of finding a higher shear strength rivet would be likely to increase the cost of the connections, and is dependent on the existence of higher shear strength rivets.
The design capacities of the connections were also calculated according to the design provisions and compared to the actual capacities in tension of the specimens, as found by experimentation. The outcomes are presented in Table 3.
As the ultimate failure mode of the wall stud connections in tension was governed by the shear strength of the rivets, the design capacities presented in Table 3 do not have a great deal of relevance. The codes and provisions recommend that the shear strength of the actual rivets is found by recommendation of the manufacturer or by testing. However, rare information from the manufacturers regarding the capacities of these rivets was available for the purpose of this research study. The design standards recommend that the capacity of the connection limited by the shear strength of the fastener should be taken as the shear strength divided by 1.25.
As the shear capacities of the rivets are unknown, the only comparison that can be made is between the design capacities themselves. The only equation provided by Eurocode 3 for the calculation of the connection capacity is for the provision of bearing resistance. The unfactored bearing resistances calculated with AISI and AS/ NZS4600 and Eurocode 3 are the same; however, Eurocode 3 applies a less conservative capacity reduction factor, which results in a higher factored design capacity than that provided by AISI and AS/NZS4600.
While AISI, AS/NZS4600, and Eurocode 3 predict that the unfactored bearing capacity of the connections is lower than the maximum load capacities of the test specimens, only the 30[degrees]Connections with type A rivets experienced significant bearing. The other connections which did not exhibit any bearing failure had maximum load capacities of between 18% and 60% greater than the unfactored bearing capacities predicted by the standards.
These results suggest that the design provisions of the code are not accurate and should not be solely relied upon when designing connections such as those tested for this study. Due to the imprecise nature of the design capacities found in Table 3, the recommendations made for these connections are based on the results of the experimentation alone.
6. Conclusions and recommendations
The paper presents the details of an experimental setup to test two common-in-use riveted wall stud connections, which are used in the housing industry in Australia and New Zealand, namely type A and type O. The tests reveal that a greater capacity in both tension and compression is obtained using the type O rivets. Therefore, the recommendation to designers is that they use the type O rivets with the 9.3 mm head.
In order to achieve a more efficient connection design, the tension capacity needs to be increased so that it is closer to the compressive capacity of the connection. The most feasible way of achieving this would be to use a rivet which had a greater shear capacity. However, this solution depends on the availability of higher-strength rivets and the magnitude of the associated increase to the connections as it may not be worth the expected higher cost of these rivets. The possibility of increasing the number of rivets per connection may be considered; however, the edge distance and spacing requirements which are required to prevent tear-out of the rivets could not be satisfied within the current flange width of the C-section used for the wall studs if more than one rivet was to be used at each flange. Increasing the flange width of the section could overcome this problem, but it would be a costly solution and therefore is unlikely to be a preferred option. It is necessary to mention that although the outcomes give some insights towards employing a better riveted CFS connection having higher capacity, more research studies are needed in order to make a more general conclusion.
Furthermore, the standard calculated design capacities of the wall stud connections show that the design codes are too conservative and can be improved considerably.
Received 24 August 2017
Accepted 10 June 2018
Nomenclatures dh rivet head diameter Fu ultimate stress Fy yield stress Abbreviations CFS cold-formed steel WSA wall stud connection with type A rivets WS wall stud connection with type O rivets T tension C compression RHF rivet head failure (popped-off) BR bearing at the rivet hole BFC buckling failure at the connection BFS buckling failure of the section away from the connection (m) moderate (<3 mm bearing) (s) severe ([greater than or equal to] 3 mm bearing)
No potential conflict of interest was reported by the authors.
Notes on contributors
Mehran Zeynalian, PhD, MSc, is Assistant professor of civil engineering at University of Isfahan, Iran. He received a master of science in civil engineering from Isfahan University of Technology in Iran in 2002. Dr Zeynalian also received his doctoral degree in the field of structures and construction management from The University of Queensland in 2012.
Hamid R Ronagh, Professor, is the director of the program on "Structures", appointed to the position in October, 2015. Prior to joining Western Sydney University, Professor Ronagh was at the University of Queensland where he led Structural Retrofit, the largest research group in the field of Structural Engineering. This group was active on many projects including FRP strengthening of reinforced concrete joints, overwrap repair of subsea pipelines and strengthening against post-earthquake fire. Hamid Ronagh is the author or co-author of over 200 scholarly publications (including over 100 journal publications) on a wide range of topics in the areas of structural engineering, in particular structural retrofit and light steel framing. He has also been involved with several major projects as a specialist consultant over the years. He received his PhD degree in Structures from the University of New South Wales in 1996.
Shahabeddin Hatami is Associate Professor of Structural Engineering at the Civil Engineering Department in Yasouj University, Iran. He received his Bachelor and MSc degrees from the Sharif University of Technology and his PhD degree from Isfahan University of Technology, Iran. He is the author of over 50 scholarly publications.
Mehran Zeynalian (iD) http://orcid.org/0000-0002-0681-2654
AISI-S100-16. 2016. North American Specification for the Design of Cold-Formed Steel Structural Members. Washington, DC: American Iron and Steel Institute.
AS/NZS4600. 2005. Cold-Formed Steel Structures. Sydney: Standards Australia.
Eurocode3. 2001. Design of Steel Structures. Part 1-3, General Rules--Supplementary Rules for Cold-Formed Members and Sheeting. Dublin: European Committee for Standardization..
Hancock, G. J., T. Murray, and D. S. Ellifrit. 2001. Cold-Formed Steel Structures to the AISI Specification. New York: CRC Press.
He, X., I. Pearson, and K. Young. 2008. "Self-Pierce Riveting for Sheet Materials: State of the Art." Journal of Materials Processing Technology 199 (1-3): 27-36. doi: 10.1016/j.jmatprotec.2007.10.071.
Lennon, R., R. Pedreschi, and B. Sinha. 1999. "Comparative Study of Some Mechanical Connections in Cold Formed Steel." Construction and Building Materials 13 (3): 109-116. doi:10.1016/S0950-0618(99)00018-5.
Mathieson, C., G. C. Clifton, and J. B. P. Lim. 2016. "Novel Pin-Jointed Connection for Cold-Formed Steel Trusses." Journal of Constructional Steel Research 116: 173-182. doi:10.1016/j.jcsr.2015.08.009.
Moss, S., and M. Mahendran. 2003. "Structural Behaviour of Self-Piercing Riveted Connections in G300 and G550 Thin Sheet Steels." In Advances in Structures, (Eds.) Hancock et al. Proceedings of the Int'l Conference on Advances in Structures, Sydney, Australia, edited by Hancock et al., 275-280, June 22-25.
Mysak, V., O. Tusnina, A. Danilov, and A. Tusnin. (2014), "The Features of Riveted Connections of Metal Elements", Proceedings of Moscow State University of Civil Engineering/Vestnik MGSU(3). 10.22227/1997-0935.2014.3.82-91
NASH. 2009. Standard for Residential and Low-Rise Steel Framing--Part 1: Design Criteria. Australia: National association of steel-framed housing.
Porcaro, R., A. Hanssen, M. Langseth, and A. Aalberg. 2006. "The Behaviour of a Self-Piercing Riveted Connection under Quasi-Static Loading Conditions." International Journal of Solids and Structures 43 (17): 5110-5131. doi:10.1016/j.ijsolstr.2005.10.006.
Rogers, C. A., and G. J. Hancock. 1999. "Screwed Connection Tests of Thin G550 and G300 Sheet Steels." Journal of Structural Engineering 125: 128. doi:10.1061/(ASCE)0733-9445(1999)125:2(128).
Standard, A. 2005. A370, Standard Test Methods and Definitions for Mechanical Testing of Steel Products. West Conshohocken, PA: ASTM International.
TI 809-07. 1998. Design of Cold-Formed Loadbearing Steel Systems and Masonry Veneer/Steel Stud Walls. USA: US Army Corps of Engineers, Engineering and Construction Division.
Yener, M., and R. N. White. 1984. "Cold-Formed Steel Panels with Riveted Connections." Journal of Structural Engineering 110 (5): 1035-1050. doi:10.1061/(ASCE)0733-9445(1984) 110:5(1035).
Zeynalian, M. 2015. "Numerical Study on Seismic Performance of Cold Formed Steel Sheathed Shear Walls." Advances in Structural Engineering 18 (11): 1819-1829. doi:10.1260/ 1369-43220.127.116.119.
Zeynalian, M. 2017. "Structural Performance of Cold-Formed Steel-Sheathed Shear Walls under Cyclic Loads." Australian Journal of Structural Engineering 18 (2): 113-124. doi:10.1080/13287982.2017.1349533.
Zeynalian, M., and H. Ronagh. 2013. "Experimental Study on Seismic Performance of Strap-Braced Cold-Formed Steel Shear Walls." Advances in Structural Engineering 16 (2): 245-257. doi:10.1260/ 1369-4318.104.22.168.
Zeynalian, M., A. Shelley, and H. R. Ronagh. 2016. "An Experimental Study into the Capacity of Cold-Formed Steel Truss Connections." Journal of Constructional Steel Research 127: 176-186. doi:10.1016/j.jcsr.2016.08.001.
Mehran Zeynalian (iD) (a), Hamid R Ronagh (b) and Shahabeddin Hatami (c)
(a) Department of Civil Engineering, The University of Isfahan, Isfahan, Iran; (b) Institute for Infrastructure Engineering, Western Sydney University, Sydney, Australia; (c) Department of Civil Engineering, Yasouj University, Yasouj, Iran
CONTACT Mehran Zeynalian ([mail]) firstname.lastname@example.org
Caption: Figure 1. Wall stud C-section.
Caption: Figure 2. Rivet types used in the testing.
Caption: Figure 3. The specimen name notation.
Caption: Figure 4. Testing machine setup.
Caption: Figure 5. Shortened test specimen.
Caption: Figure 6. Schematic diagram of the wall stud connections.
Caption: Figure 7. Load-displacement graph for a wall stud to track specimen.
Caption: Figure 8. Shear failure of type O rivet in tension.
Caption: Figure 9. Shear failure of type A rivet in tension.
Caption: Figure 10. Shear failure + bearing with type A rivets.
Caption: Figure 11. Load-displacement graph for wall stud connection with type O rivets.
Caption: Figure 12. Load-displacement graph for wall stud connection with type A rivets.
Caption: Figure 13. Buckling failure of the wall stud sections at the connection, type O riveted 30[degrees]Connection.
Caption: Figure 14. Buckling failure of the wall stud sections at the connection, type A riveted 30[degrees]Connection.
Caption: Figure 15. Buckling failure of the wall stud sections at the connection, type O riveted 90[degrees]Connection.
Caption: Figure 16. Buckling failure of the wall stud section, type A riveted 90[degrees]Connections.
Table 1. G550 material properties. Nominal Grade: 550 MPa Yield Strain: Nominal Thickness: 0.55 mm Ultimate Stress, Fu: Elastic Modulus: 168.9 GPa Ultimate Strain: Yield Stress, Fy: 592.3 MPa Fu/Fy: Nominal Grade: 0.45% Nominal Thickness: 617.25 MPa Elastic Modulus: 2.86% Yield Stress, Fy: 1.04 Table 2. Wall stud connection test results. Specimen Description Maximum Load Average Maximum Name (kN) Load (kN) Tension Tests WS-T-1 30[degrees] 6.04 6.36 wall stud to track connection WS-T-2 Type O Rivets 6.43 WS-T-3 6.61 WSA-T-1 30[degrees] 6.26 5.61 wall stud to track connection WSA-T-2 Type A Rivets 5.17 WSA-T-3 5.38 WS-T-4 90[degrees] 6.20 6.10 wall stud to track connection WS-T-5 Type O Rivets 5.80 WS-T-6 6.30 WSA-T-4 90[degrees] 4.50 4.67 wall stud to track connection WSA-T-5 Type A Rivets 5.10 WSA-T-6 4.40 Compression Tests WS-C-1 30[degrees] 9.59 9.95 wall stud to track connection WS-C-2 Type O Rivets 10.22 WS-C-3 10.04 WSA-C-1 30[degrees] 9.85 9.46 wall stud to track connection WSA-C-2 Type A Rivets 9.60 WSA-C-3 8.92 WS-C-4 90[degrees] 13.63 13.56 wall stud to track connection WS-C-5 Type O Rivets 13.98 WS-C-6 13.07 WSA-C-4 90[degrees] 13.03 12.61 wall stud to track connection WSA-C-5 Type A Rivets 11.60 WSA-C-6 13.21 Specimen Failure Name Description Tension Tests WS-T-1 RHF WS-T-2 RHF WS-T-3 RHF WSA-T-1 RHF+(s)BR WSA-T-2 RHF+(s)BR WSA-T-3 RHF+(s)BR WS-T-4 RHF WS-T-5 RHF WS-T-6 RHF+(m) BR WSA-T-4 RHF WSA-T-5 RHF WSA-T-6 RHF Compression Tests WS-C-1 BFC WS-C-2 BFC WS-C-3 BFC WSA-C-1 BFC WSA-C-2 BFC WSA-C-3 BFC WS-C-4 BFC WS-C-5 BFC WS-C-6 BFC WSA-C-4 BFC WSA-C-5 BFC + BFS WSA-C-6 BFC Table 3. Wall stud to track connection tension capacities. Tests (kN) Average Maximum Load - Specimen Name Description [P.sub.max] Failure Mode (kN) WS-T-(1-3) 30[degrees] 6.36 RHF wall stud to track connection Type O Rivets WSA-T-(1-3) 30[degrees] 5.61 RHF + (s)BR wall stud to track connection Type A Rivets WS-T-(4-6) 90[degrees] 6.1 RHF wall stud to track connection Type O Rivets WSA-T-(4-6) 90[degrees] 4.67 RHF wall stud to track connection Type A Rivets AISI & AS/NZS 4600 (kN) Net Section Tension Bearing Capacity Capacity [phi] [phi] Specimen Name [N.sub.t] [N.sub.t] [V.sub.b] [V.sub.b] WS-T-(1-3) 56.1 31.1 3.16 1.58 WSA-T-(1-3) 56.4 31 3.33 1.67 WS-T-(4-6) 56.1 31.1 3.16 1.58 WSA-T-(4-6) 56.4 31 3.33 1.67 AISI & AS/NZS 4600 (kN) Connection Shear Limiting Failure [phi] (Factored Capacity Specimen Name [V.sub.fv] [V.sub.fv] & Failure Mode) WS-T-(1-3) 3.95 2.37 1.58 BR WSA-T-(1-3) 4.17 2.5 1.67 BR WS-T-(4-6) 3.95 2.37 1.58 BR WSA-T-(4-6) 4.17 2.5 1.67 BR Eurocode (kN) Bearing Capacity [phi] Specimen Name [V.sub.b] [V.sub.b] WS-T-(1-3) 3.16 2.53 WSA-T-(1-3) 3.32 2.66 WS-T-(4-6) 3.16 2.53 WSA-T-(4-6) 3.32 2.66
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|Author:||Zeynalian, Mehran; Ronagh, Hamid R.; Hatami, Shahabeddin|
|Publication:||Australian Journal of Structural Engineering|
|Article Type:||Author abstract|
|Date:||Jul 1, 2018|
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