Chapter 6. Sensitivity Studies and Design Implications
6.4 Ductility considerations
Figure 6.4 Moment-displacement response for CFT members
Eurocode 4 [73] considers the full plastic moment (Mmax,Rd) based on the steel section.
The confined/unconfined strength of concrete was also depicted in the above figures.
It was found that the predictions are generally similar to the numerical results within the practical range of member sizes and material properties.
6.4 Ductility considerations
6.4.1 Ductility of PEC members
It is necessary to consider limiting criteria for ductility for the purpose of seismic assessment and design. From the investigation on the tests and previous numerical simulations of PEC members, the onset of local buckling was typically followed by a reduction in load capacity, particularly in the presence of larger axial loads. Therefore, it seems plausible to consider local buckling as a possible limiting criterion for PEC members. On the other hand, in the case of CFT members, it has been found that local buckling would be a more conservative limiting criterion.
In the inelastic range, local flange or tube buckling could be assessed in terms of the critical strain corresponding to its onset in the inelastic range. A reasonable treatment of this complex problem can be obtained by combining fundamental plate buckling
99
theories with experimental results [33]. It has also been suggested that the relationship between critical buckling strain (εcr) and yield strain (εy) can be expressed as [33]:
2
Where ν is the Poisson’s ratio, E is the modulus of elasticity, σy is the yield stress, t is the thickness of flange or square box, c is the flange outstand or box width, and k is the buckling coefficient. The value of k is a function of plate geometry and boundary conditions. The variation in k with the aspect ratio (a/b) for plates with different edge conditions is shown in Figure 6.5 [74].
Figure 6.5 Buckling coefficient versus aspect ratio [74]
Equation 6.1 could also be rearranged and presented in terms of the commonly used plate slenderness parameter defined as:
cr
Another suggested relationship between εcr/εy and λp which was based on experimental investigations on the local buckling of the flange can be expressed as [27]: respectively. The results from the experimental specimens and numerical assessment are depicted in Figure 6.6 and 6.7 respectively, according to the definition of Eurocode 4 [73]. The details of steel profiles used in finite element model, including PEC members and CFT members are shown in Table 6.1 and Table 6.2.
Figure 6.6 depicts the influence of flange slenderness on normalized critical strain of PEC members. Only the experimental results from IC1, IC6 and IC7 were presented in figures because the other specimens failed without local buckling due to stress concentration around the holes employed for placing the transverse bars. It has been found in Chapter 4 that the strain gauges location coincided with the buckled flange or box. It is evident from Figure 6.6 that Equation 6.3 provides a closer prediction of the critical buckling strain obtained from both the experiments and finite element estimations in the case of PEC members while Equation 6.2 seems to be conservative.
Table 6.1 Details of steel profiles of PEC members used in Figure 6.6
No. 1 2 3 4 5 6 7 8 Configuration
of members HEA240 HEA200 HEA240 HEA200 450×450 450×450 450×450 450×450 Slenderness of
flange 9.7 9.7 9.7 9.7 23.2 23.2 23.3 35.4
Details of
links 8@96 8@80 No links No links 8@225 [email protected] 8@450 8@450
101 Figure 6.6 Influence of flange slenderness on normalized critical strain for PEC
members
As discussed in previous chapters, the additional straight bars welded to the flanges within the potential plastic hinge zone could delay local buckling. It could also be observed from Figure 6.6 that the ductility of PEC members with links could be increased. A simplified approach to consider the effect of the additional bars could be undertaken by reconsidering Curve D in Figure 6.5. The aspect ratio ‘a/b’ would represent the ratio between the length of the local buckling and the outstand width.
The value for ‘a’ and ‘b’ obtained in IC1 was about 140 mm and 75 mm, respectively.
In Figure 6.5, the buckling coefficient k corresponding to a/b equalling to 1.9 is about 1.3. The additional bars were effective in limiting the length of the buckle to the spacing between the links. This means that the additional bars can be considered as an approach of reducing the ratio ‘a/b’ by limiting the value of ‘a’ to the distance between the links which would lead to higher values for the buckling coefficient ‘k’.
In the experiments, the spacing of the welded bars was either 40 mm or 60 mm.
Hence, ratio ‘a/b’ was 0.53 or 0.8, respectively. Compared to conventional specimen, the enhancement in ductility was at least 361% and 160% for PEC members where the aspect ratio was equal to 0.53 and 0.8, respectively, if no premature weld and transverse bar failure happened. Based on the above discussions, significant benefit in
ductility would occur if a/b is smaller than 1.0. As discussed in Chapter 4, premature weld or link failure limited the effectiveness of the additional bars in some tests, especially when the distance between two links was small, being equal to 40 mm. On the other hand, for IC2 and IC3, it was evident that the transverse bars were not fully effective because the flange around bars fractured earlier. However, the use of 8 mm by 6 mm transverse plate with an improved welding procedure provided a considerable improvement in performance.
6.4.2 Ductility of CFT members
Table 6.2 lists the details of CFT members which are depicted in Figure 6.7. Previous studies have suggested that a buckling coefficient k of 10.3 can be employed to determine the slenderness of column cross-sections [75]. In Figure 6.7, the grey squares gave the critical strain corresponding to local buckling, while the red squares represented the critical strain according to flange fracture. Figure 6.7 also presents the results related to the cross-section classification of Eurocode 3 [61]. In the case of CFT members, Equation 6.2 was still over-conservative for determining the buckling strain. It was suggested that Equation 6.3 should be applied to determine the buckling strain of CFT members. Furthermore, it was also noted from Figure 6.7 that in the case of SY0, SY1 and SY2, members failed in thirteen cycles, sixteen cycles and thirteen cycles, respectively, since local buckling occurred. Hence in the case of CFT members, it was conservative to consider local buckling as the limiting criterion.
Table 6.2 Details of steel profiles of CFT members used in Figure 6.7
No. 1 2 3 4 5 6 7 8 Configuration
of members (mm×mm)
180×6.3 180×12.5 180×16 250×8 250×16 300×8 300×12 300×20
103 Figure 6.7 Influence of flange slenderness on normalized critical strain for CFT
members
6.5 Concluding remarks
The experimental and numerical studies offered an assessment of key response parameters and design consideration. In terms of the stiffness of PEC members with lower axial loads, the estimated experimental value was largely in agreement with the assumption based on 20% of the concrete contribution. In terms of local buckling of PEC members, the critical strain could be predicted from simple equations dependent on the normalised plate slenderness parameters. The enhancement of ductility by delaying local buckling through the provision of additional links could be captured by a corresponding enhancement in the buckling coefficient. However, the ductility of CFT members based on the local buckling of the steel tube was over-conservative.