EFFECTS ON PREVIOUS FAR-FIELD FINDINGS
5.3.4 F RAMEWORK OF THE IDA S TUDY
5.4.2.5 Structural Dynamic Magnification factors of Storey Accelerations
The values of ) ( 1 90 T S A a r th
are provided in Table 5-12, which are the dynamic magnification
ratios of maximum storey accelerations when compared with their structural spectral acceleration.
These values can be transformed to
PGA Ar90th
by multiplying them by the ratio of the suite’s median
spectral accelerations at Sa(T1=2s) and PGA,
PGA T Sa( 1)
= 0.833. Table 5-12 manifests values of Ar90th
normalised by Sa(T1=2s) and PGA that were markedly high; generally > 5×Sa(T1), and > 4×PGA.
They exhibited maximum values at elastic behaviour and gradually decreased with damage-
induced structural softening up to LS (≈ 0.3g) performance level. Afterwards, they increased again as structural collapse was inevitable. These magnifying factors can be utilised in calculating the
90th percentile value of maximum storey acceleration by multiplying them by the structural
Table 5-13 lists dynamic magnification ratios of maximum storey accelerations for the far-
field study. Comparison of
PGA Ar90th
values show lower magnitudes for the near-field study up to LS
level. At CP and higher levels, near-field values markedly exceed those of the far-field study
following the trend of θr90th. This suggests that collapse onset occurs at the level/location of
maximum θr90th.
Table 5-16 presents median values of peak storey accelerations at different damage levels
for non-structural components, Kircher et al (1997A). These values incorporated factors for
appropriate fixation and anchorage. On comparing values of Table 5-16 with Ar90th in Table 5-
12; one can establish that at DBE (0.88g) a near extensive damage of acceleration-sensitive
contents and non-structural components is expected, which is better performance than that of the
far-field study. Likewise and at MCE and up to CP near-field structural performance is better
than that of the far-field. For levels higher than CP, near-field damage exponentially increases
with severe adverse consequences.
5.4.3 STOREYS’MAXIMUM RESPONSE PROFILES
Seeking completeness of depicting structural response, the results of storeys’ maxima are
presented for the IM levels of MCE and 3DBE. For the complete set of θ, δ and A figures, refer to Appendix A2 – Figures A2-1 to -15 for radial, X- and Y-axes plots of θ, δ and A for Tx/Ty =
{1.0, 1.5, 2.0, 2.5, 3.0} at Elastic behaviour, DBE, MCE and 3DBE. Likewise, Figures A2-16 to
-30 provide the IM normalised ones. As with Chapter 4 and on inspecting these plots of θ, δ and A, MCE was found to be a credible representative of behaviour IM levels of elastic, DBE and
MCE. Consequently, figures for MCE (response) and 3DBE (collapse onset) are only displayed
in this chapter.
Figure 5-10 displays the medians and 90th percentiles of θr/IM, θx /IM and θy/IM at IM =
Sa(T1=2s) = Sa(MCE). Figure 5-10 (a, c and e) show typical interstorey median responses for
upper mid-section where typically the bisymmetrical structure (Tx/Ty =1) exhibit maximum
values for radial and Y-directions. Figure 5-10 (c) depicts the responses of the weak X-axis
where they all nearly coincide as X-axis maintained its capacity properties for the set of
considered structures. Figure 5-10 (b and d) and (f) [for the bisymmetrical structure] display the
90th percentile code-maximum behaviour when the structure’s mid-section exhibit degrading
strength. Subsequently interstorey maxima shift upwards where stabilising normal forces are
somewhat lower, as seen when comparing Figures 5-11 (a) [median demands] and (b) [higher
90th percentile demands]. Nevertheless, Figure 5-10 (f) show that in the strong direction the
structure retained its undegraded character of maximum interstorey drift values at mid-height,
except for the more flexible bisymmetrical structure. In essence, maximum interstorey drifts
occur in the weak direction of the structure in its upper storeys that subsequently affect the radial
(vectorial) total. Consequently, interstorey drifts of the strong direction slightly shifts upwards
due to degradation of columns in the orthogonal (weak) axis but stay within the mid-height
section of the structure. In contrast with the far-field study, the imminent collapse, due to the
loss of lateral structural strength, occurs in the mid-height section (particularly about the mid-
height) of the considered structure except for the bisymmetrical case; as confirmed by Figures 5-
11 (b, d and f) at 3DBE (collapse onset) level. The figures suggest that the collapse is shown by
a large flat increase in a brittle manner. This can be ascribed to the local failure of columns due
to higher mode (vibration) effects that rapidly degrade the already degraded elements to a
“complete” collapse. On the other hand, the more flexible bisymmetrical structure follows the
far-field trend the onset of collapse starts in the upper storeys. Furthermore, Figure 5-11
reported markedly high standard deviation values, particularly in the mid-height storeys, thus
confirming the collapse is taking place in the upper storeys.
Figures 5-12 and 5-13 exhibit the medians and 90th percentile values of radial, X- and Y-
axes lateral storey drifts at MCE and 3DBE levels respectively. Both figures display the same
structures approached incipient collapse. Once again, the bisymmetrical structure expectedly
showed the highest magnitudes. Predominantly, the values of lateral storey drifts were, more or
less, linear with maxima occurring at roof level, except for the unisymmetrical structure (Tx/Ty
=3) at collapse onset. Thus, values of lateral drifts to each storey can be linearly assigned,
starting with zero at ground level and ending with the maximum at the roof level. This can be
achieved via implementation of equations (4.6) [of Chapter 4] and either of (5.4) or (5.5) to
evaluate the maximum lateral storey drift of the studied structure. In essence, the results suggest
marked resemblance between the near- and far-field response albeit of different magnitudes.
Figures 5-14 and 5-15 show the medians and 90th percentiles of radial, X- and Y-axes
storey accelerations at MCE and 3DBE levels respectively. For MCE or less, Figure 5-14 shows
that maximum values exist at the top level, while storey accelerations increase linearly from
level 1 up to about mid-height then almost maintain the same value for the remaining storeys up
to the storey below the roof. The median of this maintained value for radial storey acceleration
was 64% of the roof acceleration, 65% for X- (weak) and 64% for Y-accelerations (strong)
respectively, as listed in Table 5-17. On studying (conservatively) the curves and their data, it is
suggested to use the following:
Aroof = A90th = a Sabγbiaxial,Aε (5.6) Aiu = 0.69 Aroof (5.7) Ail = Aroof m i 69 . 0 94 . 0 ⎟ ⎠ ⎞ ⎜ ⎝ ⎛ (5.8) A0 = PGA (5.9)
where A0 is the ground storey acceleration, Ail is the lower storey accelerations up to the mid-height
of building, Aiu is the median value of storey accelerations for the remaining upper storeys except
roof, i is the storey number, m = integer(n/2) – 1 , n = number of building storeys, Aroof is the roof
acceleration. Again, structures with deep foundations or those that have multi-storey deep
unreliable. This resembles the behaviour reported in the far-field study albeit of lesser magnitudes.
On the other hand and when inspecting Figures 5-15 (b) and (d), storey accelerations show a very
large flat increase in the mid-height section of the considered structure following the trend of
interstorey drifts at collapse, except for the more flexible bisymmetrical structure. This is a stark
difference with the far-field pattern showing maximum accelerations only occur at the roof level
even at collapse.
Again, the reported overall median values of γ-percent, previously investigated in Chapter 3, for the 90th percentile code demands of θ, δ and A were generally above 35%. Moreover, individual Tx/Ty median values of γ-percent for θ, δ and A demands peaked to as high as 39.8%, thus
vindicating the suggested use of 40-percent rule in combining the bidirectional X and Y demands.
5.5 IMPLEMENTATION INTO DCFD FORMAT AND FRAGILITY CURVES
In keeping with Chapter 4, capacity and dispersions values for the far-field are compared
with the near-field results, except for βDR that are calculated from the near-field demand values
plus values of b and γbiaxial,D developed in this chapter. This decision is made to facilitate
appropriate comparison between near- and far-field demands without being affected by other
factors. Consequently value of βor is taken as in the far-field study βor = 0, also relying on the
previously-used assumption that the stronger components applied to (weak) X-axis represent the
fault-normal ones thus eliminating the need for βor. Values of βDR were calculated from values
of the median and 90th percentile demands and are listed in Table 5-14. When comparing Tables
5-14 and 5-15 (far-field), values of βDR of the near-field study are markedly less than those of the
far-field one, ascribed to relatively lower βEQ of the near-field.
Consequently and as in Chapter 4, values of the correction factors of demand (γ) and capacity (φ) are evaluated for θr, δr and Ar, and listed respectively in Tables 5-18, 5-19 and 5-20.
Accordingly, values factored demands (γD) and factored capacities (φC) are computed and used to calculate the confidence factor λ=γD / φC. Subsequently values of λ are used to compute the
Gaussian confidence variate KX = KX =½k βUT – ln(λ)/bβUT that is used to obtain the confidence
level of non-exceedance CL% = Φ (KX). These values are listed in Tables 5-18, 5-19 and 5-20.
On comparing values of CL% of the near- and far-field studies, the considered structures
under the near-field suite achieved higher values of CL% inferring higher confidence in
achieving the studied performance levels of IO, LS and CP, particularly in roof accelerations.
This finding can be attributed to the relatively lower magnitudes of θr, δr and Ar for the near-field
study when compared with the far-field study. This outcome can be also seen in Figure 5-9
where it displays lower θr values for the near-field study up to CP performance level.
Endeavouring completeness of this comparative study and as in Chapter 4, fragility
curves for the near-field study were generated and plotted (thick solid lines) in Figure 5-17
among the far-field fragilities (thin dashed lines) to facilitate comparison. Figure 5-17 show
(visually) less fragility values for the near-field study than those of the far-field. For example,
Figure 5-17 shows that the hazard intensity of Sa(T1) = 0.3g has 19% and 81% probability of
(failure) exceedance of expected θr demands respectively at LS and CP performance levels which
are significantly better than the fragility values of 28% and 87% at LS and CP for the far-field
study. In other words, this indicates that the structural performance under the near-field suite is
generally better than under the far-field suite.
5.6 CONCLUSIONS AND RECOMMENDATIONS
This chapter has investigated the response and behaviour of three-dimensional 2-second-
period 12-storey structures by means of an IDA method under the previously-developed near-
field earthquake suite. A comparative study was conducted and the drawn conclusions can be
summarised as follows:
1. Values of the log-log regression coefficients a and b for the near-field were not that
different from the far-field suite. This suggests universal applicability of the expression
2. There was a marked drop in values of the amplification factors of the biaxial interaction
for structure with Tx/Ty≤ 1.5 for the interstorey and storey drifts. Almost no change in
values of storey accelerations was recorded for these structures when compared with the
overall median of all considered structures.
3. Overall the structural response and demands were less than for this near-field study than
those of the far-field study up to collapse prevention (CP) performance level. This
reduction was particularly large in storey accelerations. Consequently, the dynamic
magnification factors for storey accelerations dropped significantly to 5×Sa(T1), and >
4×PGA, about half those of the far-field suite. However and upon exceeding the CP
level, the response and demands of the near-field study increase in a brittle explosive-like
fashion, particularly for interstorey drifts and storey accelerations.
4. As with the far-field study, the structure degrades around its mid-height section and
subsequently the degradation shifts slightly upwards. However, and with the exception
of the bisymmetrical structure, the considered structures exhibit a blow-out increase in
interstorey drifts in the degraded mid-height section that causes a sudden explosive-like
failure accompanied with extremely high magnitudes in storey accelerations.
5. Values of storey accelerations for the upper half of the considered building would result
in moderate damage to contents and non-structural components at DBE levels and
extensive damage is anticipated at MCE, which results in better performance than the far-
field response.
6. Collapse displayed by 90th percentile demands at 3DBE suggests high mode effects
failure that can be attributed to severe degrading vibrations of the structural members
with reduced strength in the mid-height section of the considered structures. One can
postulate that this damage starts at the onset of strength degradation markedly increases
when plastic mechanisms are formed, thus allowing for larger vibration amplitudes. This
explosive-like nature. Nevertheless, this higher mode induced damage only progresses to
soft-storey failure when the seismic intensity exceeds the MCE level for seismically-
designed structures.
5.7 REFERENCES
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Table 5-1. Comparative values of Spectral accelerations (Sa) for Near- and Far-Field Suites.
Period T Near-Field Suite Sa (g) Far-Field Suite Sa (g) (sec) Medians Maxima βEQ Medians Maxima βEQ
0.00 (PGA) 0.551 1.226 0.363 0.382 0.883 0.618 0.25 1.007 2.123 0.414 0.939 2.285 0.673 0.50 1.025 2.455 0.439 0.822 1.761 0.705 0.60 1.000 1.932 0.387 0.725 1.239 0.691 1.00 0.873 2.269 0.449 0.483 1.008 0.774 1.50 0.598 1.278 0.392 0.306 0.511 0.659 2.00 (T1) 0.459 0.840 0.408 0.206 0.380 0.646 2.50 0.347 0.755 0.431 0.164 0.250 0.674 3.00 0.258 0.570 0.486 0.126 0.337 0.718 Overall Median for T= 0–2.5s 0.690 1.381 0.415 0.382 0.612 0.681 Sa(2s)/PGA 0.833 0.539
Table 5-2. Comparative values of Spectral accelerations (Sa) for Near- and Far-Field Suites
Normalised by Sa(T1=2s).
Period T Near-Field Suite Sa (g) Far-Field Suite Sa (g) (sec) Medians Maxima βEQ Medians Maxima βEQ
0.00 (PGA) 1.202 2.673 0.363 1.849 4.279 0.618 0.25 2.195 4.628 0.414 4.549 11.070 0.673 0.50 2.234 5.352 0.439 3.984 8.532 0.705 0.60 2.239 6.695 0.436 2.969 5.198 0.762 1.00 1.903 4.947 0.449 2.339 4.884 0.774 1.50 1.304 2.786 0.392 1.483 2.474 0.659 2.00 (T1) 1.000 1.831 0.408 1.000 1.842 0.646 2.50 0.757 1.647 0.431 0.794 1.213 0.674 3.00 0.563 1.242 0.486 0.609 1.634 0.718 Overall Median for T= 0–2.5s 1.504 3.011 0.415 1.852 2.965 0.681
Table 5-3. Increases in Median Responses and 90th percentile Demands of aθ = θr/IM due to
Biaxial interaction for Near- and Far-Field Suites. Median Responses
aθr = θr/IM
90th percentile Demands
aθr90th = θr90th/IM