Limitations of Simplified Analyses ...

Limitations of Simplified Methods for Evaluating Liquefaction,

Settlement and Lateral Spreading in Earthquakes

?by Robert Pyke Ph.D., G.E.

June 18, 2023

The first step in any evaluation of liquefaction and its consequences should be studies of the regional geology and seismicity, and answering the question “is there any evidence of earthquake-induced liquefaction and settlement or lateral spreading of similar soils in a similar tectonic environment? See Pyke (1995, 2003, 2015) and Semple (2013). While there are rare instances of liquefaction being reported in Pleistocene age sands, the vast majority of well-documented case histories have occurred in geologically recent cohesionless soils and man-made fills, particularly hydraulically placed fills.

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The widespread belief that “one has to show a calculation” tends not to promote better geotechnical engineering practice but rather worse practice. A good screening analysis should emphasize common-sense and experience. For example, it did not require any analyses to conclude that the Marina District in San Francisco was susceptible to liquefaction prior to the 1989 Loma Prieta earthquake, or that the areas with recent alluvium along the Avon River in Christchurch NZ were susceptible to liquefaction prior to the 2010-11 Christchurch earthquakes.

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In conducting simplified analyses, all soil layers that have “clayey” descriptors should be tested as necessary to confirm that description and should be excluded. Also, soils that are older than several thousand years or are known to be over-consolidated should be excluded. See for instance Arango et al. (2000). Sand “layers” seen in individual borings or soundings that are shown to be discontinuous by adjacent borings or soundings should also be excluded. See Pyke (1995) for further discussion of this point. Layers with a measured shear wave velocity exceeding 720 fps should also be excluded. See Andrus et al. (2009). And, if measured shear wave velocities are not available, layers that would be otherwise characterized as “dense” or “very dense” should also be excluded. Layers that are characterized as “dense” may still be susceptible to excess pore pressure development under strong shaking, but few if any adverse consequences of liquefaction have been observed when the normalized clean sand SPT blowcount exceeds 15. See for instance Ishihara (1993), Youd et al. (2002). Finally, a clayey or otherwise non-liquefiable crust overlying a potentially liquefiable layer will limit any adverse consequences of liquefaction. See for instance Ishihara (1985) and Youd and Garris (1995). While lateral spreads triggered by liquefaction of layers under a crust as thick as 10 m are not totally unknown, they are limited to very large earthquakes such as have occurred off the coasts of Alaska, Chile and Japan.

This technical note only addresses liquefaction, settlement and lateral spreading in earthquakes under a horizontal or gently sloping ground surface. Simplified procedures for estimating displacements of steep slopes also have shortcomings, but these are not addressed in this note.

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More specifically the following seven factors limit the accuracy of simplified methods of analysis and tend to make the results very conservative:

?1.?????The “cyclic stress ratio” computed working backwards from the peak ground surface acceleration is invariably much greater than that computed using a nonlinear effective stress site response analysis which takes into account the specific layering and properties of the site. Pyke (2019) and Crawford et al. (2019) show examples of the comparison of the cyclic stress ratios computed using the simplified and more accurate methods and found that the cyclic stress ratio computed using site response analyses can be as low as one half that computed using the simplified method. This reduction also results from the fact that triggering of liquefaction in one layer reduces the amplitudes of the ground motion above that layer.

?2.????The correction of penetration resistance measured using either the SPT or the CPT for fines content is generally inadequate. Fines, and especially clayey fines, can dramatically reduce the penetration resistance while improving the behavior under cyclic loadings.

?3.????The correction of penetration resistance measured using either the SPT or the CPT for “thin layer” and “transition” effects is also generally inadequate. Recent work by Boulanger and DeJong (2018) and Yost et al. (2021) provides some guidance on these effects, but they are rarely accounted for in practice.

?4.????Penetration resistance also does not fully reflect the effect of aging, including pre-straining, and over-consolidation on improving the soil response to cyclic loading.

?5.????There is growing evidence that the cyclic stress ratios that trigger liquefaction in a given number of cycles that are obtained as a function of the penetration resistance from the standard charts that are used in simplified methods may be overly conservative. See for instance Stokoe (2023), whose team conducted relatively well-controlled studies of silty sands following the earthquakes in Christchurch NZ. By comparison, the case histories on which the standard curves are based are not so well documented.

?6.????There is also a growing consensus that the occurrence of liquefaction can only be understood by conducting nonlinear effective stress analyses in which excess pore pressure development and dissipation is tracked.?See Ntritsos et al. (2018), Cubrinovski (2019), Hutabarat and Bray (2019), Kramer (2019), Olson et al. (2020) and Cubrinovski and Ntritsos (2023). It further turns out that such analyses, which account for the re-distribution and dissipation of excess pore pressures, are needed to properly understand the case histories on which the simplified, empirical methods are based.

?7.????Analyses based on a single boring or CPT may include what are logged as “layers” but are in fact lenses or portions of a sinuous backfilled channel. Such lenses or segments of a channel cannot respond to earthquake ground motions independently as is assumed in simplified analyses or even in standard one-dimensional site response analyses. If such “layers” do tend to develop excess pore pressures, they become “soft inclusions” and the shear strains and hence the shear stresses are controlled by the response of the surrounding materials (again, see Pyke (1995)).

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The last of the issues listed above can now be simulated in my own site response analysis program, TESS2, by applying the displacement histories generated by running the site response calculation representing the site as a whole to a column that includes the potentially liquefiable “layer”. This level of effort is not justified for all projects as knowledge of the depositional environment and common-sense should be sufficient on most projects, but where regulators require a “calculation”, it is now possible to provide a meaningful calculation.

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I might never have finished writing this note were I to spell out all the particular limitations of individual simplified methods of analysis, and you would never read them, but I will just note that there are many of them and it is not easy for practicing engineers to be fully cognizant of them or to evaluate their impact on any given analysis. These limitations include the applicability of correlations of properties to penetration resistance measured in calibration chamber tests on “baby” sands, that is freshly placed, clean washed sands, and other correlations involving the shear modulus and penetration resistance. It is important that practicing engineers understand that not everything that is published in a journal is correct for every situation, and that some methods of simplified analysis may never be correct!

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Simplified methods for estimating settlements caused by earthquake loadings are particularly bad. For sands above the water table, this results in part from the use of inappropriate laboratory test results. Neither Harry Seed nor I thought that settlements of non-saturated sands was a major problem, but I nonetheless completed my Ph.D. thesis on this subject for other reasons, including achieving a better general understanding of the behavior of sands under cyclic loadings, as explained in Pyke (2022). For sands below the water table, this excessive conservatism results not only from the use of Ishihara and Yoshimine (1992), which was later noted by Ishihara et al. (2016) to be conservative, but even more because simplified analyses often suggest that a great depth of sand may reach the point of initial liquefaction and trigger larger settlements on reconsolidation, whereas a nonlinear effective stress analysis will indicate a much more limited depth of liquefaction. See Crawford et al (2019) and Pyke (2020) for further observations.

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The relationships for lateral spreading developed by Les Youd and his colleagues - Bartlett and Youd (1995) and Youd et al. (2002) are the best of the simplified methods, if applied with respect for the geology of the site as illustrated in Youd (2018). But even so, there is quite large scatter in the empirical data and a more advanced analysis, such as is described in Pyke (2019b), may be justified on more critical or high value projects. The method of Zhang et al. (2004) makes no sense at all as it relies on a correlation between cyclic strains in tests conducted without initial shear stresses and permanent strains developed in the field where there are initial shear stresses, and it should not be used.

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It is for all these reasons that Dr. Peter Robertson, the technical advisor to the developer of the program C-LIQ, has suggested to me that C-LIQ should be used primarily as a screening tool and should not be used to obtain the final answer on “high value” projects.

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A good example of the limitations of simplified analyses of liquefaction and its effects is provided by Jon Bray in his 2022 Seed Lecture. Much of that lecture is based on studies conducted in Christchurch, New Zealand, following the 2010 and 2011 earthquakes. Most of those studies were conducted along the Avon River whose floodplain is underlain by recent alluvial deposits. While of some academic interest, those studies did not add a great deal of practical value because many of these alluvial deposits were clear examples of soils that would be susceptible to liquefaction in even moderate earthquakes. However, there were some sites that showed no surface manifestations of liquefaction, and in particular no ejecta.

The sites that did show surface effects of liquefaction, which are shown in red, were found to contain thick deposits of clean sands that had a typical Ic, the soil behavior index obtained from cone penetration tests, of consistently less than 1.8. The sites that did not show surface effects of liquefaction, shown in blue, may have contained some cleaner sand layers but were generally finer grained and typically had Ic values greater than 1.8. The CPT traces in blue in the above figure were obtained on sites which showed no surface effects or ejecta. As Professor Bray correctly says in his lecture, it all goes back to geology. The sites that generally had Ic values greater than 1.8 had been deposited in back-water swamps. While these soils were often characterized as silty sands, they in fact tended to have clean sand layers separated by silt or even clay layers. For some years I have advised clients that they should not rely on simplified analyses if the values of Ic generally exceed 2.05, the generally accepted upper limit for clean sands (see Pyke and North (2019)). This does not necessarily mean that there is no problem, but it does mean that you should not automatically accept the results of simplified analyses and that more detailed site investigations, and more advanced analyses, are called for on high-value or critical projects. Further support for Bray’s limiting value of Ic of 1.8 for classic liquefaction with surface effects was provided by Robertson and Wride (1998) who found that the need to correct penetration resistances to “clean sand” values started at an Ic value of 1.7, and by the case history described by Boulanger et al. (2016), so that applying a limit on Ic of 2.05 in simplified analyses may be too conservative.

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The following figure shows a typical pattern of SPT blowcounts interpreted from the results of CPTs in the San Francisco Bay Area where the Bayshore deposits typically result from overlapping alluvial fans with occasional creek and levee deposits cutting through them. This results in lenses or pockets of cleaner sands within layers of more silty and clayey materials that are not continuous over a wide area. Such deposits have not exhibited historic liquefaction, but simplified methods often predict liquefaction and significant seismic settlements.

The figure shows that below an Ic value of about 1.8 the interpreted SPT blowcounts are between about 15 and 30. These materials, which are cleaner sands, may develop some excess pore pressures under cyclic loading but are most unlikely to produce consequences of engineering significance – as noted previously, engineering consequences are unlikely above a clean sand blowcount of 15. Beyond Ic values of about 1.8, the interpreted SPT blowcounts fall off as a function of the Ic value but this reflects increasing fines content and lower penetration resistance rather than increased susceptibility to adverse performance under cyclic loadings.

Additional factors that should be considered, even in screening analyses, include greater emphasis on the method of deposition, and the significance of complete saturation. The method of deposition is critical for both natural and man-made deposits. As illustrated in Pyke (1973) and in basic texts on sedimentary geology, a denser rain of particles leads to looser packing and lower densities. Conversely a less dense rain of particles leads to tighter packing. Some of the effects of this are reflected in penetration resistance, but not all of the effects. The impact of the method of deposition can be seen for instance in the difference in the behavior of older hydraulic fills, which were largely dumped, or otherwise deposited with a dense rain of particles, compared with newer hydraulically placed fills in which the placement is more distributed and there is some compactive effort applied by spreading and grading operations using heavy equipment. See Pyke et al. (1978) for an example. ?Also, it should not be assumed that sands below the current water table are “fully saturated”. In the laboratory it takes quite high back pressures to obtain a B factor approaching 1.0. This can easily be checked in the field by measuring the compression wave velocity – see Stokoe (2023) for an example. And see Banister et al. (1976) for an example where simplified analyses had predicted liquefaction in a large-scale field experiment, but liquefaction did not occur. This was attributed in part at least to seasonal fluctuations in the water table and the need for longer-term saturation, ideally with some flow of water, to obtain complete saturation.

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Finally, it should be noted that the academic criteria for delineating soils that are susceptible to liquefaction are likely very conservative. These criteria define soils that might develop significant excess pore pressures under the worse conditions such as those under buildings constructed on shallow foundations on loose sands or silts with a high water table, when rocking of the building can add to the cyclic shear stresses and strains in the soil. Figure 3 was sent to me by Jim Gingery of Keller North America in response to a question that I asked during his recent presentation to CalGeo.

Jim’s accompanying text said: “There is a domain of soil types that are liquefiable but not densifiable.?This is generally non-plastic to low-plasticity silty sands and sandy silts.?I tried to illustrate this in the figure above from Armstrong and Malvick (2016) by overlaying zones of compactable (green hatch - small rectangle in bottom left-hand corner), marginal (orange hatch – longer rectangle in bottom left-hand corner) and not compactible (red hatch).?The figure’s proportions make it look like the domain for compactible and marginally compactible liquefiable soils is small, but there are plenty of such soils.” Jim’s statement that there are soils that are liquefiable but not densifiable, is likely correct in a general sense, but it might be more correct to say that such soils can still develop some excess-pore pressures and show “cyclic mobility”, but they are unlikely to exhibit significant consequences ?of this ?behavior unless, for instance, they are under buildings with shallow or no foundations, or are adjacent to large diameter piles of piers, which apply additional cyclic shear stresses and strains to the underlying or adjacent soils.?

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Concluding Remarks

When I was a graduate student in the early nineteen-seventies it was possible to read essentially all the relevant published papers on any topic of interest, but that is impossible for practicing engineers today. Further, the published papers are largely by academics who have to publish or perish. And both research funding and publication processes tend to favor new findings and problems, rather than consolidation of existing knowledge and focusing on practical engineering solutions. For example, the big lesson from the Niigata and Kocaeli earthquakes is not so much the details of the liquefaction processes that were involved, but that it is not a good idea to construct apartment blocks with shallow foundations on loose sands or silts when there is a high water table! But in this note I have tried to pull together some important findings that come out of both research papers and observations of performance in earthquakes over a period of 50 years.

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These can be broadly summarized as follows:

·??????It’s the geology stupid! Or, more specifically, the method of deposition. Don’t forget to ask: “is there any evidence of earthquake-induced liquefaction and settlement or lateral spreading of similar soils in a similar tectonic environment?

?·??????The simplified methods for evaluation liquefaction, settlement and lateral spreading are at best screening tools. No responsible engineer should use any of the simplified methods for evaluating liquefaction or settlement unless they are familiar with each step in the procedure, the limits of applicability of that step and whether the site in question fits within the limits of the overall applicability of the method.

?·??????More detailed site investigations and more advanced analyses are called for on high-value or critical projects.

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References

Andrus, R.D., Hayati, H. and Mohanan, N.P., “Correcting Liquefaction Resistance for Aged Sands Using Measured to Estimated Velocity Ratio”, Journal of Geotechnical and Geoenvironmental Engineering, ASCE, Vol 135, No 6, pp 735-744, 2009

Andrus, R.D., Stokoe, K.H., and Juang, C.H., “Guide for Shear-Wave Based Liquefaction Potential Evaluation”, Earthquake Spectra, Vol. 20, No. 2, May 2004

Arango, I., Lewis, M.R. and Kramer, C., “Updated Liquefaction Potential Analysis Eliminates Foundation Retrofitting at Two Critical Structures”, Soil Dynamics and Earthquake Engineering, Vol 20, pp 17-25, 2000

Armstrong, R.J., and Malvick, E.J., “Practical Considerations in the Use of Liquefaction Susceptibility Criteria”, Earthquake Spectra, Vol. 32, No. 3, August 2016

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Banister, J.R., Pyke, R., Ellet, D.M., and Winters, L., "In-situ Pore Pressure Measurements at Rio Blanco," Journal of the Geotechnical Engineering Division, ASCE, Volume 102, No. GT10, October 1976

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Bartlett, S. F., and Youd, T. L., ‘‘Empirical prediction of liquefaction-induced lateral spread’’, Journal of Geotechnical Engineering, ASCE, Vol. 121, No. 4, 1995

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Boulanger, R.W., et al., “Evaluating Liquefaction and Lateral Spreading in Interbedded Sand, Silt and Clay Deposits Using the Cone Penetrometer”, Geotechnical and Geophysical Site Characterization 5, Australian Geomechanics Society, Sydney, Australia, 2016

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