Further Updated Limitations of Simplified Methods for Evaluating the Potential for Liquefaction etc 4.pdf

Further Updated Limitations of Simplified Methods for Evaluating
Earthquake-Induced Liquefaction and its Consequences
by Robert Pyke PhD, GE
September 12, 2024
This note originated as a two-page document that summarized the main limitations of simplified
methods for evaluating earthquake-induced liquefaction and its consequences. These
consequences include loss of bearing capacity, uplift of structures embedded below the water
table, settlement that occurs when excess pore pressures are dissipated, and lateral spreading.
Most of these limitations also apply to simplified methods for estimating earthquake-induced
settlements of dry or moist sands. Some of the same limitations also apply to simplified methods
for estimating earthquake-induced displacement of slopes, but those methods also involve other
issues which were not covered in my original note and are not addressed here. The original note
also assumed that the reader was familiar with the methods that were being discussed and focused
mostly on their limitations. The previous update of this article described some of the history of
development of these simplified methods in order both to place some context around their current
use and to note the significant approximations that are involved, but individual methods were still
not described in every detail. This further update adds some more detail about individual methods
and adds sub-headings in the hope of making the lengthier document more readable.
The Origin of Simplified Methods
The basis of the original “simplified procedure for evaluating soil liquefaction potential” described
by Seed and Idriss (1971) was that the average cyclic shear stress on a horizontal plane, τav, could
be obtained as a function of the peak ground surface acceleration by the following equation:
τav = 0.65 ∙ amax/g ∙ σv ∙ rd
where amax is the peak ground surface acceleration, g is the acceleration due to gravity, σv is the
total vertical stress at the depth in question, and rd is an empirical factor that accounts for the fact
that the soil column above the depth in question is not rigid. The average, that is uniform, cyclic
shear stress for a number of cycles of loading, taken to be a function of the earthquake magnitude,
can then be compared to the uniform cyclic shear stress that causes “initial liquefaction”, or 100
percent excess pore pressure, in the same number of cycles. This combination of the uniform
cyclic shear stress and number of cycles that triggers initial liquefaction was originally obtained
from cyclic triaxial tests and were subsequently “corrected” to simple shear conditions.
The original simplified procedure was a brilliant way of making it possible for engineers to better
understand and to make preliminary evaluations of the potential for liquefaction at a time when
there were only a few people who could run even equivalent linear one-dimensional site response
analyses. However, times change, and fifty years later, there are many more geotechnical

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engineers with graduate education who are capable of running nonlinear effective stress site
response analyses in order to obtain better estimates of the cyclic shear stresses and strains in
critical layers. In addition, the importance of the depositional history and soil fabric to the
behavior of real soils under cyclic loading is now better understood and there should be more
emphasis on taking this into account.
In any case the original simplified procedure caught on and evolved into a variety of more complex
but still simplified procedures that are available today. One of the first steps was that both sides of
the above equation were divided by the initial vertical effective stress, σv', at the depth in question
and the value τav / σv' became known as the cyclic stress ratio (CSR). Likewise, the cyclic shear
stress causing initial liquefaction, or a prescribed single or double amplitude cyclic shear strain,
was “normalized” by dividing it by the vertical effective stress and became referred to as the cyclic
resistance ratio (CRR). However, because the cyclic shear stress causing initial liquefaction does
not vary linearly with the initial vertical effective stress, this was an imperfect normalization and
an overburden correction factor, Kσ, later had to be added in order to address this issue.
Then, rather than basing the cyclic stress ratio causing liquefaction in a given number of cycles on
laboratory test data, practice shifted to using correlations between the observation of the
occurrence, or non-occurrence, of liquefaction in various earthquakes and Standard Penetration
Test (SPT) blowcounts and Cone Penetration Test (CPT) tip resistances. These correlations were
normally developed for magnitude 7.5 earthquakes, so that then a magnitude scaling factor was
introduced to provide for other magnitudes. Variations of the depth reduction curve that were
magnitude dependent were also developed. Thus, the simplified procedure for “triggering” of
liquefaction, that is reaching 100 percent excess pore pressure, became based on a series of charts
like this:
Figure 1 – Simplified Method for Evaluating Triggering of Liquefaction

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Unfortunately, as one of my colleagues has put it, academics then spent about two decades shifting
these curves around the page rather than focusing on the overall limitations of the procedure.
Extensions of the procedure developed for “triggering” liquefaction were also developed for
evaluating settlement of both saturated and non-saturated sands and for predicting the magnitude
of possible lateral spreading. These extensions, as well as the interpretation of SPT and CPT data,
generally involve a number of correlations which are both very approximate and poorly
understood. Most, if not all, of these correlations are based on laboratory tests, including
calibration chamber tests, conducted on washed and screened clean sands that have not been
subjected to pre-straining or overconsolidation, what Professor Michele Jamiolkowski referred to
as “baby” sands. The behavior of real soils may be different. Thus, the simplified procedures
gradually became more complicated, so that few, if any, practicing engineers fully understand each
step in the various procedures, the limits of applicability of that step and whether the site in
question fits within the limits of the overall applicability of the method. The current procedures
have ended up being both over-simplified and too complicated to understand!
I am not sure how or why it became common practice to rely so much on simplified methods of
analysis, especially for evaluating the consequences of liquefaction such as settlement and lateral
spreading, but likely the paper by Youd et al. (2001), which reported the conclusions of two
academic / industry workshops, and the subsequent paper by Seed et al. (2003) and monograph
by Idriss and Boulanger (2008), all of which focused more on the triggering of initial liquefaction,
had at least something to do with it. I wrote a critical discussion of the Youd et al. paper and my
discussion and the authors’ closure, which largely agreed with my comments, were published as
Pyke (2003). However, I suspect that most graduate students and practicing geotechnical
engineers have read the original Youd et al. paper and not my discussion! Readers of this note
might like to read my discussion and the authors’ closure, but the six points that I made at that
time are still valid and are covered in the following text. However, practicing geotechnical
engineers should not have to read my previous writings or this note to understand that the
simplified methods are not very accurate. It is just wrong if you calculate 30 feet of lateral
spreading of a levee built on a sandy, but relatively dense, foundation, or if you calculate 10-15
inches of settlement for a profile that lacks horizontally continuous sand or silty sand layers, and
that kind of profile has never been known to exhibit any consequences of liquefaction. If that
happens you should seek a better method of evaluation if the project has any importance at all.
The Critical First Step
The first step in any evaluation of liquefaction and its consequences should in fact be 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 for instance 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.
A significant early contribution on the susceptibility of various types of sedimentary deposits to
liquefaction was made by Youd and Perkins (1978) and excellent recent contributions on this
subject are provided by Kayen (2024) and Abayo et al. (2023).

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Do You Always Need to Show a Calculation?
The widespread use of simplified methods of analysis has been driven in part by the belief that
“one has to show a calculation.” However, that belief does not promote better geotechnical
engineering practice, and often leads to the opposite, which is worse practice. The policy of having
to show a calculation promotes the use of computer programs which the user may not understand
and can lead to blind acceptance of the output of computer programs, rather than checking the
results of calculations against common sense and experience. A good screening analysis, for
example, should emphasize common sense and experience rather than calculations. 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 and
a high water table along the Avon River in Christchurch NZ were susceptible to liquefaction prior
to the 2010-11 Christchurch earthquakes.
Common Sense Rules
But, if you have to do a quick simplified analysis, here are some common sense rules that should
be applied:
1. All soil layers that have “clayey” descriptors should be tested as necessary to confirm that
description and should be excluded. (The question of the effect of silty and clayey fines
on liquefaction is discussed in more detail below.)
2. 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).
3. Sand and silty sand layers with measured shear wave velocities exceeding about 720 fps
should also be excluded, see Andrus et al. (2009). However, gravelly soils which contain
sufficient sand to be susceptible to liquefaction may have shear wave velocities up to
about 1050 fps, see Rollins et al. (2022). Broadly speaking, gravels with less than about
30 percent sand are not susceptible to liquefaction because, unless they are confined by
impermeable layers, any excess pore pressures that tend to develop will dissipate too
quickly, but gravels that contain more than 30 percent sand or are otherwise confined
should not be excluded unless the shear wave velocity exceeds 1050 fps. If measured
shear wave velocities are not available, layers that would be otherwise characterized as
“dense” or “very dense” should 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) and Youd et al. (2002).
4. 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.

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General Limitations
More specifically the following eight factors limit the accuracy of simplified methods of analysis
and tend to make the results very conservative:
1. The “cyclic stress ratio” computed working down from the peak ground surface
acceleration assuming no liquefaction 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 liquefaction in any one layer reduces the
amplitudes of the ground motion above that layer.
2. Duration varies with the type of source and the distance as well as with magnitude. Site
specific ground motions have to be much more accurate than using averages of worldwide
data. See item 7 below for additional advantages of using site-specific input motions.
3. 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 resistance to developing excess pore
pressures under cyclic loadings. If simplified analyses suggest that there are significant
consequences of liquefaction, it is desirable to take samples and run hydrometer tests in
order to obtain the full particle size distribution. Measurement of the PI is not
unimportant but viewing the full particle size distribution is even more informative.
Because the interpretation of soil type from CPTs is not foolproof, it is good practice to
always conduct at least several CPT / boring pairs at the outset of an investigation
whenever liquefaction is a possibility. The CPTs should be advanced first, and then SPT
blowcounts can be targeted in the upper portion of any sufficiently thick sandy layers.
Regardless of whether the weighted average shear wave velocity is required in order to
determine the site class according to the building code, measurement of the shear wave
velocities in seismic CPTs is always valuable because the shear wave velocity also serves as
a good index test.
4. The correction of penetration resistance measured using either the SPT or the CPT for
“transition”, “thin layer”, “multiple thin layer”, and “underlying soft layer” 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.
5. Penetration resistance also does not fully reflect the effect of ageing, including
cementation, pre-straining, and over-consolidation on improving the soil response to
cyclic loading.
6. 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

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standard charts that are used in simplified methods may be overly conservative - see for
instance Stokoe (2023). Stokoe’s team conducted relatively well-controlled studies of
silty sands following the earthquakes in Christchurch NZ. By comparison, the soil
profiles and properties in the case histories on which the standard curves are based are
not so well documented.
7. There is also a growing consensus that the occurrence of liquefaction can only be
understood by conducting nonlinear effective-stress site response analyses in which excess
pore pressure development, redistribution and dissipation are 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 are
needed to determine which intervals in the deposit are most susceptible to liquefaction and
to properly understand the case histories on which the simplified, empirical methods are
based.
8. Nonlinear effective-stress site response analyses also use site-specific acceleration histories
as input so that the overall duration and the number of cycles of loading are much better
represented than they are in simplified analyses. And, because settlement and excess pore
pressure development are a function of the prior loading up to the current point in time,
these quantities can be calculated much more accurately using site-specific acceleration
histories as input, rather than assuming a number of cycles of a uniform loading. Just as
equivalent linear site response analyses were a brilliant idea at the time, so was the concept
of a number of uniform cycles being equivalent to an irregular loading, but that was more
than fifty years ago!
While nonlinear effective-stress site response analyses are usually considered to be more
difficult to perform than simplified methods of analysis, and may require more detailed
field and laboratory studies, they are actually easier to understand and they can provide
the engineer with more insight into both liquefaction issues and design ground motions.
9. 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)).
The lens / sinuous channel issue 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 general site conditions 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 possible to provide a meaningful calculation without
running two or three-dimensional response analyses.

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Particular Limitations
It is not possible in a relatively short note to spell out all the particular limitations of individual
simplified methods of analysis, but 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 project. It is
important that practicing engineers understand that not everything that is published in a
journal is correct for every situation, and that simplified analyses of some sites may never be
accurate enough to be useful. Some of the more important limitations are discussed below.
Settlement
Simplified methods for estimating settlements caused by earthquake loadings are particularly
bad. 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).
The conservatism of simplified methods for evaluating settlements of non-saturated sands
results from the use of inappropriate laboratory test results, as well as the limitations associated
with working down from the peak ground surface acceleration. Tokimatsu and Seed (1987) used
data from Silver and Seed (1971), who had used a particularly angular sand which might be
expected to settle much more than natural soils. Tokimatsu and Seed also recommended
doubling the computed settlements of non-saturated sands based on Pyke et al. (1975). But that
is a conservative interpretation of the Pyke et al. paper and doubling the computed settlements
when they are already conservative is unnecessary. More accurate estimation of earthquake-
induced excess pore pressures and settlements requires use of bi-directional nonlinear effective
stress site response analyses in which two horizontal input motions are applied simultaneously
and the increments in excess pore pressures and latent or immediate settlements are then
calculated each half cycle for the motions in each of the two horizontal directions and added to
the total.
The computer program CLiq includes a procedure for estimating settlements of dry, that is non-
saturated sands, due to Robertson and Shao (2010). This is a legitimate, although very
approximate procedure, which is based on Pradel (1998), which in turn is based on Tokimatsu
and Seed (1987), thus the results are likely quite conservative.
There is no hard limit on the total amount of settlement that can occur, but it can be seen from
data presented in Pyke (2020) and Pyke (2022) that additional settlements will be small once
the accumulated settlement reaches a value on the order of 0.5 percent of the layer thickness
under uni-directional loading. Since the settlements caused by each component of motion are
additive, once the accumulated settlement reaches about 0.5 percent of the layer thickness,
additional settlements caused by motion in either direction will also be small. In other words,
accounting for the second component of motion increases the rate at which settlements or latent
settlement accumulate but does not have much effect on the maximum settlement.
For sands below the water table, the problem is more complicated because, as noted for instance
by Macedo and Bray (2018), there are three possible contributors to ground surface or building

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settlements: the settlement that occurs when excess pore pressures are dissipated during or
after the earthquake; the settlement that occurs as a result of material being ejected by excess
pore pressures; and the settlement under or adjacent to buildings and other structures that
results from rocking of the building adding to the cyclic shear stresses and strains in the soil.
For the first of these mechanisms, that is free-field settlement that occurs when excess pore
pressures are dissipated during or after the earthquake, excessive conservatism again results
partly from the use of conservative laboratory data. Tokimatsu and Seed (1987) used old data
from Lee and Albaisa (1974), obtained from cyclic triaxial tests, which are inappropriate for this
purpose. Ishihara and Yoshimine (1992) used a variety of laboratory tests, but as noted
subsequently by Ishihara et al. (2016), they tested only freshly placed washed sands instead of
naturally occurring soils. But simplified analyses of the settlement of saturated sands are even
more conservative because they are both conservative with respect to reaching the point of
initial liquefaction in a particular layer, and conservative regarding the depth of sand that
reaches the point of initial liquefaction. Both these factors lead to larger settlements on
reconsolidation. A nonlinear effective stress site response analysis will generally indicate both
slower development of excess pore pressures and a much more limited depth of liquefaction. See
Crawford et al (2019) and Pyke (2020) for examples of this issue.
The method of Ishihara and Yoshimine (1992) is also conservative for other another reason.
Based on earlier studies by Ishihara’s students, Ishihara and Yoshimine (1992) developed a
relationship between volumetric strain on reconsolidation and the maximum amplitude of the
shear strain as a function of relative density on the basis of various stress-controlled cyclic tests
with uniform cycles of loading. They then settled on a definition of the “factor of safety” against
liquefaction defined as 5 percent double amplitude axial strain in 20 cycles, and, following a
complicated process which I doubt any practitioners have ever read, constructed relationships
between this factor of safety and the maximum amplitude of the shear strain as a function of
relative density. They then concluded that “if the factor of safety is known by means of the
conventional method of liquefaction analysis, it will be possible to circumvent the determination
of the maximum shear strain and to directly estimate the amount of post-liquefaction
volumetric strain.” They cite references from that time on how to evaluate the conventional
factor of safety against liquefaction, but who knows how this is implemented in current
spreadsheets and computer programs and whether it is consistent with how the settlement data
was processed. Moreover, fundamentally the concept of a factor of safety against liquefaction
makes no sense. The value of the expected excess pore pressures as a percentage of the initial
vertical effective stress (the excess pore pressure ratio) makes sense and might be of some value,
but the processes of settlement on reconsolidation after liquefaction or lateral spreading do not
lend themselves to a single value of the factor of safety. Back when laboratory tests were
generally conducted using stress-controlled cyclic tests with uniform cycles and irregular shear
stress histories were converted to an equivalent number of uniform cycles, it was
understandable that Ishihara and Yoshimine (1992) developed this procedure, but Pyke (1973),
Pyke et al. (1975) and Pyke (2022) describe a much better way of computing the accumulation of
latent settlement for irregular cyclic loadings. These procedures compute latent settlements
every half cycle as a function of the peak strain in that half cycle and the accumulated latent
settlement to that point. The accumulated settlement is used as a measure of the extent of
previous cyclic loading as suggested by Martin et al. (1975) and Seed et al. (1978). Pyke (2019)

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and Crawford et al. (2019) show examples of settlements of saturated sands computed in this
manner which are much less than those computed using Ishihara and Yoshimine (1992) and are
also much more consistent with observed settlements.
As an example of the limitations of the Ishihara and Yoshimine (1992) procedure, let me
summarize part of a recent analysis conducted for a site in San Jose, California, focusing on a
deep sand layer which is about 20 feet thick. This sand layer showed tip resistances ranging
from 200 to 600 tsf, which convert to interpreted normalized SPT blow counts in order of 25 to
50, and had measured shear wave velocities in the order of 1200 to 1300 fps. But, even though
“transitions” were excluded and an Ic cutoff of 2.05 was applied, the computed “factors of
safety”, while they are above 2.0 in some parts of this layer, drop to as little as 0.25 in other
parts of the layer, so that volumetric strains in the order of 2 percent and settlement on
reconsolidation of 2 inches were computed. While it would be possible to just ignore this result
because of the depth of the layer, the high measured shear wave velocities, and also the facts
that in this case there is a 20-foot deep clay cap and no history of liquefaction at similar sites,
the computed factors of safety make no sense and the engineer should reject the analysis on that
basis alone.
The method of Zhang et al. (2002), which uses CPT data, basically relies on Ishihara and
Yoshimine (1992) and is conservative for the same reasons, but it does have the advantage of
offering a correction for fines due to Robertson and Wride (1998), which is one of the better
corrections for fines content.
Cetin et al. (2009) recommended a “correction” for the settlement of saturated sands on
reconsolidation which is unconservative. While this work was well-intentioned, the basis for the
authors’ “linear weighting scheme” is not well supported and should not be used. If the
computed settlements are thought to be excessive, there are better ways to get a more accurate
estimate.
There is an even more shocking misunderstanding in the unconservative direction. I have
recently seen some engineers, presumably trying to offset the conservatism of simplified
methods of analysis, argue that sands which are dilative under static loadings, as indicated by
either laboratory tests or CPT-based correlations, do not exhibit any consequences of
earthquake-induced liquefaction. This is simply not correct. It is true that such sands are likely
not susceptible to “static liquefaction” and the generation of flow slides, but it was clearly
established by the original studies by Seed and Lee, and many subsequent studies and field
observations, that under cyclic loading otherwise dilative sands could develop excess pore
pressures and at least reach the point of initial liquefaction, or 100% excess pore pressure
development. However, as noted above, the consequences of this excess pore pressure
development diminish with increasing density, and other factors such as ageing and pre-
straining. Ishihara (1993) suggested that clean sands are dilative above a normalized SPT
blowcount of about 9 and do not exhibit significant consequences of liquefaction above a
normalized SPT blowcount of about 15, at which point sands at normal confining pressures are
strongly dilative. Development of any significant excess pore pressures is unlikely for sands with
normalized SPT blowcounts above 30. But simply assuming that clean sands which are dilative
under static loadings are not susceptible to earthquake-induced liquefaction and its
consequences is unconservative!

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Prediction of the exact amount of settlement that may results from the second and third
mechanisms described by Macedo and Bray (2018), namely, the settlement that occurs as a
result of material being ejected by excess pore pressures and the settlement under or adjacent to
buildings and other structures that results from rocking of the building adding to the cyclic shear
stresses and strains in the soil, is quite difficult. There is no question that these mechanisms are
real, but these problems are best avoided by not using shallow foundations on sites that are highly
susceptible to liquefaction.
Lateral Spreading
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). Youd (2018) also discusses checks that
should be applied to each layer in the soil profile which includes the exclusion of sand layers
under 1 m in thickness and those having a normalized clean sand SPT blowcount of greater than
15. The importance of ruling out sites with significant fines is also noted. See Youd et al. (2009)
for three case histories involving sites where lateral spreading might have been anticipated but
where significant deformations did not in fact develop.
The large scatter in the empirical data, which is not unexpected given the dependence of lateral
spreading on the character of the ground motions and the inherent variability of ground motions
and other factors, suggests that more advanced analyses, such as that described by Pyke (2019b)
which use site-specific input motions, 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 an inappropriate
correlation between cyclic strains in tests conducted without initial shear stresses and
permanent strains developed in the field where there are initial shear stresses, which is what
causes permanent deformations. These are two quite different mechanisms, and the method of
Zhang et al. (2004) should not be used.
Summary comment
It is for all these reasons that Dr. Peter Robertson, the technical advisor to the developer of the
program CLiq, has suggested to me that such tools should be used primarily for screening and
should not be used to obtain the final answer on “high value” projects.
A Good Example
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 quite recent alluvial deposits.
Those studies did not add a great deal of practical value to existing knowledge because most of
these alluvial deposits were clear examples of soils that would be susceptible to liquefaction in

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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 in Figure 2, were found to contain thick deposits of clean
sands that had a typical value of 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.
Figure 2 – From Bray (2022)
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)), rather than the value of 2.6, which is the conventional point of
separation of sandy and clayey soils. This does not necessarily mean that there is no problem at
sites that include silty soils that have Ic values greater than 1.8 or 2.05, but it does mean that you
should not automatically accept the results of simplified analyses and that more detailed site
investigations or 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 even a limit on Ic of 2.05 in simplified analyses may be too
conservative.

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Further Support for Using a Lower Ic Cutoff
Figure 3 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 because, as noted previously, engineering
consequences are unlikely above a normalized clean sand blowcount of 15.
Figure 3 – Sensitivity of SPT Blowcounts to Ic
For Ic values greater than 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 to Consider
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 once the fill is above
water - see Pyke et al. (1978) for an example.
0
10
20
30
40
0 0.5 1 1.5 2 2.5 3 3.5
SPT
N160
(blows/ft)
Ic

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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 unity. 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.
Criteria for Eliminating Soils with Silty and Clayey Fines
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 worst conditions, such as those under buildings constructed on
shallow foundations on loose sands or silts with a high water table, where rocking of the building
can add to the cyclic shear stresses and strains in the soil. Although the criteria of Bray and Sancio
(2006) are widely cited and used, there are important points made in their conclusions which
should be emphasized: “Moreover, some soils with PI < 12 may not be susceptible to liquefaction.
Other factors such as soil minerology, void ratio, overconsolidation ratio, age, etc. are also
contributing factors to liquefaction susceptibility”, and “the proposed criteria should be applied
with engineering judgment.”
Figure 4 was sent to me by Jim Gingery of Keller North America in response to a question that I
asked during a 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
Figure 4 – From Gingery (2021)

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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
foundations, or are adjacent to large diameter piles or piers, which apply additional cyclic shear
stresses and strains to the underlying or adjacent soils. Case histories which show larger
earthquake-induced settlements usually involve lateral spreading or pumping under buildings
with shallow foundations that have been built on loose saturated sands and silts. Analysis of
these case histories is of legitimate academic interest but the most important lesson for
practicing engineers is to spend the money on a deep foundation!
Methods of Analysis
There are basically four kinds of geotechnical analysis:
1. One that does not involve any calculations but relies on common sense and experience.
Experience includes both the engineer’s personal experience and the collected
experience of the profession including observations of behavior in past earthquakes.
2. “Back of the envelope” calculations, that is a simple calculation that gives the engineer a
sense of the expected behavior.
3. Simplified analyses, which might involve more calculations and require a spreadsheet or
a computer program.
4. More complete analyses which include more detailed modeling of the geometry and all
the relevant variables. These can be 1, 2 or 3D models.
In many cases I believe that Option 1 is all that is necessary, although a simple “back of the
envelope” calculation such as the Youd et al. procedure for estimating lateral spreads, or my
suggestion that settlement of non-saturated sands will not exceed about 0.5 percent of the layer
height, may help quantify the potential problem. The late Professor Bob Whitman was of the
opinion that beyond Options 1 and 2, there was no point in running a simplified analysis and
that you should try to model the problem as accurately as possible in a more complete analysis. I
concur with that but note that a more complete analysis will still involve a number of
approximations, so one should not assume that the numerical answer will be precisely correct.
However, the engineer should obtain a much better understanding of the problem and obtain
answers that are “in the ballpark” if adequate field and laboratory data are available.

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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 written 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, especially 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.
These can be broadly summarized as follows:
• It’s the geology stupid!
Or, more specifically, the method of deposition and the resulting soil fabric. And 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?
• It is not wise to completely rely on automated processes involving spreadsheets or
computer programs in geotechnical engineering. The variability in the depositional
history and current geometry of soil deposits necessarily calls for some judgment to be
exercised by the responsible engineer who should always ask, “does this answer make
any sense?”
• The simplified methods for evaluating liquefaction, settlement and lateral spreading are at
best screening tools. Ideally, no responsible engineer should rely on any of the simplified
methods for evaluating liquefaction and its consequences unless they have some
familiarity 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.
And reviewers should not automatically either accept or reject the results of simplified
analyses on projects of any importance.
• More detailed site investigations and more advanced analyses may be called for on high-
value or critical projects. Even for simplified analyses it is good practice to initially conduct
seismic CPTs and, if there are any potentially liquefiable layers, to follow-up with
companion borings in which SPTs are targeted towards the top of those layers and full
grain size curves are obtained.
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