This is a further updated version of my presentation on the subject topic. It includes additional material on seismic settlements. The calculation of seismic settlements using simplified methods has been out of control for some time and is leading to expensive ground improvement when it is not necessary. An improved method of evaluation is described and illustrated by case histories.

1. Improved Analyses of Earthquake-Induced
Liquefaction and Settlement
By Robert Pyke Ph.D., G.E.
Robert Pyke, Consulting Engineer, Walnut Creek, CA, USA
March 2020

2. This presentation is based on:
• Pyke, R., “Improved analyses of earthquake-induced liquefaction and settlement”,
Proc. 7th International Conference on Earthquake Geotechnical Engineering,
Rome, June 2019
• Pyke, R. and North, J., “Shortcomings of simplified analyses of earthquake-induced
liquefaction and settlement”, Proc. 7th International Conference on Earthquake
Geotechnical Engineering, Rome, June 2019
• Crawford, C., Tootle, J., Pyke, R. and Reimer, M., “Comparison of simplified and
more refined analyses of seismic settlements”, Proc. 7th International Conference
on Earthquake Geotechnical Engineering, Rome, June 2019

3. This version has been annotated to provide the continuity that would
be given orally in a live presentation.
And, some additional material been added regarding various
complications such as partial saturation or partial drainage, silty and
clayey fines, thin layers, the effect of a non-liquefiable crust, and fabric
or “ageing”. These factors create greater uncertainty and
conservatism in the application of simplified methods of evaluation
and argue for wider use of site-specific evaluations with careful site
investigations and nonlinear effective stress analyses.
And, additional material on the limitations of simplified methods for
computing seismic settlements, both for materials above and below
the water table, has been added.

4. Simplified Analyses of Liquefaction and Settlement
• Began with Seed and Idriss (1969) when running cyclic laboratory tests and site response analyses were
very specialized activities. 50 years ago! When 3 people in the world could run site response analyses and
input motions were passed around as punched card decks!
• (Subsequently) obtain cyclic stress ratio causing liquefaction from penetration resistance rather than from
laboratory tests
• (Later still) for saturated sands procedures were developed for obtaining settlements as a function of the
factor of safety against liquefaction – e.g. ishihara and Yoshimine (1992), Zhang et al. (2002). For sands
above the water table simplified procedures for estimating seismic settlement were developed by
Tokimatsu and Seed (1987) and Pradel (1998). Robertson and Shiao (2010) adapted Pradel’s method for
use with CPT data
• The next slide shows the key elements of the simplified procedure graphically. These figures are taken from
Boulanger and Idriss (2014). Basically the procedure uses a simple formula to computed the induced cyclic
shear stresses from the ground surface peak acceleration and uses penetration resistance and case
histories to evaluate the cyclic shear stresses causing liquefaction.

6. Issues with Simplified Analyses
• Duration varies with type of source and distance as well as with magnitude. Site specific motions are
much more accurate than averages of worldwide data.
• Ground surface acceleration and induced shear stresses are strongly impacted by the development,
redistribution and dissipation of excess pore pressures and the induced shear stresses are generally
overpredicted by simplified methods – see the examples in this presentation
• Penetration resistance is a poor indicator of soil behavior under cyclic loading and is overly sensitive
to the presence of silty and clayey fines
• The failure to properly account for ageing effects – which include over-consolidation, pre-straining,
improved packing and chemical bonding – since the empirical data that is used in simplified methods
largely comes from recent deposits
• See for instance Boulanger et al. (2016), Ntritsos et al. (2018), Pyke and North (2019), Beyzaei et al
(2019), Olson et al. (2020) for examples and more details

7. Particular Issues with Simplified
Analyses of Seismic Settlements
• The most obvious issue is the simple fact that seismic
settlements are commonly wildly over-predicted relative
to what has actually been observed in similar ground
conditions in previous earthquakes
• For saturated sands perhaps the biggest problem is that
simplified methods for analysis of liquefaction over
predict the number of layers in the profile that will
actually reach initial liquefaction and trigger larger
settlements – see the examples later n this presentation
• But also, settlements following liquefaction obtained
from Japanese data (Ishihara and Yoshimine, 1992) are
fine for layers that reach initial liquefaction but may be
exaggerated for layers that do not reach initial
liquefaction
• A good example of the over-conservatism of simplified
methods in predicting seismic settlements is given by
Geyin and Maurer (2019) using reasonably good data
from Christchurch NZ. This data is for saturated sands
where there was widespread liquefaction including
ejection of sands and silts which should add to the
observed settlements, however, use of the method of
Zhang et al. (2002), which relies on CPT measurements,
grossly over-predicts the observed settlements

8. Seismic settlements of non-saturated sands
• For sands above the water table there are multiple reasons for this
excessive conservatism including the fact that the laboratory data that
Tokimatsu and Seed (1987) relied on was inappropriate and overly
conservative. Tokimatsu and Seed used data from Silver and Seed
(1971) and Lee and Albaisa (1974), but Lee and Albaisa’s data was
obtained from triaxial tests, which are inappropriate, and Silver and
Seed, in addition to having measurement errors, used Crystal Silica No.
20 sand, a particularly angular sand, which, notwithstanding the
findings of Duku et al (2008), shows greater settlements than the
sands encountered in practice. Pradel (1998) then relied on the data
from Silver and Seed and the Tokimatsu and Seed paper
• In the case of the method of Robertson and Shao (2010) for estimating
seismic settlements from CPT data, which is based on Pradel (1998)
and included in the computer program C-LIQ, the correction of
interpreted N160cs values for fines using the parameter Ic is grossly
inadequate. Simply plotting the interpreted N160cs values or relative
densities against Ic, as shown on the right, will demonstrate this. This
leads to two compounding errors: (1) the low-strain shear modulus is
underestimated so that the computed cyclic shear strains are too
great; and (2) the settlement as a function of a given cyclic shear strain
is also too great
0
10
20
30
40
50
60
70
80
90
100
0 0.5 1 1.5 2 2.5 3
RelativeDensity(%)
Ic

9. Other issues re seismic settlements
• The estimated settlements are then commonly doubled citing Seed, Pyke and
Martin (1978), but the concept of adding the settlements caused by two
orthogonal components of shaking applies only to more accurate calculations. It is
not necessary when the calculation is very approximate and conservative in the
first place!
• The estimated settlements are, however, sometimes reduced using a “depth
weighting factor” suggested by Cetin et al. (2009). This is a completely arbitrary
fudge factor and should never be used

10. Additional factors that complicate evaluation of the potential
for excess pore pressure development and liquefaction
• Partial saturation and partial drainage
• The presence of silty and especially clayey fines
• Thin layers and lenses
• A non-liquefiable crust
• Soil fabric, resulting from the time under sustained pressure, or age, the previous shear strain history,
the coefficient of lateral earth pressure and overconsolidation.
All of these factors create greater uncertainty and conservatism in the application of simplified methods
of evaluation and argue for wider use of site-specific evaluations with careful site investigations and
nonlinear effective stress analyses.
More generally, the engineer should be aware of both local geology and history of the occurrence of
liquefaction and settlement in the same tectonic environment. See Pyke (1995, 2003, 2016) and Semple
(2015).

11. Partial saturation and partial drainage
• Full development of excess pore pressures under cyclic loading requires fully saturated, undrained
conditions
• In the laboratory considerable effort including the use of de-aired water and back pressure is required
to fully saturated test specimens. Therefore in the field full saturation is unlikely above the depth of a
permanent water table. Seasonal wetting greatly reduces the chance of liquefaction. See for instance
Banister, Pyke et al. (1976)
• Redistribution and dissipation of excess pore pressures will obviously impact the development of
excess pore pressures and whether of not initial liquefaction (triggering) is reached. This effect must
be included in some averaged, approximate way in empirical procedures but both greater accuracy
and insight can be obtained by conducting nonlinear effective stress site response analyses

12. Effects of silty and clayey fines
• Both silty and clayey fines reduce the penetration resistance of soils because penetration resistance is
a function of both strength and compressibility
• Most silt particles are essentially very fine sands so they may not have much impact on excess pore
pressure development under cyclic loading except for how they impact packing of particles or “fabric”
• Various corrections for silty fines have been proposed for both SPT blowcounts and CPT tip resistances
but their general applicability and accuracy is questionable. For important projects regional
correlations or site-specific laboratory tests are desirable. Dobry and Abdoun (2017) note that shear
wave velocity is less impacted by the presence of silty fines than penetration resistance and may be
preferred for use in liquefaction “triggering” relationships.
• Clay particles have electric charges and are inherently cohesive. Basically any soil that has a “clayey”
descriptor is non-liquefiable. In routine practice perhaps more errors in evaluating liquefaction
potential result from this issue than any other. SPT samples with low blowcounts should always be
carefully examined and particle size distributions and liquid and plastic limits obtained for each
material type. CPT interpreted soil types cannot be relied on in silty and clayey soils and should always
be checked by taking at least some samples.

13. Particular Issues Regarding Thin Layers
• If only seen in a single boring or CPT they may be lenses rather than layers, and the induced stresses in a
lenses may be lower than in the surrounding material – see Pyke (1995)
• The CPT may report thin sand layers above the actual location of the layer. Sites with thin sand layers
should have a follow-up investigation with SPT samples taken continuously through and below the
indicated thin layers with particle size distributions and liquid and plastic limits being obtained for each
material type. See also Beyzaei et al (2019).
• Both the SPT and the CPT “have a nose”. They can detect softer layers even before they get to them. See
the following slide taken from Boulanger and DeJong (2018). A sand layer underlain by a soft clay has to
be in the order of one meter thick in order that the true peak tip resistance be recorded. Similarly, an SPT
sampler needs to have several feet of the same material below the tip to record a meaningful blowcount.

15. Effects of a non-liquefiable crust
A non-liquefiable and less pervious crust overlying a potential liquefiable deposit can do two things:
1. Increase the peak excess pore pressures that are developed in the liquefiable layer because drainage is
impeded;
2. Limit the occurrence and observation of the effects of liquefaction at the ground surface
Some relevant references are Ishihara (1985), Youd and Garris (1995) and Green et al. (2018)
However, the use of a Liquefaction Potential Index as suggested in the last of these papers is no substitute for
a site-specific nonlinear effective stress analysis which can directly evaluate the effect of a less pervious crust
and other “system effects” resulting from less pervious thin layers. See “supporting publications” at the end
of this presentation and Ntritsos et al. (2018)

16. How to account for soil fabric and ageing
• Ideally these effects might be accounted for by taking high quality undisturbed samples and conducting
laboratory tests to evaluate the effect on the cyclic stress ratio causing liquefaction, but this is difficult and
time consuming.
• Therefore empirical methods or generic data from research studies may be the best way to account for
these effects – see the following three slides.
• The first of the following three slides shows the effect of just overconsolidation on freshly deposited test
specimens. There is some experimental evidence that the significant gains in resistance caused by
overconsolidation are diminished if liquefaction subsequently occurs but the design intent should be to
avoid that happening
• “Ageing” averages the long-term effects of sustained pressure, pre-straining, overconsolidation and
cementation or better seating at grain contacts. The second slide, from Lewis et al. (2005) shows various
estimates of the increase in resistance to triggering of liquefaction with age. The scatter is understandable
since different factors might be involved in each case. Ideally engineers will develop and use appropriate
regional factors as suggested by Dobry and Abdoun (2017).
• An alternate approach, shown on the third slide, has been suggested by Bwambale and Andrus (2019) in
which MEVR, the ratio of the measured shear wave velocity to the estimated shear wave velocity using
penetration resistance and correlations for recently deposited sands, is used as an alternative to time.

17. Effect of overconsolidation:
Effect is greater than
suggested just by
the increase in the
coefficient of lateral
earth pressure!
From Pyke (2015)

18. Effect of Ageing
From Lewis et al. (2005)

19. Effect of Ageing
From Bwambale and Andrus (2019)

20. Beware of automation!
• It is not only poor practice, but can be dangerous, to automate the process of
processing field data and conducting analyses without human oversight and
intervention!
• See my companion presentation on “Improved Analysis of Potential Lateral Spreading
Displacements in Earthquakes”, for an excellent example of the need for human
intervention and interpretation of a soil profile that is taken from Youd (2018), or
better yet, read the Youd paper!

21. First Case History:
This case history involves an elementary school site in Alameda, CA, close to the
Hayward fault, which can be seen in the next slide. The slide after that shows a close
up of the site which is situated on a made made island with about 15 feet of
hydraulically placed sand fill. The sand fill is obviously susceptible to liquefaction, but
all sand fills are not the same and this one showed no signs of liquefaction and
settlement in the 1989 Loma Prieta earthquake.

22. Hayward Fault

24. As part of a required seismic safety evaluation a local geotechnical consultant
evaluated the potential for liquefaction and seismic settlement using CPT data and
the program C-LIQ*. Neither this application nor the subsequent slide, which is
taken from a poor-quality .pdf that is part of the public record, does justice to C-LIQ,
which is a nice program for looking at data but contains some methods of analysis
which are questionable at best . The methods of Zhang et al. (2002), which is based
in part on Ishihara and Yoshimine (1992), and is included in C-LIQ for estimating the
settlement of saturated sands, and of Robertson and Shao (2010), which is based in
part on Pradel (1998), and is included in C-LIQ for estimating the settlement of sands
above the water table, rely on correlations which are approximate and apply at best
only to freshly deposited clean sands, do not apply at all to sands with fines, and
overall tend to be very conservative for estimating earthquake-induced settlements.
* https://geologismiki.gr/

25. As already noted, the next slide is hard to read but it shows a typical CPT sounding
that was pushed at the direction of the geotechnical consultant. No borings were
made, or samples taken. However, it is known from accumulated local experience
that the profile basically has five strata:
1. Sand fill
2. young Bay Mud, lightly OC clayey silts and silty clays
3. Merritt Sand, late Pleistocene wind-blown sands
4. Old Bay Clay, OC clayey silts and silty clays
5. Lower Alameda Formation, very dense sands and gravels

27. However, the previous slide shows liquefaction in four different strata
down to a depth of more than 60 feet and an estimated 10 inches of
settlement of the ground surface. In reality, only one of these materials,
the hydraulically placed sand fill, is susceptible to liquefaction and
seismic settlement. This is a big problem!

28. Because the structural engineer advised that is was uneconomical to
retrofit the school buildings to accommodate 10 inches of settlement,
the school was closed, and the children are being bussed to other
schools, instead of walking to their neighborhood school.
Thus, over-conservatism led not to increased safety but to adverse
social consequences.

29. The subsurface profile is idealized in the following slide. The ranges of
normalized SPT blowcounts are interpreted from the CPT data and the
shear wave velocities are estimated by the author based on his
extensive local experience. On the basis of this data and local
knowledge, the sand fill might be susceptible to liquefaction and
settlement, but the other layers are not.

31. The profile on the previous slide is simplified in that it averages soil properties over
each depositional unit, but it captures the essential depositional units and site
response in earthquakes will be dominated by the average properties. It will now be
used in a site response analysis using the new computer program TESS2. TESS2 uses
the same explicit finite difference scheme as the earlier program TESS, but it has
been rewritten to incorporate many new features including the ability to
simultaneously analyze the response of a soil column to two horizontal input
motions. Some further details regarding TESS2 follow presentation of the results.

32. The following slide provides an indication of the amplitudes of the input motions,
which were fitted to the target spectrum, labelled Site Class C. Five pairs of
horizontal motions representing a magnitude 7+ earthquake on the Hayward fault
were input at the top of the Lower Alameda Formation at a depth of 100 feet.
It can be seen that the ground surface motions are significantly reduced as a result
of both damping in the mud layers and the development of excess pore pressures
and softening of the sand fill.

33. Input motions
Ground surface motions

34. And the next slide shows the printed output from TESS2 with three added columns.
The basic results are just for one horizontal component of one motion, but the excess
pore pressures and the settlements are the sum of the contributions from each of the
two input motion components, and are necessarily the same for both components.
The column headed Rumax is the maximum excess pore pressure ratio, and the
column headed Rufinal is the excess pore pressure ratio at the end of the specified
input motion. Note that liquefaction only occurs in Layer 5. Layer 4 also has very high
excess pore pressures, but it does not quite liquefy because of simultaneous
dissipation of excess pore pressures towards the water surface.

35. TESS2 Results

36. Column 3 of the previous slide shows the peak shear stresses generated by this
component of motion, and column 9 shows the peak shear stresses that would have
been computed using Boulanger and Idriss (2014). The simplified method values are
approximately twice the actual values. This is one of the reasons that the simplified
analysis of liquefaction is conservative. Because the liquefaction analysis is
conservative, the evaluation of potential seismic settlements using the method
embodies in C-LIQ is very conservative. This run, plus the 4 other runs, suggest
seismic settlements of the ground surface in the order of an inch or two. Not ten
inches! This, more correct, finding would have made an enormous difference to the
feasibility of retrofitting the school buildings and avoided the need to bus the
children to other schools.

37. TESS2 - bi-directional, nonlinear, effective stress site
response analyses
• The same explicit finite difference solutions for response and
redistribution and dissipation of excess pore pressures as TESS
• Simple hyperbolic soil model – see Pyke (1979, 1993, 2004,2020)
• Excess pore pressures following Seed, Martin and Lysmer (1976)
• Settlements following Pyke (1973) and Seed, Pyke and Martin (1978)
• Runs two horizontal components simultaneously and adds excess
pore pressures and settlements or latent settlements
From Pyke (2015)From Pyke (2015)

38. Ease of use features of TESS2
• Selection and modification of site-specific input motions now much easier thanks to the PEER
strong motion database etc.
• Users basically only have to specify shear wave velocity, soil type and undrained shear strength or
apparent relative density for each soil layer
• Most or all of the other required soil properties have defaults that are built in, although,
importantly, the user can specify site specific data if this is available
• In particular, default relationships for the increase in resistance to liquefaction with various aspects
of “ageing” are built into the program as a function of either time or MEVR
• However, the user still has to think about the depositional environment, the age of the deposit,
and how to subdivide the profile into layers

39. Calculation of seismic settlements
• Settlements in non saturated sand layers and latent settlements in fully saturated sand layers in
TESS2 are computed using the methodology of Martin, Finn and Seed (1975), Seed, Pyke and
Martin (1978) and data from Pyke (1973), which is supplied as a default. Users can substitute their
own site-specific data if desired. The results of a typical cyclic simple shear test are shown on the
following slide. There are two points worth noting:
1. The settlement per cycle clearly decreases with an increasing number of cycles;
2. The shear modulus increases with the number of cycles
[While this test was intended to be at a constant cyclic shear strain, because the sample becomes
stiffer under cyclic loading and there is compliance in the testing device, the cyclic strains are not
entirely constant and the results have to be reduced on a cycle by cycle basis. Also, similar effects
of the number of cycles were also seen in bi-directional shaking table tests so the decrease in
settlement per cycle and the increase in shear modulus is not the result of unidirectional shaking.]

41. A summary of the data obtained for freshly pluviated Monterey No. 0 sand is shown
on the following slide. It may be seen that even for this “baby sand”*, the
settlement is unlikely to exceed about 0.5 % of the layer height, even for loose sands
and strong shaking. Thus, one may conclude that for natural soils, the settlements,
even for loose sands and strong shaking, are unlikely to exceed about 0.5% of the
layer height unless liquefaction occurs.
* A washed and graded sand freshly deposited in the laboratory without over-consolidation, pre-
straining or ageing.

43. The key to the Martin and Pyke procedure for computing settlements
caused by irregular cyclic loadings is to know the settlement per cycle as
a function of cyclic shear strain and the accumulated settlement or
latent settlement to that point, as shown for one relative density on the
following slide.

44. Basis for calculation of settlements

45. In TESS2, if any layer reaches 100 percent excess pore pressure, the latent
settlements for fully saturated layers by default jump to the post-liquefaction
settlements suggested by Ishihara and Yoshimine (1992). However, prior to that
point, the settlements estimated by the Seed, Pyke and Martin (1978) procedure
and using the default data from Pyke (1973) are generally lower than the Ishihara
and Yoshimine values. The user can of course over-ride the default data with site-
specific values if these have been obtained.

46. Ishihara and Yoshimine (1992):

47. Second Case History
This case history involves a typical site in “Silicon Valley”, on the margins of San
Francisco Bay. The eight CPT soundings shown on the next slide all showed some
sand, but always at different depths, consistent with late Pleistocene alluvial fan
deposits. Thus the deposit is lensed, not layered. There is no record of such
materials ever liquefying, and if any excess pore pressures develop in one of the
sand lenses, that will cause softening and the lens becomes a soft inclusion and
carries less load. See Pyke (1995).

49. Detailed results for CPT-2 are shown on the following two slides using the methods
for evaluating liquefaction based on Boulanger and Idriss (2014) and for estimating
seismic settlements based on Zhang et al. (2002), as implemented in C-LIQ .
The results shown apply the standard option that is built into C-LIQ for eliminating
transitions and assume an “Ic cutoff” of 2.6, which is standard boundary between
“clayey” and “non-clayey” soils. Occasional liquefaction is shown to a depth of 28 m
and the accumulated seismic settlements at the surface are 10 cm.

52. Parametric study:
The next slide shows the results of a limited parametric study using
C-LIQ and CPT soundings 2 and 5. Result of the estimated seismic settlements are
shown for four cases.
> Without detection and elimination of transitions and an Ic cutoff of 2.6
> With elimination of transitions and an Ic cutoff of 2.6
> With elimination of transitions and an Ic cutoff of 2.05, the limiting value for clean
sands, which is more applicable to the correlations used internally in C-LIQ
> With an “ageing factor” of 1.5, which produces some odd results in this case

53. Settlements in cm computed using C-LIQ:
Assuming: CPT-2 CPT-5
No transitions
17 25
Default transitions
10 23
Ic cutoff = 2.05
4 1
Ageing factor = 1.5
4 2

54. The most correct answer for the estimated seismic settlement at this
site, even under strong ground motions, is likely zero. The above results
suggest that the simplified methods of analysis implemented in C-LIQ
can give reasonable results if the user makes good choices regarding the
input parameters, but that if used blindly, these methods can give very
conservative results.

55. The following slide shows a summary of the results obtained using bi-directional,
nonlinear, effective stress analyses using TESS2. Conservatively assuming that the
sand layers are horizontally continuous, there is some excess pore pressure
development in the cleaner sand layers, the peak accelerations at the ground
surface are reduced from 0.55 g to approximately 0.3 or 0.4 g depending on the
values assumed for the hydraulic conductivities, and the estimated settlements are
reduced to 0.6 mm, close to the likely correct answer of zero.
Again, this is still a conservative approximation, but in this case and in many others,
it is not worth doing the necessary site investigations and a 2D or 3D analysis that
models the lensed deposit in detail.

56. Settlements in cm for CPT-2 calculated using TESS2

57. Third Case History – River Islands
• On Stewart Tract within City of Lathrop, CA
• Mostly underlain by Pleistocene Sands
• But one patch, outlined on the following slide, was reworked about
3000 years ago and shows lower penetration resistances and implied
densities

59. Preliminary analyses, as shown on the next slide, using the methods of
Boulanger and Idriss (2014) and Zhang et al. (2002), as implemented in
C-LIQ, indicate liquefaction for the full depth of the reworked sands and
suggest settlements in the order of 1 foot or more. This would mean
that expensive ground improvement would be required for the planned
high-tech office and R&D use, making that use non-competitive with
alternate locations.

61. Thus, the developer elected to fund improved site investigations,
laboratory testing and analyses in order to ascertain whether site
improvement was really required. This involved pushing new CPTs with
shear wave velocity measurements, measurement of SPT blowcounts,
and use of a piston sampler in adjacent borings. Some of the key data is
shown on the following slide.

63. Samples were drained in the field and then carefully transported to the
laboratory. Stress-controlled, constant height, cyclic simple shear tests
were then conducted in order to determine the cyclic stress ratios
causing liquefaction using a new cyclic simple shear device designed by
Dr Michael Riemer of the University of California, Berkeley. With very
careful handling and practice it was possible to get relatively
undisturbed samples into the test device.

69. The results of the cyclic simple shear tests are summarized on
the following slide. These results justified the use of a cyclic
resistance ratio higher than that which would have been
obtained from the CPT or SPT data and back-calculation of case
histories.

71. Potential seismic sources and design response spectra are shown on the
following slide. It was decided to conservatively adopt the use of median
+ one standard deviation ground motion generated by a magnitude 7
earthquake on Segment 7 of the Great Valley fault for this evaluation.

73. Selected results from analyses using TESS2 are shown on the following
slides. Results are shown for just one of the five pairs of input motions
that were used and for two different values of the hydraulic
conductivity. Run No. 60202 used a hydraulic conductivity in the
reworked sands of 10-2 cm/sec and Run No. 60203 used 10-3 cm/sec.
Although it had been intended that site-specific data on seismic
settlement be obtained by running strain-controlled cyclic simple shear
tests, the schedule did not allow that, and so settlements were
estimated using the default data built into the program based on
extensive tests run on Monterey No. 0 sand by Pyke (1973). This is
believed to be conservative.

75. Note that in the previous slide, liquefaction occurs only in one layer of
the profile and not for the full 40 feet depth of the saturated reworked
sands. For the best estimate of CRR10 of 0.24 used in these runs, the
difference in the values of hydraulic conductivity made little difference
And note in the following slide that the peak induced shear stresses are
only about two-thirds of those obtained using
C-LIQ and Boulanger and Idriss (2014).

78. As in the previous case histories, the estimated seismic settlements are
much less than those obtained using C-LIQ for multiple reasons, but
most notably because the induced cyclic stresses are less than those
estimated using the simplified method and because less layers are found
to liquefy.

80. The results of limited parametric study are summarized on the following
slide. For the more conservative value of CRR10 and for a hydraulic
conductivity of 10-3 cm/sec, two layers in the model liquefy and the
estimated settlement of the ground surface is about 3 inches. The more
likely estimate of the maximum ground surface settlement is 2 inches.
Because redistribution and dissipation of excess pore pressures can be
significant, measurement or assumption of appropriate values for the
hydraulic conductivity is important. However, the default values that are
built into the program as a function of soil type will be adequate for
most routine analyses.

81. Relative Density 50 % 50 % 60 % 60 %
CRR10 0.20 0.24 0.20 0.24
K = 10-2cm/sec 1.8 1.8 1.7 1.7
K = 10-3cm/sec 3.4 1.8 2.8 1.7
Limited parametric study - settlement in inches

82. Conclusions
• Simplified analyses at best should only be used for screening analyses
• Depositional history and historic performance may actually be a better screening
tool
• If you have to show a calculation, or if there are real economic or safety issues, do
a decent calculation
• It’s not that hard to do a decent calculation these days (if you have the right tools)

83. Supporting publications
• In addition to the references listed below, there were two excellent invited papers presented at the
recent 7th International Conference on Earthquake Geotechnical Engineering held in Rome in June 2019,
that illustrate how nonlinear effective stress analyses are required to understand the development and
dissipation of excess pore pressures in real soil profiles. These are:
Keynote Lecture 08, “The use of numerical analysis in the interpretation of liquefaction case
histories”, by Steve Kramer, and
Keynote Lecture 09, “Key aspects in the engineering assessment of soil liquefaction”, by Misko
Cubrinovski
• See also:
Hutabarat, D., and Bray, J.D., “Effective stress analysis of liquefiable site in Christchurch to discern
the characteristics of sediment ejecta”, which won the best student paper award

84. Companion presentations
• “Improved Analysis of Potential Lateral Spreading Displacements in Earthquakes”, an
expanded version of an invited presentation given at the 2nd Ishihara Colloquium sponsored by the San
Diego Chapter of EERI and held at San Diego State University, August 2019
https://www.linkedin.com/pulse/improved-analysis-potential-lateral-spread-earthquakes-robert-pyke
• “Limitations of Vs30 for Characterizing Sites for Ground Motion Studies and Guidance on
the Conduct of Nonlinear Site Response Analyses”, which focusses more on the use of nonlinear
effective stress analyses to obtain improved estimates of site effects in earthquakes and discusses the
implications of the new seismic loading requirements in ASCE 7-16. Posting on LinkedIn is imminent.

85. References 1
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and Soil Dynamics V ASCE Geotechnical Special Publication 290, 2018
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Division, ASCE, Vo. 146, No. 2, 2020

86. References 2
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87. The End!
If you have any comments or questions, write to me at: bobpyke@attglobal.net