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Evaluating cyclic liquefaction potential using thecone
penetration test
P.K. Robertson and C.E. (Fear) Wride
Abstract: Soil liquefaction is a major concern for structures
constructed with or on sandy soils. This paper describesthe
phenomena of soil liquefaction, reviews suitable definitions, and
provides an update on methods to evaluate cyclicliquefaction using
the cone penetration test (CPT). A method is described to estimate
grain characteristics directly fromthe CPT and to incorporate this
into one of the methods for evaluating resistance to cyclic
loading. A worked exampleis also provided, illustrating how the
continuous nature of the CPT can provide a good evaluation of
cyclic liquefactionpotential, on an overall profile basis. This
paper forms part of the final submission by the authors to the
proceedingsof the 1996 National Center for Earthquake Engineering
Research workshop on evaluation of liquefaction resistance
ofsoils.
Key words: cyclic liquefaction, sandy soils, cone penetration
test.
Rsum : La liqufaction des sols est un risque majeur pour les
structures faites en sable ou fondes dessus. Cetarticle dcrit les
phnomnes de liqufaction des sols, fait le tour des dfinitions sy
rapportant et fournit une mise jour des mthodes permettant dvaluer
la liqufaction cyclique partir de lessai de pntration au cne (CPT).
Ondcrit une mthode qui permet destimer les caractristiques
granulaires partir du CPT et dincorporer directement lesrsultats
dans une des mthodes dvaluation de la rsistance au chargement
cyclique. Un exemple avec solution estaussi prsent pour illustrer
comment la nature continue du CPT peut donner une bonne ide du
potentiel deliqufaction cyclique sur la base dun profil densemble.
Cet article fait partie de la contribution finale des auteurs
auxcomptes-rendus du sminaire NCEER tenu en 1996 sur lvaluation de
la rsistance des sols la liqufaction.
Mots cls : liqufaction cyclique, sols sableux, CPT
[Traduit par la Rdaction] Robertson and Wride 459
Soil liquefaction is a major concern for structures con-structed
with or on saturated sandy soils. The phenomenonof soil
liquefaction has been recognized for many years.Terzaghi and Peck
(1967) referred to spontaneous liquefac-tion to describe the sudden
loss of strength of very loosesands that caused flow slides due to
a slight disturbance.Mogami and Kubo (1953) also used the term
liquefaction todescribe a similar phenomenon observed during
earthquakes.The Niigata earthquake in 1964 is certainly the event
thatfocused world attention on the phenomenon of soil
liquefac-tion. Since 1964, much work has been carried out to
explainand understand soil liquefaction. The progress of work
onsoil liquefaction has been described in detail in a series
ofstate-of-the-art papers, such as those by Yoshimi et al.(1977),
Seed (1979), Finn (1981), Ishihara (1993), and Rob-ertson and Fear
(1995). The major earthquakes of Niigata in1964 and Kobe in 1995
have illustrated the significance andextent of damage that can be
caused by soil liquefaction.Liquefaction was the cause of much of
the damage to theport facilities in Kobe in 1995. Soil liquefaction
is also a
major design problem for large sand structures such as
minetailings impoundments and earth dams.
The state-of-the-art paper by Robertson and Fear (1995)provided
a detailed description and review of soil liquefac-tion and its
evaluation. In January 1996, the National Centerfor Earthquake
Engineering Research (NCEER) in theUnited States arranged a
workshop in Salt Lake City, Utah,to discuss recent advances in the
evaluation of cyclic lique-faction. This paper forms part of the
authors final presenta-tion (Robertson and Wride 1998) to the
proceedings of thatworkshop (Youd and Idriss 1998). The objective
of this pa-per is to provide an update on the evaluation of cyclic
lique-faction using the cone penetration test (CPT).
Severalphenomena are described as soil liquefaction. In an effort
toclarify the different phenomena, the mechanisms will bebriefly
described and definitions for soil liquefaction will
bereviewed.
Before describing methods to evaluate liquefaction poten-tial,
it is important to first define the terms used to explainthe
phenomena of soil liquefaction.
Figure 1 shows a summary of the behaviour of a granularsoil
loaded in undrained monotonic triaxial compression. Invoid ratio
(e) and mean normal effective stress (p) space, asoil with an
initial void ratio higher than the ultimate stateline (USL) will
strain soften (SS) at large strains, eventually
Can. Geotech. J. 35: 442459 (1998) 1998 NRC Canada
442
Received April 7, 1997. Accepted March 5, 1998.
P.K. Robertson and C.E. (Fear) Wride. GeotechnicalGroup,
University of Alberta, Edmonton, AB T6G 2G7,Canada. e-mail:
[email protected]
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reaching an ultimate condition often referred to as critical
orsteady state. The term ultimate state (US) is used here,
afterPoorooshasb and Consoli (1991). However, a soil with aninitial
void ratio lower than the USL will strain harden (SH)at large
strains towards its ultimate state. It is possible tohave a soil
with an initial void ratio higher than but close tothe USL. For
this soil state, the response can show limitedstrain softening
(LSS) to a quasi steady state (QSS) (Ishihara1993), but eventually,
at large strains, the response strainhardens to the ultimate state.
For some sands, very largestrains are required to reach the
ultimate state, and in some
cases conventional triaxial equipment may not reach theselarge
strains (axial strain, a > 20%).
During cyclic undrained loading (e.g., earthquake load-ing),
almost all saturated cohesionless soils develop positivepore
pressures due to the contractive response of the soil atsmall
strains. If there is shear stress reversal, the effectivestress
state can progress to the point of essentially zero ef-fective
stress, as illustrated in Fig. 2. For shear stress rever-sal to
occur, ground conditions must be generally level orgently sloping;
however, shear stress reversal can occur insteeply sloping ground
if the slope is of limited height
1998 NRC Canada
Robertson and Wride 443
Fig. 1. Schematic of undrained monotonic behaviour of sand in
triaxial compression (after Robertson 1994). LSS,
limitedstrain-softening response; qST, static gravitational shear
stress; Su, ultimate undrained shear strength; SH, strain-hardening
response; SS,strain-softening response; US, ultimate state.
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(Pando and Robertson 1995). When a soil element reachesthe
condition of essentially zero effective stress, the soil hasvery
little stiffness and large deformations can occur duringcyclic
loading. However, when cyclic loading stops, the de-formations
essentially stop, except for those due to localpore-pressure
redistribution. If there is no shear stress rever-sal, such as in
steeply sloping ground subjected to moderatecyclic loading, the
stress state may not reach zero effectivestress. As a result, only
cyclic mobility with limited defor-mations will occur, provided
that the initial void ratio of thesand is below the USL and the
large strain response is strainhardening (i.e., the material is not
susceptible to a cata-strophic flow slide). However, shear stress
reversal in thelevel ground area beyond the toe of a slope may lead
tooverall failure of the slope due to softening of the soil in
thetoe region.
Based on the above description of soil behaviour in un-drained
shear, Robertson and Fear (1995), building on ear-lier work by
Robertson (1994), proposed specific definitionsof soil liquefaction
which distinguished between flow lique-faction (strain-softening
behaviour; see Fig. 1) from cyclicsoftening. Cyclic softening was
further divided into cyclicliquefaction (see Fig. 2) and cyclic
mobility. A full descrip-tion of these definitions is given by
Robertson and Wride(1998) in the NCEER report (Youd and Idriss
1998). Fig-ure 3 presents a flow chart (after Robertson 1994) for
theevaluation of liquefaction according to these definitions.
Thefirst step is to evaluate the material characteristics in
terms
of a strain-softening or strain-hardening response. If the
soilis strain softening, flow liquefaction is possible if the
soilcan be triggered to collapse and if the gravitational
shearstresses are larger than the ultimate or minimum strength.The
trigger mechanism can be either monotonic or cyclic.Whether a slope
or soil structure will fail and slide will de-pend on the amount of
strain-softening soil relative tostrain-hardening soil within the
structure, the brittleness ofthe strain-softening soil, and the
geometry of the ground.The resulting deformations of a soil
structure with bothstrain-softening and strain-hardening soils will
depend onmany factors, such as distribution of soils, ground
geometry,amount and type of trigger mechanism, brittleness of
thestrain-softening soil, and drainage conditions. Soils that
areonly temporarily strain softening (i.e., experience a mini-mum
strength before dilating to US) are not as dangerous asvery loose
soils that can strain soften directly to ultimatestate. Examples of
flow liquefaction failures are Fort PeckDam (Casagrande 1965),
Aberfan flowslide (Bishop 1973),Zealand flowslide (Koppejan et al.
1948), and the Stava tail-ings dam. In general, flow liquefaction
failures are not com-mon; however, when they occur, they take place
rapidlywith little warning and are usually catastrophic. Hence,
thedesign against flow liquefaction should be carried out
cau-tiously.
If the soil is strain hardening, flow liquefaction will
gen-erally not occur. However, cyclic softening can occur due
tocyclic undrained loading, such as earthquake loading. The
1998 NRC Canada
444 Can. Geotech. J. Vol. 35, 1998
Fig. 2. Schematic of undrained cyclic behaviour of sand
illustrating cyclic liquefaction (after Robertson 1994). qcy,
cycling shear stress.
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amount and extent of deformations during cyclic loadingwill
depend on the density of the soil, the magnitude and du-ration of
the cyclic loading, and the extent to which shearstress reversal
occurs. If extensive shear stress reversal oc-curs, it is possible
for the effective stresses to reach zeroand, hence, cyclic
liquefaction can take place. When thecondition of essentially zero
effective stress is achieved,large deformations can result. If
cyclic loading continues,deformations can progressively increase.
If shear stress re-versal does not take place, it is generally not
possible toreach the condition of zero effective stress and
deformationswill be smaller, i.e., cyclic mobility will occur.
Examples ofcyclic softening were common in the major earthquakes
inNiigata in 1964 and Kobe in 1995 and manifested in theform of
sand boils, damaged lifelines (pipelines, etc.), lateralspreads,
slumping of small embankments, settlements, andground-surface
cracks. If cyclic liquefaction occurs anddrainage paths are
restricted due to overlying less permeablelayers, the sand
immediately beneath the less permeable soilcan loosen due to
pore-water redistribution, resulting in pos-sible subsequent flow
liquefaction, given the right geometry.
Both flow liquefaction and cyclic liquefaction can causevery
large deformations. Hence, it can be very difficult toclearly
identify the correct phenomenon based on observeddeformations
following earthquake loading. Earth-quake-induced flow liquefaction
movements tend to occurafter the cyclic loading ceases due to the
progressive natureof the load redistribution. However, if the soil
is sufficientlyloose and the static shear stresses are sufficiently
large, theearthquake loading may trigger essentially spontaneous
liq-uefaction within the first few cycles of loading. Also, if
the
soil is sufficiently loose, the ultimate undrained strengthmay
be close to zero with an associated effective confiningstress very
close to zero (Ishihara 1993). Cyclic liquefactionmovements, on the
other hand, tend to occur during the cy-clic loading, since it is
the inertial forces that drive the phe-nomenon. The post-earthquake
diagnosis can be furthercomplicated by the possibility of
pore-water redistributionafter the cyclic loading, resulting in a
change in soil densityand possibly the subsequent triggering of
flow liquefaction.Identifying the type of phenomenon after
earthquake loadingis difficult and, ideally, requires
instrumentation during andafter cyclic loading together with
comprehensive site charac-terization.
The most common form of soil liquefaction observed inthe field
has been cyclic softening due to earthquake load-ing. Much of the
existing research work on soil liquefactionhas been related to
cyclic softening, primarily cyclic lique-faction. Cyclic
liquefaction generally applies to level orgently sloping ground
where shear stress reversal occursduring earthquake loading. This
paper is concerned primar-ily with cyclic liquefaction due to
earthquake loading and itsevaluation using results of the CPT.
Much of the early work related to earthquake-induced
soilliquefaction resulted from laboratory testing of
reconstitutedsamples subjected to cyclic loading by means of
cyclictriaxial, cyclic simple shear, or cyclic torsional tests.
Theoutcome of these studies generally confirmed that the
resis-tance to cyclic loading is influenced primarily by the state
ofthe soil (i.e., void ratio, effective confining stresses, and
soilstructure) and the intensity and duration of the cyclic
load-ing (i.e., cyclic shear stress and number of cycles), as
wellas the grain characteristics of the soil. Soil structure
incorpo-rates features such as fabric, age, and cementation.
Graincharacteristics incorporate features such as grain-size
distri-bution, grain shape, and mineralogy.
Resistance to cyclic loading is usually represented interms of a
cyclic stress ratio that causes cyclic liquefaction,termed cyclic
resistance ratio (CRR). The point of liquefac-tion in a cyclic
laboratory test is typically defined as thetime at which the sample
achieves a strain level of either 5%double-amplitude axial strain
in a cyclic triaxial test or34% double-amplitude shear strain in a
cyclic simple sheartest. For cyclic simple shear tests, CRR is the
ratio of the cy-clic shear stress to cause cyclic liquefaction to
the initialvertical effective stress, i.e., (CRR)ss = cyc/vo . For
cyclictriaxial tests, CRR is the ratio of the maximum cyclic
shearstress to cause cyclic liquefaction to the initial effective
con-fining stress, i.e., (CRR)tx = dc/23c . The two tests
imposedifferent loading conditions and the CRR values are
notequivalent. Cyclic simple shear tests are generally consid-ered
to be better than cyclic triaxial tests at closely repre-senting
earthquake loading for level ground conditions.However, experience
has shown that the (CRR)ss can be esti-mated quite well from
(CRR)tx, and correction factors havebeen developed (Ishihara 1993).
The CRR is typically takenat about 15 cycles of uniform loading to
represent an equiva-lent earthquake loading of magnitude (M) 7.5,
i.e., CRR7.5.
1998 NRC Canada
Robertson and Wride 445
Fig. 3. Suggested flow chart for evaluation of soil
liquefaction(after Robertson 1994).
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The CRR for any other size earthquake can be estimatedusing the
following equation:[1] CRR CRR MSF)7.5=( ) (where MSF is the
magnitude scaling factor (recommendedvalues are provided in the
report by NCEER (Youd andIdriss 1998), which summarizes the results
of the 1996NCEER workshop).
When a soil is fine grained or contains some fines, somecohesion
or adhesion can develop between the fine particlesmaking the soil
more resistant at essentially zero effectiveconfining stress.
Consequently, a greater resistance to cyclicliquefaction is
generally exhibited by sandy soils containingsome fines. However,
this tendency depends on the nature ofthe fines contained in the
sand (Ishihara 1993). Laboratorytesting has shown that one of the
most important index prop-erties influencing CRR is the plasticity
index of the finescontained in the sand (Ishihara and Koseki 1989).
Ishihara(1993) showed that the (CRR)tx appears to increase with
in-creasing plasticity index. Studies in China (Wang 1979) sug-gest
that the potential for cyclic liquefaction in silts andclays is
controlled by grain size, liquid limit, and water con-tent. The
interpretation of this criterion as given byMarcuson et al. (1990)
and shown in Fig. 4 can be useful;however, it is important to note
that it is based on limiteddata and should be used with caution.
Figure 4 suggests thatwhen a soil has a liquid limit less than 35%
combined with awater content greater than 90% of the liquid limit,
it is un-clear if the soil can experience cyclic liquefaction, and
thesoil should be tested to clarify the expected response to
un-drained cyclic loading.
Although void ratio (relative density) has been recognizedas a
dominant factor influencing the CRR of sands, studiesby Ladd
(1974), Mulilis et al. (1977), and Tatsuoka et al.(1986) have
clearly shown that sample preparation (i.e., soilfabric) also plays
an important role. This is consistent withthe results of monotonic
tests at small to intermediate strainlevels. Hence, if results are
to be directly applied with anyconfidence, it is important to
conduct cyclic laboratory testson reconstituted samples with a
structure similar to that in
situ. Unfortunately, it is very difficult to determine the
insitu fabric of natural sands below the water table. As a re-sult,
there is often some uncertainty in the evaluation ofCRR based on
laboratory testing of reconstituted samples,although, as suggested
by Tokimatsu and Hosaka (1986), ei-ther the small strain shear
modulus or shear wave velocitymeasurements could be used to improve
the value of labora-tory testing on reconstituted samples of sand.
Therefore,there has been increasing interest in testing
high-quality un-disturbed samples of sandy soils under conditions
represen-tative of those in situ. Yoshimi et al. (1989) showed
thataging and fabric had a significant influence on the CRR ofclean
sand from Niigata. Yoshimi et al. (1994) also showedthat sand
samples obtained using conventional high-qualityfixed piston
samplers produced different CRR values thanthose of undisturbed
samples obtained using in situ groundfreezing. Dense sand samples
showed a decrease in CRRand loose sand samples showed an increase
in CRR whenobtained using a piston sampler, as compared with the
re-sults of testing in situ frozen samples. The difference inCRR
became more pronounced as the density of the sand in-creased.
Based on the above observations, for high-risk projectswhen
evaluation of the potential for soil liquefaction due toearthquake
loading is very important, consideration shouldbe given to a
limited amount of appropriate laboratory testson high-quality
undisturbed samples. Recently, in situground freezing has been used
to obtain undisturbed samplesof sandy soils (Yoshimi et al. 1978,
1989, 1994; Sego et al.1994; Hofmann et al. 1995; Hofmann 1997).
Cyclic simpleshear tests are generally the most appropriate tests,
althoughcyclic triaxial tests can also give reasonable results.
The above comments have shown that testing
high-qualityundisturbed samples will give better results than
testing poorquality samples. However, obtaining high-quality
undis-turbed samples of saturated sandy soils is very difficult
andexpensive and can only be carried out for large projects
forwhich the consequences of liquefaction may result in largecosts.
Therefore, there will always be a need for simple, eco-nomic
procedures for estimating the CRR of sandy soils,particularly for
low-risk projects and the initial screeningstages of high-risk
projects. Currently, the most popular sim-ple method for estimating
CRR makes use of penetration re-sistance from the standard
penetration test (SPT), although,more recently, the CPT has become
very popular because ofits greater repeatability and the continuous
nature of its pro-file.
The late Professor H.B. Seed and his coworkers devel-oped a
comprehensive SPT-based approach to estimate thepotential for
cyclic softening due to earthquake loading. Theapproach requires an
estimate of the cyclic stress ratio(CSR) profile caused by a design
earthquake. This is usuallydone based on a probability of
occurrence for a given earth-quake. A site-specific seismicity
analysis can be carried outto determine the design CSR profile with
depth. A simplifiedmethod to estimate CSR was also developed by
Seed andIdriss (1971) based on the maximum ground surface
acceler-
1998 NRC Canada
446 Can. Geotech. J. Vol. 35, 1998
Fig. 4. Graphical representation of liquefaction criteria for
siltsand clays from studies by Seed et al. (1973) and Wang (1979)
inChina (after Marcuson et al. 1990): < 15% finer than 0.005
mm,liquid limit (LL) < 35%, and water content > 0.9 liquid
limit.
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ation (amax) at the site. This simplified approach can be
sum-marized as follows:
[2] CSR avvo
max vo
vo
=
=
065. ag rd
where av is the average cyclic shear stress; amax is the
maxi-mum horizontal acceleration at the ground surface; g =9.81
m/s2 is the acceleration due to gravity; vo and vo arethe total and
effective vertical overburden stresses, respec-tively; and rd is a
stress-reduction factor which is dependenton depth. Details about
the rd factor are summarized byRobertson and Wride (1998), based on
the recommendationsby Seed and Idriss (1971) and Liao and Whitman
(1986).The CSR profile from the earthquake can be compared tothe
estimated CRR profile for the soil deposit, adjusted tothe same
magnitude using eq. [1]. At any depth, if CSR isgreater than CRR,
cyclic softening (liquefaction) is possible.This approach is the
most commonly used technique in mostparts of the world for
estimating soil liquefaction due toearthquake loading.
The approach based on the SPT has many problems, pri-marily due
to the inconsistent nature of the SPT. The mainfactors affecting
the SPT have been reviewed (e.g., Seed etal. 1985; Skempton 1986;
Robertson et al. 1983) and aresummarized by Robertson and Wride
(1998). It is highlyrecommended that the engineer become familiar
with the de-
tails of the SPT to avoid or at least minimize the effects
ofsome of the major factors, the most important of which isthe
energy delivered to the SPT sampler. A full descriptionof the
recommended modifications to the simplified SPTmethod to estimate
cyclic liquefaction is given by Robertsonand Wride (1998) in the
NCEER report (Youd and Idriss1998).
Cone penetration test (CPT)Because of the inherent difficulties
and poor repeatability
associated with the SPT, several correlations have been
pro-posed to estimate CRR for clean sands and silty sands
usingcorrected CPT penetration resistance (e.g., Robertson
andCampanella 1985; Seed and de Alba 1986; Olsen 1988;Olsen and
Malone 1988; Shibata and Teparaska 1988;Mitchell and Tseng 1990;
Olsen and Koester 1995; Suzuki etal. 1995a, 1995b; Stark and Olson
1995; Robertson and Fear1995).
Although cone penetration resistance is often just cor-rected
for overburden stress (resulting in the term qc1), trulynormalized
(i.e., dimensionless) cone penetration resistancecorrected for
overburden stress (qc1N) can be given by
[3] q qP
C qPQc1N
c
a2
c1
a2=
=
where qc is the measured cone tip penetration resistance; CQ=
(Pa/vo )n is a correction for overburden stress; the expo-nent n is
typically equal to 0.5; Pa is a reference pressure inthe same units
as vo (i.e., Pa = 100 kPa if vo is in kPa);and Pa2 is a reference
pressure in the same units as qc (i.e.,Pa2 = 0.1 MPa if qc is in
MPa). A maximum value of CQ = 2is generally applied to CPT data at
shallow depths. The nor-malized cone penetration resistance, qc1N,
is dimensionless.
Robertson and Campanella (1985) developed a chart forestimating
CRR from corrected CPT penetration resistance(qc1) based on the
Seed et al. (1985) SPT chart andSPTCPT conversions. Other similar
CPT-based charts werealso developed by Seed and de Alba (1986),
Shibata andTeparaska (1988), and Mitchell and Tseng (1990). A
com-parison between three of these CPT charts is shown inFig. 5. In
recent years, there has been an increase in avail-able field
performance data, especially for the CPT (Ishihara1993; Kayen et
al. 1992; Stark and Olson 1995; Suzuki et al.1995b). The recent
field performance data have shown thatthe existing CPT-based
correlations to estimate CRR aregenerally good for clean sands. The
recent field performancedata show that the correlation between CRR
and qc1N byRobertson and Campanella (1985) for clean sands provides
areasonable estimate of CRR. Based on discussions at the1996 NCEER
workshop, the curve by Robertson andCampanella (1985) has been
adjusted slightly at the lowerend to be more consistent with the
SPT curve. The resultingrecommended CPT correlation for clean sand
is shown inFig. 6. Included in Fig. 6 are suggested curves of
limitingshear strain, similar to those suggested by Seed et al.
(1985)for the SPT. Occurrence of liquefaction is based on
levelground observations of surface manifestations of cyclic
liq-uefaction. For loose sand (i.e., qc1N < 75) this could
involvelarge deformations resulting from a condition of
essentiallyzero effective stress being reached. For denser sand
(i.e.,
1998 NRC Canada
Robertson and Wride 447
Fig. 5. Comparison between three CPT-based charts forestimating
cyclic resistance ratio (CRR) for clean sands (afterIshihara 1993).
D50, average grain size; FC, fines content.
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qc1N > 75) this could involve the development of large
porepressures, but the effective stress may not fully reduce tozero
and deformations may not be as large as those in loosesands. Hence,
the consequences of liquefaction will vary de-pending on the soil
density as well as the size and durationof loading. An approximate
equation for the clean sand CPTcurve shown in Fig. 6 is given later
in this paper (eq. [8]).
Based on data from 180 sites, Stark and Olson (1995)
alsodeveloped a set of correlations between CRR and cone
tipresistance for various sandy soils based on fines content
andmean grain size, as shown in Fig. 7, in terms of qc1N. TheCPT
combined database is now larger than the originalSPT-based database
proposed by Seed et al. (1985).
The field observation data used to compile the CPT data-base are
apparently based on the following conditions, simi-lar in nature to
those for the SPT-based data: Holocene age,clean sand deposits;
level or gently sloping ground; magni-tude M = 7.5 earthquakes;
depth range from 1 to 15 m (345ft) (84% is for depths < 10 m (30
ft)); and representative av-erage CPT qc values for the layer that
was considered tohave experienced cyclic liquefaction.
Caution should be exercised when extrapolating the
CPTcorrelation to conditions outside of the above range. An
im-portant feature to recognize is that the correlation appears
tobe based on average values for the inferred liquefied
layers.However, the correlation is often applied to all measuredCPT
values, which include low values below the average, aswell as low
values as the cone moves through soil layer in-terfaces. Therefore,
the correlation can be conservative invariable deposits where a
small part of the CPT data couldindicate possible liquefaction.
Although some of the re-corded case histories show liquefaction
below the suggestedcurve in Fig. 6, the data are based on average
values and,
hence, the authors consider the suggested curve to be
consis-tent with field observations. It is important to note that
thesimplified approach based on either the SPT or the CPT hasmany
uncertainties. The correlations are empirical and thereis some
uncertainty over the degree of conservatism in thecorrelations as a
result of the methods used to select repre-sentative values of
penetration resistance within the layersassumed to have liquefied.
A detailed review of the CPTdata, similar to those carried out by
Liao and Whitman(1986) and Fear and McRoberts (1995) on SPT data,
wouldbe required to investigate the degree of conservatism
con-tained in Figs. 6 and 7. The correlations are also sensitive
tothe amount and plasticity of the fines within the sand.
For the same CRR, SPT or CPT penetration resistance insilty
sands is smaller because of the greater compressibilityand
decreased permeability of silty sands. Therefore, onereason for the
continued use of the SPT has been the need toobtain a soil sample
to determine the fines content of thesoil. However, this has been
offset by the poor repeatabilityof SPT data. With the increasing
interest in the CPT due toits greater repeatability, several
researchers (e.g., Robertsonand Campanella 1985; Olsen 1988; Olsen
and Malone 1988;Olsen and Koester 1995; Suzuki et al. 1995a, 1995b;
Starkand Olson 1995; Robertson and Fear 1995) have developeda
variety of approaches for evaluating cyclic liquefaction po-tential
using CPT results. It is now possible to estimate
graincharacteristics such as apparent fines content and grain
sizefrom CPT data and incorporate this directly into the
evalua-tion of liquefaction potential. Robertson and Fear
(1995)recommended an average correction, which was dependenton
apparent fines content, but not on penetration resistance.This
paper provides modifications to and an update of theCPT approach
suggested by Robertson and Fear (1995). Theproposed equation to
obtain the equivalent clean sand nor-
1998 NRC Canada
448 Can. Geotech. J. Vol. 35, 1998
Fig. 6. Recommended cyclic resistance ratio (CRR) for cleansands
under level ground conditions based on CPT. l, limitingshear
strain.
Fig. 7. Summary of variation of cyclic resistance ratio
(CRR)with fines content based on CPT field performance data
(afterStark and Olson 1995).
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malized CPT penetration resistance, (qc1N)cs, is a function
ofboth the measured penetration resistance, qc1N, and the
graincharacteristics of the soil, as follows:[4] ( )q K qc1N cs c
c1N=where Kc is a correction factor that is a function of the
graincharacteristics of the soil, as described later in this
paper.
Grain characteristics from the CPTIn recent years, charts have
been developed to estimate
soil type from CPT data (Olsen and Malone 1988; Olsen andKoester
1995; Robertson and Campanella 1988; Robertson1990). Experience has
shown that the CPT friction ratio (ra-tio of the CPT sleeve
friction to the cone tip resistance) in-creases with increasing
fines content and soil plasticity.Hence, grain characteristics such
as apparent fines content ofsandy soils can be estimated directly
from CPT data usingany of these soil behaviour charts, such as that
by Robertson(1990) shown in Fig. 8. As a result, the measured
penetra-tion resistance can be corrected to an equivalent clean
sandvalue. The addition of pore-pressure data can also
providevaluable additional guidance in estimating fines
content.Robertson et al. (1992) suggested a method for
estimatingfines content based on the rate of pore-pressure
dissipation(t50) during a pause in the CPT.
Based on extensive field data and experience, it is possi-ble to
estimate grain characteristics directly from CPT re-sults using the
soil behaviour type chart shown in Fig. 8.The boundaries between
soil behaviour type zones 27 canbe approximated as concentric
circles (Jefferies and Davies1993). The radius of each circle can
then be used as a soilbehaviour type index. Using the CPT chart by
Robertson(1990), the soil behaviour type index, Ic, can be defined
asfollows:
[5] I Q Fc - log= + +[( . ) ( . ) ] .347 1222 2 0 5
where Q qP
Pn
=
c vo
a2
a
vo
is the normalized CPT penetration resistance(dimensionless); the
exponent n is typically equal to 1.0; F =[fs/(qc vo)]100 is the
normalized friction ratio, in percent;fs is the CPT sleeve friction
stress; vo and vo are the totaland effective overburden stresses,
respectively; Pa is a refer-ence pressure in the same units as vo
(i.e., Pa = 100 kPa ifvo
is in kPa); and Pa2 is a reference pressure in the sameunits as
qc and vo (i.e., Pa2 = 0.1 MPa if qc and vo arein MPa).
The soil behaviour type chart by Robertson (1990) uses
anormalized cone penetration resistance (Q) based on a sim-ple
linear stress exponent of n = 1.0 (see above), whereasthe chart
recommended here for estimating CRR (see Fig. 6)is essentially
based on a normalized cone penetration resis-tance (qc1N) based on
a stress exponent n = 0.5 (see eq. [3]).Olsen and Malone (1988)
correctly suggested a normaliza-tion where the stress exponent (n)
varies from around 0.5 insands to 1.0 in clays. However, this
normalization for soiltype is somewhat complex and iterative.
The Robertson (1990) procedure using n = 1.0 is recom-mended for
soil classification in clay type soils when Ic >2.6. However, in
sandy soils when Ic 2.6, it is recom-mended that data being plotted
on the Robertson chart bemodified by using n = 0.5. Hence, the
recommended proce-dure is to first use n = 1.0 to calculate Q and,
therefore, aninitial value of Ic for CPT data. If Ic > 2.6, the
data should beplotted directly on the Robertson chart (and assume
qc1N =Q). However, if Ic 2.6, the exponent to calculate Q shouldbe
changed to n = 0.5 (i.e., essentially calculate qc1N usingeq. [3],
since vo < < qc) and Ic should be recalculated basedon qc1N
and F. If the recalculated Ic remains less than 2.6, thedata should
be plotted on the Robertson chart using qc1Nbased on n = 0.5. If,
however, Ic iterates above and below avalue of 2.6, depending which
value of n is used, a value ofn = 0.75 should be selected to
calculate qc1N (using eq. [3])and plot data on the Robertson chart.
Note that if the in situeffective overburden stresses are in the
order of 50150 kPa,the choice of normalization has little effect on
the calculatednormalized penetration resistance.
The boundaries of soil behaviour type are given in termsof the
index, Ic, as shown in Table 1. The soil behaviour typeindex does
not apply to zones 1, 8, or 9. Along the normallyconsolidated
region in Fig. 8, soil behaviour type index in-creases with
increasing apparent fines content and soil plas-ticity, and the
following simplified relationship is suggested:
1998 NRC Canada
Robertson and Wride 449
Fig. 8. Normalized CPT soil behaviour type chart, as proposedby
Robertson (1990). Soil types: 1, sensitive, fine grained; 2,peats;
3, silty clay to clay; 4, clayey silt to silty clay; 5, siltysand
to sandy silt; 6, clean sand to silty sand; 7, gravelly sand
todense sand; 8, very stiff sand to clayey sand
(heavilyoverconsolidated or cemented); 9, very stiff, fine
grained(heavily overconsolidated or cemented). OCR,
overconsolidationratio; , friction angle.
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[6a] if apparent fines content FCcI < =126 0. (%)[6b] if c126
35. . I
apparent fines content FC c(%) . ..= 175 373 25I[6c] if apparent
fines content FCcI > =35 100. (%)
The range of potential correlations is illustrated in Fig.
9,which shows the variation of soil behaviour type index (Ic)with
apparent fines content and the effect of the degree ofplasticity of
the fines. The recommended general relation-ship given in eq. [6]
is also shown in Fig. 9. Note that thisequation is slightly
modified from the original work by Rob-ertson and Fear (1995) to
increase the prediction of apparentFC for a given value of Ic.
The proposed correlation between CPT soil behaviour in-dex (Ic)
and apparent fines content is approximate, since theCPT responds to
many other factors affecting soil behaviour,such as soil
plasticity, mineralogy, sensitivity, and stress his-tory. However,
for small projects, the above correlation pro-vides a useful guide.
Caution must be taken in applying
eq. [6] to sands that plot in the region defined by 1.64 < Ic
3.60 2 Organic soils: peats
Table 1. Boundaries of soil behaviour type (after Robertson
1990).
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The proposed correction factor, Kc, is approximate, sincethe CPT
responds to many factors, such as soil plasticity,fines content,
mineralogy, soil sensitivity, and stress history.However, for small
projects or for initial screening on largerprojects, the above
correlation provides a useful guide. Cau-tion must be taken in
applying the relationship to sands thatplot in the region defined
by 1.64 < Ic < 2.36 and F 0.5%so as not to confuse very loose
clean sands with sands con-taining fines. In this zone, it is
suggested that the correctionfactor Kc be set to a value of 1.0
(i.e., assume that the sandis a clean sand).
Note that the relationship between the recommended cor-rection
factor, Kc, and soil behaviour type index, Ic, is shownin Fig. 10
as a broken line beyond an Ic of 2.6, which corre-sponds to an
approximate apparent fines content of 35%.Soils with Ic > 2.6
fall into the clayey silt, silty clay, andclay regions of the CPT
soil behaviour chart (i.e., zones 3and 4). When the CPT indicates
soils in these regions (Ic >2.6), samples should be obtained and
evaluated using criteriasuch as those shown in Fig. 4. It is
reasonable to assume, ingeneral, that soils with Ic > 2.6 are
nonliquefiable and thatthe correction Kc could be large. Soils that
fall in the lowerleft region of the CPT soil behaviour chart (Fig.
8), definedby Ic > 2.6 and F 1.0%, can be very sensitive and,
hence,possibly susceptible to both cyclic and (or) flow
liquefac-tion. Soils in this region should be evaluated using
criteriasuch as those shown in Fig. 4 combined with additional
test-ing.
Figure 11 shows the resulting equivalent CRR curves forIc values
of 1.64, 2.07, and 2.59 which represent approxi-mate apparent fines
contents of 5, 15, and 35%, respectively.
Influence of thin layersA problem associated with the
interpretation of penetra-
tion tests in interbedded soils occurs when thin sand layersare
embedded in softer deposits. Theoretical as well as labo-ratory
studies show that the cone resistance is influenced bythe soil
ahead of and behind the penetrating cone. The conewill start to
sense a change in soil type before it reaches thenew soil and will
continue to sense the original soil evenwhen it has entered a new
soil. As a result, the CPT will notalways measure the correct
mechanical properties in thinlyinterbedded soils. The distance over
which the cone tipsenses an interface increases with increasing
soil stiffness. Insoft soils, the diameter of the sphere of
influence can be assmall as two to three cone diameters, whereas in
stiff soilsthe sphere of influence can be up to 20 cone
diameters.Hence, the cone resistance can fully respond (i.e., reach
fullvalue within the layer) in thin soft layers better than in
thinstiff layers. Therefore care should be taken when interpret-ing
cone resistance in thin sand layers located within softclay or silt
deposits. Based on a simplified elastic solution,Vreugdenhil et al.
(1994) have provided some insight as tohow to correct cone data in
thin layers. Vreugdenhil et al.have shown that the error in the
measured cone resistancewithin thin stiff layers is a function of
the thickness of the
1998 NRC Canada
Robertson and Wride 451
Fig. 10. Recommended grain characteristic correction to obtain
clean sand equivalent CPT penetration resistance in sandy
soils.
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1998 NRC Canada
452 Can. Geotech. J. Vol. 35, 1998
layer as well as the stiffness of the layer relative to that
ofthe surrounding softer soil. The relative stiffness of the
lay-ers is reflected by the change in cone resistance from thesoft
surrounding soil to the stiff soil in the layer.Vreugdenhil et al.
validated the model with laboratory andfield data.
Based on the work by Vreugdenhil et al. (1994), Robert-son and
Fear (1995) suggested a conservative correction fac-tor for cone
resistance. The corrections apply only to thinsand layers embedded
in thick, fine-grained layers. A fulldescription of the proposed
correction factor is given byRobertson and Wride (1998) in the
NCEER report (Youdand Idriss 1998). Thin sand layers embedded in
soft clay de-posits are often incorrectly classified as silty sands
based onthe CPT soil behaviour type charts. Hence, a slightly
im-proved classification can be achieved if the cone resistanceis
first corrected for layer thickness before applying the
clas-sification charts.
Cyclic resistance from the CPTIn an earlier section, a method
was suggested for estimat-
ing apparent fines content directly from CPT results, usingeq.
[6]. Following the traditional SPT approach, the esti-mated
apparent fines content could be used to estimate thecorrection
necessary to obtain the clean sand equivalent pen-etration
resistance. However, since other grain characteris-tics also
influence the measured CPT penetration resistance,it is recommended
that the necessary correction be estimatedfrom the soil behaviour
type index, as described above.Hence, eqs. [4], [5], and [7] can be
combined to estimate theequivalent clean sand normalized
penetration resistance,
(qc1N)cs, directly from the measured CPT data. Then, usingthe
equivalent clean sand normalized penetration resistance(qc1N)cs,
the CRR (for M = 7.5) can be estimated using thefollowing
simplified equation (which approximates the cleansand curve
recommended in Fig. 6):[8a]if CRRc1N cs c1N cs50 160 93 1000
0083
< = +( ) ,
( ).q q
[8b]if CRRc1N cs c1N cs( ) , . ( ) .q q< =
+50 0833 1000 005
3
In summary, eqs. [4][7] and [8] can be combined to pro-vide an
integrated method for evaluating the cyclic resis-tance (M = 7.5)
of saturated sandy soils based on the CPT. Ifthin layers are
present, corrections to the measured tip resis-tance in each thin
layer may be appropriate. The CPT-basedmethod is an alternative to
the SPT or shear wave velocity(Vs) based in situ methods; however,
using more than onemethod is useful in providing independent
evaluations ofliquefaction potential. The proposed integrated CPT
methodis summarized in Fig. 12 in the form of a flow chart. Theflow
chart clearly shows the step-by-step process involvedin using the
proposed integrated method based on the CPTfor evaluating CRR and
indicates the recommended equa-tions for each step of the
process.
Although the proposed approach provides a method tocorrect CPT
results for grain characteristics, it is not in-tended to remove
the need for selected sampling. If the userhas no previous CPT
experience in the particular geologicregion, it is important to
take carefully selected samples toevaluate the CPT soil behaviour
type classification.Site-specific modifications are recommended,
where possi-ble. However, if extensive CPT experience exists within
thegeologic region, the modified CPT method with no samplescan be
expected to provide an excellent guide to evaluateliquefaction
potential.
Application of the proposed CPT methodAn example of this
proposed modified CPT-based method
is shown in Fig. 13 for the Moss Landing site that
sufferedcyclic liquefaction during the 1989 Loma Prieta
earthquakein California (Boulanger et al. 1995, 1997). The
measuredcone resistance is normalized and corrected for
overburdenstress to qc1N and F and the soil behaviour type index
(Ic) iscalculated. The final continuous profile of CRR at N =
15cycles (M = 7.5) is calculated from the equivalent clean
sandvalues of qc1N (i.e., (qc1N)cs = Kcqc1N) and eq. [8].
Includedin Fig. 13 are measured fines content values obtained
fromadjacent SPT samples. A reasonable comparison is seen be-tween
the estimated apparent fines contents and the mea-sured fines
contents. Note that for Ic > 2.6 (i.e., FC > 35%;see eq.
[6]), the soil is considered to be nonliquefiable; how-ever, this
should be checked using other criteria (e.g.,Marcuson et al. 1990)
(see Fig. 4). The estimated zones ofsoil that are predicted to
experience cyclic liquefaction arevery similar to those observed
and reported by Boulanger etal. (1995, 1997).
Fig. 11. CPT base curves for various values of soil
behaviourindex, Ic (corresponding to various apparent fines
contents, asindicated).
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1998 NRC Canada
Robertson and Wride 453
The predicted zones of liquefaction are slightly conserva-tive
compared to the observed ground response. For exam-ple, when the
CPT passes from a sand into a clay (e.g., at adepth of 10.5 m), the
cone resistance decreases in the transi-
tion zone and the method predicts a very small zone of pos-sible
liquefaction. However, given the continuous nature ofthe CPT and
the high frequency of data (typically every20 mm), these erroneous
predictions are easy to identify. In
Fig. 12. Flow chart illustrating the application of the
integrated CPT method of evaluating cyclic resistance ratio (CRR)
in sandy soils.Vs, shear wave velocity.
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NRC
Canada
454Can.
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J.Vol.
35,1998
Fig. 13. Application of the integrated CPT method for estimating
cyclic resistance ratio (CRR) to the State Beach site at Moss
Landing which suffered cyclic liquefactionduring the 1989 Loma
Prieta earthquake in California.
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-
general, given the complexity of the problem, the
proposedmethodology provides a good overall prediction of the
po-tential for liquefaction. Since the methodology should be
ap-plied to low-risk projects and in the initial screening stagesof
high-risk projects, it is preferred that the method pro-duces, in
general, slightly conservative results.
Comparison with other CPT methodsOlsen (1988), Olsen and Koester
(1995), and Suzuki et al.
(1995b) have also suggested integrated methods to estimatethe
CRR of sandy soils directly from CPT results with thecorrelations
presented in the form of soil behaviour charts.The Olsen and
Koester method is based on SPTCPT con-versions plus some
laboratory-based CRR data. The methodby Suzuki et al. is based on
limited field observations. Themethods by Olsen and Koester and
Suzuki et al. are shownin Fig. 14. The Olsen and Koester method
uses a variablenormalization technique, which requires an iterative
process
to determine the normalization. The method by Suzuki et al.uses
the qc1N normalization suggested in this paper (eq. [3]with n =
0.5). The Olsen and Koester method is very sensi-tive to small
variations in measured friction ratio and theuser is not able to
adjust the correlations based onsite-specific experience. The
friction sleeve measurement forthe CPT can vary somewhat depending
on specific CPTequipment and tolerance details between the cone and
thesleeve and, hence, can be subject to some uncertainty. Themethod
proposed in this paper is based on field observationsand is
essentially similar to those of Olsen and Koester andSuzuki et al.;
however, the method described here is slightlymore conservative and
the process has been broken downinto its individual components.
Built into each of the CPT methods for estimating CRR isthe step
of correcting the measured cone tip resistance to aclean sand
equivalent value. It is the size of this correctionthat results in
the largest differences between predicted val-
1998 NRC Canada
Robertson and Wride 455
Fig. 14. Comparison of estimating CRR from the CPT. (a) CRR =
0.05 to 0.3, after Olsen and Koester (1995). c, stress exponent.(b)
CRR = 0.15 and 0.25, after Suzuki et al. (1995b). v , vertical
effective stress.
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1998 NRC Canada
456 Can. Geotech. J. Vol. 35, 1998
ues of CRR from the various methods. Figure 15 providesan
approximate comparison between the methods by Starkand Olson
(1995), Olsen and Koester (1995), and Suzuki etal. (1995b), in
terms of Kc from soil behaviour type index,Ic. Also superimposed in
Fig. 15 is the recommended rela-tionship, based on the equations
given in this paper. Thecomparisons are approximate because the
different authorsuse different normalizations when correcting CPT
tip resis-tance for plotting data on a soil classification chart.
How-ever, for effective overburden stresses in the order of100 kPa
(-1 tsf (tsf= ton-force per square foot)), all of thenormalization
methods should give similar values of normal-ized penetration
resistance for the same value of measuredpenetration resistance.
The boundaries between different soilbehaviour types are also
indicated in Fig. 15, as a guide tothe type of soil in which
corrections of certain magnitudesare suggested by the various
methods.
The methods by Suzuki et al. (1995b) and Olsen andKoester (1995)
appear to be consistent with each other, indi-cating that Kc
increases with increasing CRR. Very largecorrections result
especially in soils of high Ic. The methodby Stark and Olson (1995)
indicates that Kc decreases withincreasing CRR for a given Ic and
does not fit in with thetrend of the combined Suzuki et al. and
Olsen and Koesterlines. The magnitudes of the corrections suggested
by Stark
and Olson are generally smaller than those of the
othermethods.
Figure 15 indicates that the recommended correction isgenerally
more conservative than the corrections proposedby the other
authors. The other methods generally predicthigher values of Kc and
suggest that corrections should beapplied beginning at lower values
of Ic, particularly forhigher values of CRR. Note that the
recommended relation-ship between Kc and Ic is shown in Fig. 15 as
a broken linebeyond Ic = 2.6, which corresponds to an apparent FC
of35% (eq. [6]). This shows that the integrated CPT methodfor
evaluating CRR, as outlined here, does not apply to soilsthat would
be classified as clayey silt, silty clay, or clay. Asexplained
earlier, when interpretation of the CPT indicatesthat these types
of soils are present, samples should be ob-tained and evaluated
using other criteria, such as those givenin Fig. 4 (Marcuson et al.
1990). It is logical that innonliquefiable clay soils, the
equivalent correction factor,Kc, could be very large for Ic >
2.6.
For low-risk, small-scale projects, the potential for
cyclicliquefaction can be estimated using penetration tests such
asthe CPT. The CPT is generally more repeatable than the SPT
Fig. 15. Approximate comparison of various methods for
correcting CPT tip resistance to clean sand equivalent values,
based on soilbehaviour type.
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and is the preferred test, where possible. The CPT
providescontinuous profiles of penetration resistance, which are
use-ful for identifying soil stratigraphy and for providing
contin-uous profiles of estimated cyclic resistance ratio
(CRR).Corrections are required for both the SPT and CPT for
graincharacteristics, such as fines content and plasticity. For
theCPT, these corrections are best expressed as a function ofsoil
behaviour type index, Ic, which is affected by a varietyof grain
characteristics.
For medium- to high-risk projects, the CPT can be usefulfor
providing a preliminary estimate of liquefaction potentialin sandy
soils. For higher risk projects, and in cases wherethere is no
previous CPT experience in the geologic region,it is also preferred
practice to drill sufficient boreholes adja-cent to CPT soundings
to verify various soil types encoun-tered and to perform index
testing on disturbed samples. Aprocedure has been described to
correct the measured coneresistance for grain characteristics based
on the CPT soil be-haviour type index, Ic. The corrections are
approximate,since the CPT responds to many factors affecting soil
behav-iour. Expressing the corrections in terms of soil
behaviourindex is the preferred method of incorporating the effects
ofvarious grain characteristics, in addition to fines content.When
possible, it is recommended that the corrections beevaluated and
modified to suit a specific site and project.However, for
small-scale, low-risk projects and in the initialscreening process
for higher risk projects, the suggested gen-
eral corrections provide a useful guide. Correcting CPT re-sults
in thin sand layers embedded in softer fine-grained de-posits may
also be appropriate. The CPT is generally limitedto sandy soils
with limited gravel contents. In soils with highgravel contents,
penetration may be limited.
A summary of the CPT method is shown in Fig. 16, whichidentifies
the zones in which soils are susceptible to cyclicliquefaction
(primarily zone A). In general, soils with Ic >2.6 and F >
1.0% (zone B) are likely nonliquefiable. Soilsthat plot in the
lower left portion of the chart (zone C; Ic >2.6 and F <
1.0%) may be susceptible to cyclic and (or) flowliquefaction due to
the sensitive nature of these soils. Soilsin this region should be
evaluated using other criteria. Cau-tion should also be exercised
when extrapolating the sug-gested CPT correlations to conditions
outside of the rangefrom which the field performance data were
obtained. Animportant feature to recognize is that the correlations
appearto be based on average values for the inferred liquefied
lay-ers. However, the correlations are often applied to all
mea-sured CPT values, which include low values below theaverage for
a given sand deposit and at the interface be-tween different soil
layers. Hence, the correlations could beconservative in variable
stratified deposits where a smallpart of the penetration data could
indicate possible liquefac-tion.
As mentioned earlier, it is clearly useful to evaluate CRRusing
more than one method. For example, the seismic CPTcan provide a
useful technique for independently evaluatingliquefaction
potential, since it measures both the usual CPTparameters and shear
wave velocities within the same bore-hole. The CPT provides
detailed profiles of cone tip resis-tance, but the penetration
resistance is sensitive to graincharacteristics, such as fines
content and soil mineralogy,and hence corrections are required. The
seismic part of theCPT provides a shear wave velocity profile
typically aver-aged over 1 m intervals and, therefore, contains
less detailthan the cone tip resistance profile. However, shear
wave ve-locity is less influenced by grain characteristics and few
orno corrections are required (Robertson et al. 1992; Andrusand
Stokoe 1998). Shear wave velocity should be measuredwith care to
provide the most accurate results possible, sincethe estimated CRR
is sensitive to small changes in shearwave velocity. There should
be consistency in the liquefac-tion evaluation using either method.
If the two methods pro-vide different predictions of CRR profiles,
samples shouldbe obtained to evaluate the grain characteristics of
the soil.
A final comment to be made is that a key advantage of
theintegrated CPT method described here is that the algorithmscan
easily be incorporated into a spreadsheet. As illustratedby the
Moss Landing example presented here, the result is astraightforward
method for analyzing entire CPT profiles ina continuous manner.
This provides an useful tool for the en-gineer to review the
potential for cyclic liquefaction across asite using engineering
judgement.
The authors appreciate the contributions of the membersof the
1996 NCEER Workshop on Evaluation of Liquefac-tion Resistance of
Soils (T.L. Youd, Chair) which was heldin Salt Lake City, Utah.
Particular appreciation is extended
1998 NRC Canada
Robertson and Wride 457
Fig. 16. Summary of liquefaction potential on soil
classificationchart by Robertson (1990). Zone A, cyclic
liquefaction possible,depending on size and duration of cyclic
loading; zone B,liquefaction unlikely, check other criteria; zone
C, flowliquefaction and (or) cyclic liquefaction possible,
depending onsoil plasticity and sensitivity as well as size and
duration ofcyclic loading.
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