Recent Advances in Soil Liquefaction Engineering and Seismic Site Response Evaluation
Recent Advances in Soil Liquefaction Engineering and Seismic Site Response Evaluation
Recent Advances in Soil Liquefaction Engineering and Seismic Site Response Evaluation
Engineering and Soil Dynamics and Symposium in Honor of Professor W.D. Liam Finn
San Diego, California, March 26-31, 2001
ABSTRACT VOLUME
ABSTRACT
Over the past decade, major advances have occurred in both understanding and practice with regard to engineering treatment of
seismic soil liquefaction and assessment of seismic site response. Seismic soil liquefaction engineering has evolved into a sub-field in
its own right, and assessment and treatment of site effects affecting seismic site response has gone from a topic of controversy to a
mainstream issue addressed in most modern building codes and addressed in both research and practice. This rapid evolution in the
treatment of both liquefaction and site response issues has been pushed by a confluence of lessons and data provided by a series of
earthquakes over the past eleven years, as well as by the research and professional/political will engendered by these major seismic
events. Although the rate of progress has been laudable, further advances are occurring, and more remains to be done. As we enter a
“new millennium”, engineers are increasingly well able to deal with important aspects of these two seismic problem areas. This paper
highlights a number of major recent and ongoing developments in each of these two important areas of seismic practice, and offers
insights regarding work/research in progress, as well as suggestions regarding further advances needed. The first part of the paper
addresses soil liquefaction, and the second portion (briefly) addresses engineering assessment of seismic site response.
ABSTRACT
Over the past decade, major advances have occurred in both understanding and practice with regard to engineering treatment of
seismic soil liquefaction and assessment of seismic site response. Seismic soil liquefaction engineering has evolved into a sub-field in
its own right, and assessment and treatment of site effects affecting seismic site response has gone from a topic of controversy to a
mainstream issue addressed in most modern building codes and addressed in both research and practice. This rapid evolution in the
treatment of both liquefaction and site response issues has been pushed by a confluence of lessons and data provided by a series of
earthquakes over the past eleven years, as well as by the research and professional/political will engendered by these major seismic
events. Although the rate of progress has been laudable, further advances are occurring, and more remains to be done. As we enter a
“new millenium”, engineers are increasingly well able to deal with important aspects of these two seismic problem areas. This paper
will highlight a few major recent and ongoing developments in each of these two important areas of seismic practice, and will offer
insights regarding work/research in progress, as well as suggestions regarding further advances needed. The first part of the paper will
address soil liquefaction, and the second portion will (briefly) address engineering assessment of seismic site response.
0.4 0.4
0.3 0.3
CSRN
CSRN
0.2 0.2
0 0
0 10 20 30 40 0 10 20 30 40
(N1)60 (N1)60,cs
0.4
5%
0.3
CSRN
0.2
0
0 10 20 30 40
(N1)60,cs
Fig. 5: Comparison of Best Available Probabilistic Correlations for Evaluation of Liquefaction Potential
(All Plotted for Mw=7.5, σv' = 1300 psf, and Fines Content ≤ 5%)
not only provide greatly reduced uncertainty, they also help to obtained and studied. Additional cases were also obtained,
resolve a number of corollary issues that have long been including several proprietary data sets. Eventually,
difficult and controversial, including: (1) magnitude-correlated approximately 450 liquefaction (and “non-liquefaction”) field
duration weighting factors, (2) adjustments for fines content, case histories were evaluated in detail. A formal rating system
and (3) corrections for effective overburden stress. was established for rating these case histories on the basis of
data quality and uncertainty, and standards were established
As a starting point, all of the field case histories employed in for inclusion of field cases in the final data set used to
the correlations shown in Figures 4 and 5(a) through (c) were establish the new correlations. In the end, 201 of the field
a σ
CSRpeak = max ⋅ v ⋅ (rd ) (Eq. 1)
g σ v′
where
amax = the peak horizontal ground surface acceleration,
g = the acceleration of gravity,
σv = total vertical stress,
σ′v = effective vertical stress, and
rd = the nonlinear shear mass participation factor.
1 +
0.104⋅( 0.0785⋅Vs*, 40′ + 24.888)
16.258 + 0.201⋅ e
d ≥ 65 ft:
1 +
0.104⋅( 0.0785⋅Vs*, 40′ + 24.888)
16.258 + 0.201⋅ e
where
σ ε r (d) = d 0.850 ⋅ 0.0072 [for d < 40 ft], and σ ε r (d) = 40 0.850 ⋅ 0.0072 [for d ≥ 40 ft]
d d
soil strata for most of the important liquefaction (and non- unbiased correlation of Equation 2, there is no intrinsic a priori
liquefaction) earthquake field case histories occur. This, in bias associated with either approach.
turn, creates some degree of corresponding bias in
relationships developed on this basis. In these new correlations, in-situ cyclic stress ratio (CSR) is
taken as the “equivalent uniform CSR” equal to 65% of the
Cetin and Seed (2000, 2001) propose a new, empirical basis single (one-time) peak CSR (from Equation 1) as
for estimation of rd as a function of; (1) depth, (2) earthquake
magnitude, (3) intensity of shaking, and (4) site stiffness (as CSR eq = (0.65) ⋅ CSR peak (Eq. 3)
expressed in Equation 2).
Figure 7 shows the values of rd from the 2,153 site response In-situ CSReq was evaluated directly, based on performance of
analyses performed as part of these studies sub-divided into 12 full seismic site response analyses (using SHAKE 90; Idriss
“bins” as a function of peak ground surface acceleration (amax), and Sun, 1992), for cases where (a) sufficient sub-surface data
site stiffness (VS,40ft), earthquake magnitude (Mw), and depth was available, and (b) where suitable “input” motions could be
(d). [VS,40ft is the “average” shear wave velocity over the top developed from nearby strong ground motion records. For
40 feet of a site (in units of ft./sec.), taken as 40 feet divided cases wherein full seismic site response analyses were not
by the shear wave travel time in traversing this 40 feet.] performed, CSReq was evaluated using the estimated amax and
Superimposed on each figure are the mean and + 1 standard Equations 1 and 2. In addition to the best estimates of CSReq,
deviation values central to each “bin” from Equation 2. Either the variance or uncertainty of these estimates (due to all
Equation 2, or Figure 7, can be used to derive improved (and contributing sources of uncertainty) was also assessed (Cetin
statistically unbiased) estimates of rd. et al., 2001).
It is noted, however, that in-situ CSR (and rd) can “jump” or At each case history site, the critical stratum was identified as
transition irregularly within a specific soil profile, especially the stratum most susceptible to triggering of liquefaction.
near sharp transitions between “soft” and “stiff” strata, and When possible, collected surface boil materials were also
that CSR (and rd) are also a function of the interaction between considered, but problems associated with mixing and
a site and each specific excitation motion. Accordingly, the segregation during transport, and recognition that liquefaction
best means of estimation of in-situ CSR within any given of underlying strata can result in transport of overlying soils to
stratum is to directly calculate CSR by means of appropriate the surface through boils, limited the usefulness of some of
site-specific, and event-specific, seismic site response this data.
analyses, when this is feasible. As the new correlations were
developed using both directly-calculated rd values (from site
response analyses) as well as rd values from the statistically
(c) Mw<6.8, amax ≤0.12g, Vs,40 ft. ≤525 fps (d) Mw<6.8, amax ≤0.12g, Vs,40 ft. >525 fps
Fig. 7: Rd Results for Various “Bins” Superimposed with the Predictions (Mean and Mean ±1σ
σ) Based
on Bin Mean Values of Vs,40 ft , Mw, and amax (continued…)
(g) Mw<6.8, 0.12< amax ≤0.23g, Vs,40 ft. ≤525 fps (h) Mw<6.8, 0.12< amax ≤0.23g, V.s,40 ft. >525 fps
Fig. 7: Rd Results for Various “Bins” Superimposed with the Predictions (Mean and Mean ±1σ
σ) Based
on Bin Mean Values of Vs,40 ft , Mw, and amax (continued…)
(k) Mw<6.8, 0.23< amax, Vs,40 ft. ≤525 fps (l) Mw<6.8, 0.23< amax, Vs,40 ft. >525 fps
Fig. 7: Rd Results for Various “Bins” Superimposed with the Predictions (Mean and Mean ±1σ
σ) Based
on Bin Mean Values of Vs,40 ft , Mw, and amax
σ’v 0
0 .7 0 .8 0 .9 1
(Eq. 4(b))
5
where σ’v is the actual effective overburden stress at the depth
of the SPT in atmospheres.
10
Rod Length (m)
values as
20
N1,60 = N1 ⋅ C R ⋅ C S ⋅ C B ⋅ C E (Eq. 5)
25
• The best approach is to directly measure the impact energy transmitted with
each blow. When available, direct energy measurements were employed.
• The next best approach is to use a hammer and mechanical hammer release
system that has been previously calibrated based on direct energy
measurements.
• Otherwise, ER must be estimated. For good field procedures, equipment and
monitoring, the following guidelines are suggested:
• For lesser quality fieldwork (e.g. irregular hammer drop distance, sliding
friction of hammer on rods, wet or worn rope on cathead, etc.) further
judgmental adjustments are needed.
Notes: (1) Based on rope and cathead system, two turns of rope around cathead, “normal” release
(not the Japanese “throw”), and rope not wet or excessively worn.
(2) Rope and cathead with special Japanese “throw” release. (See also Note 4)
(3) For the ranges shown, values roughly central to the mid-third of the range are more
common than outlying values, but ER and CE can be even more highly variable than the
ranges shown if equipment and/or monitoring and procedures are not good.
(4) Common Japanese SPT practice requires additional corrections for borehole diameter
and for frequency of SPT hammer blows. For “typical” Japanese practice with rope
and cathead, donut hammer, and the Japanese “throw” release, the overall product of
CB ∗ CE is typically in the range of 1.0 to 1.3.
In these current studies, based on the overall (regressed) The duration weighting factors shown in Figures 10(a) and (b)
correlation, the energy- and procedure- and overburden- fall slightly below those recommended by the NCEER
corrected N-values (N1,60) are further corrected for fines Working group, and slightly above (but very close to) recent
content as recommendations of Idriss (2000). Idriss’ recommendations
are based on a judgmental combination of interpretation of
N1,60,CS = N1,60 * CFINES (Eq. 6) high-quality cyclic simple shear laboratory test data and
empirical assessment of “equivalent” numbers of cycles from
recorded strong motion time histories, and are the only other
where the fines correction was “regressed” as a part of the values shown that account for the cross-correlation of rd with
Bayesian updating analyses. The fines correction is equal to magnitude. The close agreement of this very different (and
zero for fines contents of FC < 5%, and reaches a maximum principally laboratory data based) approach, and the careful
(limiting) value for FC > 35%. As illustrated in Figure 9(b), (field data based) probabilistic assessments of these current
the maximum fines correction results in an increase of N- studies, are strongly mutually supportive.
values of about +6 blows/ft. (at FC > 35%, and high CSR).
As illustrated in this figure, this maximum fines correction is Adjustments for Effective Overburden Stress:
somewhat smaller than the earlier maximum correction of
+9.5 blows/ft proposed by Seed et al. (1984). An additional factor not directly resolved in prior studies
based on field case histories is the increased susceptibility of
The regressed relationship for CFINES is soils to cyclic liquefaction, at the same CSR, with increases in
effective overburden stress. This is in addition to the
FC normalization of N-values for overburden effects as per
C FINES = (1 + 0.004 ⋅ FC ) + 0.05 ⋅
Equation 4.
N
1, 60
The additional effects of reduction of normalized liquefaction
lim: FC ≥ 5% and FC ≤ 35% (Eq. 7) resistance with increased effective initial overburden stress
(σ’v) has been demonstrated by means of laboratory testing,
where FC = percent fines content (percent by dry weight finer and this is a manifestation of “critical state” type of behavior
Youd and Noble (1977) PL<50 These current studies were not very sensitive to Kσ, as the
range of σ’v in the case history data base was largely between
σ’v = 600 to 2,600 lb/ft2, but it was possible to “regress” Kσ as
part of the Bayesian updating. The results are shown in Figure
14, over the range of σ’v ≈ 600 to 3,600 lb/ft2 for which they
are considered valid. These are in good agreement with the
earlier recommendations of Figure 13, and it is recommended
that Kσ can be estimated as
1.5
For correctness, and to avoid ambiguity, both the earlier
relationship of Seed et al. (1984), and the correlations
developed in these current studies, need to be formally
1
normalized to σ’v = 1 atm. Accordingly, in these studies, all
data are corrected for K σ-effects (by Equations 9 and 10); not
0.5 just those data for which σ’v was greater than 1 atm. A
recommended limit is Kσ < 1.5 (at very shallow depths.)
0 Figures 12 and 13 show the proposed new correlations, this
4.5 5.5 6.5 7.5 8.5 time for σ’v =1 atm, and these figures represent the final, fully
Mw normalized recommended correlations.
Fig. 10(b): Recommended Magnitude-Correlated The overall correlation can be expressed in parts, as in the
Duration Weighting Factor as a Function previous sections (and Equations 6 - 12, and Figures 7 - 12).
of N1,60 It can also be expressed concisely as a single, composite
relationship as shown in Equation 11.
N 1, 60 ⋅ (1 + 0.004 ⋅ FC ) − 29.53 ⋅ ln (M w )
− 3.70 ⋅ ln (σ v′ ) + 0.05 ⋅ FC + 44.97 + 2.70 ⋅ Φ (PL ) (Eq. 12)
−1
where
Φ-1(PL) = the inverse of the standard cumulative normal distribution (i.e. mean=0, and standard deviation=1)
note: for spreadsheet purposes, the command in Microsoft Excel for this specific function is “NORMINV(PL,0,1)”
` Fig. 11: Recommended “Probabilistic” SPT-Based Fig. 12: Recommended “Probabilistic” SPT-Based
Liquefaction Triggering Correlation (for Liquefaction Triggering Correlation (for
Mw=7.5 and σv′=1.0 atm) and the Relationship Mw=7.5 and σv′=1.0 atm) with Adjustments
for “clean sands” Proposed by Seed et al. for Fines Content Shown.
(1984)
Fig. 13: Recommended Kσ Values for σ’v > 2 atm. CSR* = CSReq,M=7.5,1atm = CSReq,M=7.5 / Kσ (Eq 14)
50
For “deterministic” evaluation of liquefaction resistance,
40
largely compatible with the intent of the earlier relationship
30 proposed by Seed et al. (1984), the same steps can be
20 undertaken (except for the fines adjustment) to asses the fully
10 adjusted and normalized CSReq,M=7.5,1atm values, and
0
normalized N1,60 values, and these can then be used in
200 600 1000 1400 1800 2200 2600 3000 3400 3800 4200
conjunction with the recommended “deterministic”
relationship presented in Figure 14. The recommendations of
1.4 Figure 14 correspond to the new probabilistic relationships
1.3
(for PL = 20%), except at very high CSR (CSR > 0.4). At
these very high CSR; (a) there is virtually no conclusive field
1.2
data, and (b) the very dense soils (N1,60 > 30 blows/ft) of the
1.1
boundary region are strongly dilatant and have only very
1 limited post-liquefaction strain potential. Behavior in this
Kσ 0.9 region is thus not conducive to large liquefaction-related
0.8
displacements, and the heavy dashed lines shown in the upper
portion of Figure 12 represent the authors’ recommendations
0.7
____ This Study in this region based on data available at this time.
0.6
------ Recommended by NCEER
0.5 Working Group (1998) This section of this paper has presented the development of
0.4 recommended new probabilistic and “deterministic”
0 400 800 1200 1600 2000 2400 2800 3200 3600 4000 relationships for assessment of likelihood of initiation of
σv' (psf) liquefaction. Stochastic models for assessment of seismic soil
liquefaction initiation risk have been developed within a
Bayesian framework. In the course of developing the
Fig. 14: Values of Kσ Developed and Used in These proposed stochastic models, the relevant uncertainties
Studies, NCEER Working Group including: (a) measurement/estimation errors, (b) model
Recommendations (for n=0.7, DR ≈ 60%) for imperfection, (c) statistical uncertainty, and (d) those arising
Comparison from inherent variables were addressed.
Finally, a common question is “what happens at N1,60,cs values Estimates of the “large” deformations likely to occur for these
greater than about 15 blows/ft.?” The answer is that the types of cases can often be made with fair accuracy (within a
relationships of Figures 18 and 19 should be concave upwards factor of about + 2). “Large” liquefaction-induced
(to the right), so that extrapolation at constant slope to the displacements/deformations (> 1m.) are principally the result
right of N1,60,cs=15 blows/ft should provide a conservative of gravity-induced “slumping”, as geometric rearrangement of
basis for assessment of Su,r in this range. As these projected the driving soil and/or structural masses is required to re-
values represent relatively good strength behavior, this linear establish static equilibrium. A majority of the deformations,
extrapolation tends to be sufficient for most projects. It should for these cases, occur after strong shaking has ceased so that
be noted, however, that values of Su,r should generally not be cyclic inertial forces are not very important in “driving” the
taken as higher than the maximum drained shear strength. deformations (though they are very important in “triggering”
Values of Su,r higher than the fully-drained shear strength the liquefaction-induced ground softening.)
Figure 22 shows the results of this approach (both equations, Fig. 22: Predicted vs. Measured Displacements from
as applicable.) Figure 22(a) shows a plot of predicted Lateral Spreading Case Histories (after
displacement magnitude vs. the actual observed displacement Bartlett and Youd, 1995)
for the case histories studied. For (measured) displacements
greater then approximately 1.5m., the ratio of
predicted:measured displacements was generally in the range this approach is the progressive acceleration and then
of 0.5:1 to 2:1, and this is a reasonable band of accuracy for deceleration of the displacing soil (and/or structural) mass.
engineering purposes in this range of displacements. The deformations are not arrested when the geometry is
sufficiently rearranged as to produce a “static” Factor of
The third method for estimation of expected “large” Safety of 1.0 (based on post-liquefaction strengths, as
liquefaction-induced displacements is based on evaluation of appropriate.). Instead, shear strength must be employed to
the deformations/displacements required to re-establish static overcome the momentum progressively accumulated during
equilibrium. This requires careful assessment of the most acceleration of the displacing mass, so that the deforming
critical mode of failure/deformation. An important issue in mass comes to rest at a “static” Factor of Safety of greater
Figures 23 through 25 illustrate the complicated types of Further complicating the issue of prediction of liquefaction-
mechanical behaviors that control cyclic deformations in this induced deformations is the fact that, for most cases of
“small to moderate” displacement range. Figure 23 presents engineering interest, there is a directionally preferential
the results of an undrained cyclic simple shear test of “driving” shear stress due to gravity loading (in addition to
Monterey #0/30 sand at a relative density of Dr = 50%, and an cyclic inertial stresses induced by the earthquake). Figure 25
initial vertical effective stress of σ’v,i = 85 kPa. These presents the results of an undrained cyclic simple shear test
conditions correspond roughly to a soil with an N1,60,cs value of with these initial "driving” shear stresses. In this test, the
about 10 blows/ft. In this figure, (a) the bottom left figure “driving” shear stresses are aligned in the same direction as
presents evolution of cyclically-induced pore pressures the (reversing) cyclic shear stress loading, and the initial
(expressed as reduction in σ’v,i), (b) the bottom right figure (constant) driving shear stresses are equal to 0.08 times the
shows increasing shear strains with increasing numbers of initial vertical effective stress (of 85 kPa).
cycles, (c) the top right figure shows shear stress vs. shear
0.1 0.1
0.0 0.0
-0.1 -0.1
-0.2 -0.2
-0.3 -0.3
35 35
30 30
Number of Cycles
25 25
20 20
15 15
10 10
5 5
0 0
1 0.8 0.6 0.4 0.2 0 -1 0 -5 0 5 10
Fig. 23: Undrained Cyclic Simple Shear Test on Monterey #30/0 Sand (Test No. Ms15j)
Dr=50%, σv,i’=85 kPa, CSR=0.22 , α=0
0.4 0.4
N orm. Shear Stress, ( τ / σ )
c
0.2 0.2
0.0 0.0
-0.2 -0.2
-0.4 -0.4
30 30
25 25
Number of Cycles
20 20
15 15
10 10
5 5
0 0
1 0.8 0.6 0.4 0.2 0 -1 0 -5 0 5 10
Fig. 24: Undrained Cyclic Simple Shear Test on Monterey #30/0 Sand (Test No. Ms30j)
Dr=75%, σv,i’=85 kPa, CSR=0.4 , α=0
0.2 0. 2
0.0 0. 0
- 0.2 -0.2
40 40
35 35
30 30
Number of Cycles
25 25
20 20
15 15
10 10
5 5
0 0
1 0. 8 0. 6 0. 4 0. 2 0 -2 0 2 4 6 8 10 12
Fig. 25: Undrained Cyclic Simple Shear Test on Monterey #30/0 Sand (Test No. Ms10k)
Dr=55%, σv,i’=85 kPa, CSR=0.33 , α=0.18
In addition to the types of cyclic softening and dilatent re- stresses) has begun to be available, and development and
stiffening shown in the two previous figures, this test (Figure calibration of improved analytical and constitutive models for
25) also exhibits cyclic “ratcheting” or progressive this type of behavior are currently still under development.
accumulation of shear strains in the direction of the driving
shear force. It is this type of complex “ratcheting” behavior Additional complications involved in attempting to predict
that usually principally controls “small to moderate” “small to moderate” liquefaction-induced deformations and
liquefaction-induced deformations and displacements displacements include: (1) the irregular and multi-directional
(displacements in the range of about 2 to 75 cm. for field loading involved in field situations, representing a complex
cases.) and multi-directional seismic response problem, and (2) the
many types and “modes” of deformations and displacements
This problem is further complicated in field cases by the that can occur.
occurrence of cyclic shear stresses “transverse” (not parallel
to) the direction of the (static) driving shear stresses. Figures 26 and 27 illustrate a number of “modes” or
Boulanger et al. (1995) clearly demonstrated that cyclic shear mechanisms that can result in “small to moderate” lateral and
stresses transverse to driving shear forces can, in many cases, vertical displacements, respectively. These figures are
represent a more severe type of loading for “triggering” of schematic and for illustrative purposes only; they are not to
liquefaction than cyclic shear stresses aligned “parallel” with scale.
driving forces. It is only in the last few years, however, that
high quality laboratory data with “transverse” as well as Figure 26 illustrates three examples of modes of deformation
“parallel” cyclic simple shear loading (and driving shear that can produce “small to moderate” liquefaction-induced
lateral displacements (of less than about 1m.) It should be strains, as a function of SPT N-values. As shown previously
noted that these can also produce much larger deformations, if in Figures 23 through 25, the shear strain required for
the liquefiable soils are very loose, and geometry is dilational re-stiffening decreases with increased initial
sufficiently adverse. density (or increased N-value). Although there is not yet a
well-established (or well-defined) basis for selection of the
Figure 26(a) shows an example of limited lateral spreading precise shear strain corresponding to the “limiting” shear
towards a free face, and Figure 26(b) shows an example of strain (see for examples, Figures 23 through 25), the values of
limited lateral spreading downslope or downgrade. These Figure 28 represent suitable approximate values for many
modes can also give rise to large displacements, but when the engineering purposes. The recommendations of Figure 28 are
liquefiable soils have limited shear strain potential (the shear for sands with approximately 10% silty fines. Shamoto et al.
strain required for dilatent re-stiffening), then displacements also presented similar figures for 0% and 20% fines, but the
are limited. differences are less than the uncertainty in defining precisely
what is meant by “limiting” shear strain.
Figure 28 (Shamoto, Zhang and Tokimatsu, 1998) presents
engineering estimates of limiting (post-liquefaction) shear
(c) Global Rotational or Translational Site (d) “Slumping” or Limited Shear Deformations
Displacement
(e) Lateral Spreading and Resultant Pull- (f) Localized Lateral Soil Movement
Apart Grabens
(g) Full Bearing Failure (h) Partial Bearing Failure (i) Foundation Settlements
or Limited “Punching” Due to Ground Softening
Exacerbated by Inertial
“Rocking”
There is an urgent need for improvement of our ability to Table 2 presents a brief list of selected major mitigation
accurately and reliably estimate expected “limited”, or “small methods available. It should be noted that these do not have to
to moderate” liquefaction-induced deformations and be employed singly; it is often optimal to use two or more
displacements. There is a similar need to improve our ability methods in combination.
to assess the expected ramifications of these types of
displacements and deformations on the performance of It is not reasonable, within the constraints of this paper, to
buildings and other engineered facilities. In addition, there are attempt a comprehensive discussion of all available mitigation
currently no well-established standards regarding expectations methods. Instead, limited comments will be offered regarding
of “acceptable” performance for most types of structures and various aspects of some of these. It should be noted that
facilities. These are all important areas in which further mitigation of liquefaction hazard is an area subject to
progress is urgently needed. considerable controversy, and that our understanding of the
efficacy of some of these methods is still evolving. It is
During the middle portion of the 20th Century, considerable suggested that key issues to be considered in selection and
research was done to develop an understanding of the implementation of mitigation methods are: (1) applicability,
consequences of various levels of differential (static) (2) effectiveness, (3) the ability to verify the reliability of the
settlements on different types of structures. This work, which mitigation achieved, (4)cost, and (5) other issues of potential
involved considerable field studies of performance of actual concern (e.g.: environmental and regulatory issues, etc.).
structures, also required close collaboration between
geotechnical and structural professionals. The results of these More comprehensive treatments of many of the mitigation
studies led to greatly improved understanding of the methods listed in Table 2 are available in a number of
ramifications of various levels of differential (static) refeences (e.g.: Mitchell, 1995; Hausmann, 1990).
The first class or catagory of methods listed in Table 2 involve The second group of methods listed in Table 2 involve in-situ
surface compaction. When this is the case, potentially ground densification. It is recommended that these methods
liquefiable soil types should be placed in layers and be coupled with a suitably comprehensive post-treatment
compacted, using vibratory compaction, to specifications verification program to assure that suitable mitigation has
requiring not less than 95% relative compaction based on the been achieved. CPT testing is particularly useful here, as it is
maximum dry density (γd,max) as determined by a Modified rapid and continuous. When CPT is to be used for post-
AASHTO Compaction Test (ASTM 1557D). densification verification, it is a very good idea to establish
pre-densification CPT data, and to develop site-specific cross-
correlation between SPT and CPT data.
Vibrodensification is generally effective in soils with less than Compaction piles provide improvement by three mechanisms;
about 5% clay fines, but can be ineffective in soils with larger (1) by densification due to driving installation, (2) by
fractions of clay fines. It had long been thought that the increasing lateral stresses, and (3) by providing structural
diffficulty in vibrodensification of soils with high fines reinforcing elements. This method is only rarely used,
contents was related to the inability of water to escape, and however, due to its cost. It is generally employed in unusual
indeed some improvement in densification of soils with high situations where other methods cannot reliably be
fines contents has been observed with the use of wick drains to implemented.
assist in allowing egress of water. It is noted, however, that
the clay contents at which vibrodensification begins to be Blasting can be used to achieve deep densification of
ineffective are very similar to the clay contents at which potentially liquefiable soils. This method, however, tends to
classic cyclically-induced liquefaction ceases to occur (see produce less uniform densification than vibrodensification,
Figure 2). It appears likely that, as vibrodensification and generally cannot reliably produce densities as high as
essentially works by liquefying and densifying the soils, the those that can be obtained with high energy vibrodensification
limit of “treatable” soil types is largely coincident with the methods that effectively transmit high vibrational energy to
types of soils that are “liquefiable”, and thus in need of soils at depth (e.g. Vibroflotation, etc.) Blasting also raises
treatment. environmental concerns, issues regarding propagation of
vibrations across neighboring sites, and issues regarding noise
Some of the vibrodensification methods also result in and safety.
installation of dense gravel columns through the treated
ground. It has been suggested that these dense gravel Compaction grouting is the last of the “in-situ ground
columns, which have high shear moduli relative to the densification” methods listed in Table 2, and also the first of
surrounding (treated) soils, will attract a large share of the three “grouting” methods listed in Table 2. Compaction
shear stresses propagating through the composite treated grouting involves injection of very stiff (low slump) cement
ground, and thus partially shield the softer surrounding soils. grout into the ground at very high pressure, ideally forming
This, in turn, would produce the added benefit of reducing he “bulbs” of grout and displacing the surrounding soils.
cyclic shear stress ratios (CSR) to which the treated soils Compaction grouting works both by densifying soils, and by
would be subjected during an earthquake. increasing in-situ effective lateral stresses. The degree of
densification that can be achieved by the monotonic (non-
Estimates of the level of shear stresses borne by the dense cyclic) loading imposed by the growing grout mass is
gravel columns are sometimes computed by estimating the dilationally limited, however, and recent research suggests that
contributions of the stiffer columns and the softer surrounding the increased lateral stresses can relax over time. An
soil, based an assumption of a simple shear mode of additional drawback is the difficulty in verifying improvement
deformation, and using contributory areas of the gravel by means of penetration testing. Compaction grouting
columns and the surroundiing soils and their respective shear performed well at one site in San Francisco during the 1989
moduli. Unfortunately, for column height to diameter ratios of
Notes:
1. H = total (vertical) depth of soils of the type or types referred to.
2. Vs = seismic shear wave velocity (ft/s) at small shear strains (shear strain ~ 10-4%).
3. If surface soils are cohesionless, Vs may be less than 800 ft/s in top 10 feet.
4. “Cohesionless soils” = soils with less than 30% “fines” by dry weight. “Cohesive soils” = soils with more than 30% “fines” by
dry weight, and 15% ≤ PI (fines) ≤ 90%. Soils with more than 30% fines, and PI (fines) < 15% are considered “silty” soils herein,
and these should be (conservatively) treated as “cohesive” soils for site classification purposes in this Table.
5. “Soft Clay” is defined as cohesive soil with: (a) Fines content ≥ 30%, (b) PI(fines) ≥ 20%, and (c) Vs ≤ 600 ft/s.
6. Site-specific geotechnical investigations and dynamic site response analyses are strongly recommended for these conditions.
Response characteristics within this Class (E) of sites tends to be more highly variable than for Classes A0 through D, and the
response projections herein should be applied conservatively in the absence of (strongly recommended) site-specific studies.
7. Site-specific geotechnical investigations and dynamic site response analyses are required for these conditions. Potentially
significant ground failure must be mitigated, and/or it must be demonstrated that the proposed structure/facility can be engineered
to satisfactorily withstand such ground failure.
Fig. 31: Proposed Site Dependent Response Spectra Fig. 33: Calculated Normalized Response Spectra for
with 5% Damping (Modified After Seed et Oakland and Los Angeles Deep Stiff Sites
al., 1997) Compared to Current Design Spectra (Chang et
al., 1997)
Andrus, R.D. and Stokoe, K.H., (2000) “Liquefaction Finn, W. D. Liam. (1998) “Seismic Safety of Embankment
Resistance of Soils from Shear-Wave Velocity.” Journal of Dams Development in Research and Practice 1988-1998.”
Geotechnical Special Publication No. 75, Proceedings of a
Hamada, M., O'Rourke, T. D., and Yoshida, N. (1994). Jong, H.-L. and Seed, R. B. (1988). "A Critical Investigation
"Liquefaction-Induced Large Ground Displacement." of Factors Affecting Seismic Pore Pressure Generation and
Performance of Ground and Soil Structures during Post-Liquefaction Flow Behavior of Saturated Soils",
Earthquakes: Thirteenth International Conference on Soil Geotechnical Engineering Research Report No. SU/GT/88-01,
Mechanics and Foundation Engineering, New Delhi, Stanford University.
Proceedings, Japanese Society of Soil Mechanics and
Foundation Engineering Pub., Tokyo, p. 93-108. Kishida, H. (1966), "Damage to Reinforced Concrete
Buildings in Niigata City with Special Reference to
Hamada, M., Yasuda, S., Isoyama, R., and Emoto, K. (1986). Foundation Engineering", Soils and Foundations, Vol. VII,
"Study On Liquefaction Induced Permanent Ground No. 1.
Displacements", Report of the Association for the
Development of Earthquake Prediction in Japan, Tokyo Japan, Koizumi, Y (1966), "Change in Density of Sand Subsoil
Available from the Faculty of Marine Science and caused by the Niigata Earthquake", Soils and Foundations,
Technology, Tokai University. Vol. VIII, No. 2, pp. 38-44.
Harder, L. F. (1977). "Liquefaction of sand under irregular Liao, S. S. C., Lum, K. Y. (1998), "Statistical Analysis and
loading conditions." M.S. Thesis, University of California, Application of the Magnitude Scaling Factor in Liquefaction
Davis. Analysis", Geotechnical Earthquake Engineering and Soil
Dynamics III, Vol. 1, 410-421.
Harder, L. F. (1988). "Use of Penetration Tests to Determine
the Cyclic Loading Resistance of Gravelly Soils During Liao, S. S. C., Veneziano, D., Whitman, R.V. (1988),
Earthquake Shaking." Ph.D. Thesis, University of California, "Regression Models for Evaluating Liquefaction Probability",
Berkeley. Journal of Geotechnical Engineering, ASCE, Vol. 114, No. 4,
pp. 389-409.
Harder, L. F. Jr. (1997) “Application of the Becker
Penetration Test for Evaluating the Liquefaction Potential of Liao, S. S. C., Whitman, R. V. (1986), "Overburden
Gravelly Soils.” Proc., NCEER Workshop on Evaluation of Correction Factor for SPT in Sand", Journal of Geotechnical
Liquefaction Resistance of Soils, NCEER-97-0022. Engineering, ASCE, Vol. 112, No. 3, March 1986, pp. 373-
377.
Hausmann, M. R. (1990). Engineering principals of ground
modification, McGraw Hill, pp. 632. Mitchell, J. K., Baxter, C. D. P., and Munson, T. C. (1995).
"Performance of improved ground during earthquakes." Soil
Hynes, M. E. (1988). "Pore pressure generation characteristics Improvement for Earthquake Hazard Mitigation, ASCE, New
of gravel under undrained loading." Ph.D. Thesis, University York, 1-36.
of California, Berkeley.
NCEER (1997), "Proceedings of the NCEER Workshop on
Idriss, I. M. (2000), "Personal Communication" Evaluation of Liquefaction Resistance of Soils", Edited by
Idriss, I. M., Sun, J. I., (1992), "Users Manual for SHAKE91, Youd, T. L., Idriss, I. M., Technical Report No. NCEER-97-
A Computer Program for Conducting Equivalent Linear 0022, December 31, 1997.
Seismic Response Analyses of Horizontally Layered Soil
Poulos, S. J., Castro, G., and France, J. W. (1985). Seed, R. B., Cetin, K. O., Der Kiureghian, A., Tokimatsu, K.,
"Liquefaction Evaluation Procedure." Journal of Geotechnical Harder, L. F. Jr., and Kayen, R. E. (2001) “SPT-Based
Engineering, ASCE, 111(6), 772-792. Probabilistic and Deterministic Assessment of Seismic Soil
Liquefaction Potential.” Submitted to the ASCE Journal of
Riemer, M. F. (1992). "The effects of testing conditions on the Geotechnical and Geoenvironmental Engineering.
constitutive behavior of loose, saturated sands under
monotonic loading." Ph.D. Thesis, University of California, Shamoto, Y., Zhang, J.-M., and Tokimatsu, K. (1998).
Berkeley. "Methods for evaluating residual post-liquefaction ground
settlement and horizontal displacement." Soils and
Riemer, M. F., Seed, R. B., and Sadek, S. (1993). “The Foundations, Special Issue, 69-83.
SRS/RFT Soil Evaluation Testing Program.” University of
California, Berkeley Geotechnical Report No. UCB/GT-93/01. Stark, T. D. and Mesri, G. (1992). "Undrained Shear Strength
of Liquefied Sands For Stability Analysis." Journal of
Riemer, M.F., Seed, R.B. (1997), “Factors Affecting Apparent Geotechnical Engineering, ASCE, 118(11), 1727-1747.
Position of Steady-State Line”, Journal of Geotechnical and
Geoenvironmental Engineering, Vol 123, No. 3, March 1997, Toprak, S., Holzer, T. L., Bennett, M. J., Tinsley, J. C. (1999),
pp. 281. "CPT- and SPT-based Probabilistic Assessment of
Liquefaction Potential, Proceedings of Seventh U.S.-Japan
Robertson, P. K. and Wride, C. E. (1998). "Evaluating Cyclic Workshop on Earthquake Resistant Design of Lifeline
Liquefaction Potential Using The Cone Penetration Test." Facilities and Countermeasures Against Liquefaction.
Canadian Geotechnical Journal, 35(3), 442-459.
Vaid, Y.P., Chung, E.K.F., and Kuerbis, R.H. (1990). “Stress
path and steady state.” Canadian Geotechnical Journal, 27, 1-
Seed, H. B., Idriss, I. M. (1971), “Simplified Procedure for 7.
Evaluating Soil Liquefaction Potential”, Journal of the Soil
Mechanics and Foundations Division, ASCE, Vol. 97, No Von Thun, J. L. (1986). "Analysis of Dynamic Compaction
SM9, Proc. Paper 8371, September 1971, pp. 1249-1273. Foundation Treatment Requirements, Stage I, Jackson Lake
Dam", Technical Memo No. TM-JL-230-26, Bureau of
Seed, H. B., Seed, R. B., Harder, L. F., and Jong, H. L. (1989). Reclamation, Engineering and Research Center, Embankment
"Re-evaluation of the Slide in the Lower San Fernando dam in Dams Branch.
the 1971 San Fernando earthquake", Earthquake Engineering
Research Center Report No. UCB/EERC-88/04, University of Yoshimi, Y., Tokimatsu, K., Ohara, J. (1994), "In-situ
California, Berkeley. Liquefaction Resistance of Clean Sands Over a Wide Density
Range", Geotechnique, Vol. 44, No. 3, pp. 479-494.
Seed, H. B., Tokimatsu, K., Harder, L. F., and Chung, R. M.
(1985). "Influence of SPT Procedures in soil liquefaction Youd, T. L., Idriss, I. M., Andrus, R. D., Arango, I., Castro,
resistance evaluations." Journal of Geotechnical Engineering, G., Christian, J. T., Dobry, R., Finn, W. D. L., Harder, L. F.
ASCE, 111(12), 1425-1445. Jr., Hynes, M. E., Ishihara, K., Koester, J. P., Liao, S. S. C.,
Marcuson, W. F. III., Martin, G, R., Mitchell, J. K., Moriwaki,
Seed, H. B., Tokimatsu, K., Harder, L. F., Chung, R. M. Y., Power, M. S., Robertson, P. K., Seed, R. B., and Stokoe,
(1984), "The Influence of SPT Procedures in Soil Liquefaction K. H., II. (1997) Summary Paper, Proc., NCEER Workshop
Resistance Evaluations", Earthquake Engineering Research on Evaluation of Liquefaction Resistance of Soils, NCEER-
Center Report No. UCB/EERC-84/15, University of 97-0022.
California at Berkeley, October, 1984.
Youd, T. L. (2000) Personal Communication.
Seed, R. B. and Harder, L. F. (1990). "SPT-Based Analysis Of
Cyclic Pore Pressure Generation And Undrained Residual Youd et al. (2001) Personal Communication.
Strength." H.B. Seed Memorial Symposium, Berkeley, Ca.,
BiTech Publishing, Ltd., v. 2, p. 351-376. Youd, T. L., Noble, S. K. (1997), "Liquefaction Criteria Based
on Statistical and Probabilistic Analyses", Proceedings of the
Seed, R. B., Chang, S. W., Dickenson, S. E., and Bray, J. D. NCEER Workshop on Evaluation of Liquefaction Resistance
(1997) “Site-Dependent Seismic Response Including Recent of Soils, December 31, 1997, pp. 201-205.
Strong Motion Data.” Proc., Special Session on Earthquake