Development of Residual Forces in Driven Piles - Discussion

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DISCUSSIONS AND CLOSURES
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Discussion of “Settlement Ratio of Pile
Groups in Sandy Soils from Field Load
Tests” by Y. Xu and L. M. Zhang
August 2007, Vol. 133, No. 8, pp. 1048–1054.
DOI: 10.1061/�ASCE�1090-0241�2007�133:8�1048�
William J. Neely, P.E., M.ASCE1
1Vice President, The Reinforced Earth Company, 1660 Hotel Circle N.,
Suite 304, San Diego, CA 92108. E-mail: wneely@reinforcedearth.
com
The authors analyzed data from loading tests on more than 30 pile
groups and found that the mean value of settlement ratio de-
creases as the average load per pile in the group increases. Their
results indicate that the settlement ratio Rs, which is equal to the
settlement of a pile group divided by the settlement of a single
pile at the same average load per pile, is about 2 for a factor of
safety of 2.5. Also, the settlement ratio appears to be independent
of the length, diameter, spacing, and number of piles in the group,
although more than half of the authors’ cases consist of nine-pile
groups.
It is surprising that the data collected from the literature by the
authors do not include the comprehensive series of loading tests
on pile groups, which comprise as many as 25 piles in a single
group at spacings of between three- and six-pile diameters, re-
ported by Berezantsev et al. �1961�. The results from this series of
tests are used here to demonstrate that the authors’ findings re-
garding both the magnitude of Rs, the group settlement ratio, and
its roughly linear reduction with increasing normalized load
Q /Qf, may produce unconservative assessments of pile group
settlement in some situations. The results from Berezantsev et al.
�1961� are also used to illustrate good agreement between mea-
sured Rs values and those predicted from the method of Poulos
and Davis �1980�.
About half �15� of the case histories in Table 2 involve loading
tests on pile groups in which the pile length was 2 m or less; for
the remaining 16 groups, the average pile length is 4.6 m, in the
range of 2.5 to 7.64 m. For almost 60% of the pile groups exam-
ined by the authors, the failure loads for the corresponding single
piles are less than 100 kN, and in only three cases does the single
pile failure load exceed 200 kN. The stiffness factor K �defined
later� for the 14 groups of aluminum pipe piles is about an order
of magnitude less than for the groups of solid reinforced concrete
piles and steel pipe piles. In foundation engineering practice, piles
are usually much longer, are considerably stiffer, and have very
much greater failure loads than is the case for many of the piles in
the study. It is probably for these reasons that the authors did not
advocate the application of their findings to pile design practice.
The work reported by Berezantsev et al. �1961� included load-
ing tests on 267-mm diameter concrete piles driven 5.5 m into
medium dense silty sand. Piles were load-tested as single piles
and in groups of 4, 9, 16, and 25 piles at spacings ranging from
three- to six-pile diameters. Fig. 1 shows the load-settlement be-
havior of a single pile as well as the relationship between the
average load per pile and the corresponding settlement for groups
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 J. Geotech. Geoenviron. Eng., 2
of 4, 9, 16, and 25 piles at a spacing/diameter �s/d� ratio of 3. To
be consistent with the approach used by the authors, the single-
pile failure load Qf was determined using Davisson’s �1972� fail-
ure criterion. The value of Qf is 314 kN and occurs at a pile head
settlement of 5.25 mm. The most striking feature of the data in
Fig. 1 is that for the same average load per pile, the settlement of
a group increases markedly as the number of piles in the group
increases. For each of the pile group load-settlement data points
in Fig. 1, the group settlement ratio Rs was obtained by dividing
the settlement of the pile group by the settlement of the single pile
at the same load. The values of group settlement ratio are plotted
with the normalized load Q /Qf in Fig. 2 for each of the four pile
groups with a spacing/diameter ratio of 3. It can be seen that
individual values of Rs range from about 3 to more than 12,
depending on the value of Q /Qf and the number of piles in the
group. Linear regression was used to find the best straight line fit
to the data for each pile group. The results presented in Fig. 2 also
show that Rs increases linearly with increasing Q /Qf and that the
rate of increase in Rs becomes more pronounced as the number of
piles in the group increases. The linear relationships between Rs
and Q /Qf in Fig. 2 also suggest that elastic methods for evaluat-
ing the settlement of pile groups, such as that of Poulos and Davis
�1980�, may be suitable for values of Q /Qf up to 0.5, correspond-
ing to a factor of safety of 2. The relationships between settlement
ratio and normalized load in Fig. 2 contrast sharply with that from
the authors’ study shown in Fig. 5 of their paper. Values of Rs
from the authors’ analyses are less than any of the values in
Fig. 2; also, their trend of decreasing Rs with increasing
Q /Qf�reverse of that derived from the tests at a spacing diameter
ratio of 3 reported by Berezantsev et al. �1961�.
Fig. 3 compares the load-settlement response of a single pile
with that for three nine-pile groups at spacing/diameter ratios of
3, 4.5, and 6. Fig. 4 shows the corresponding values of settlement
ratio plotted with normalized load. At a spacing/diameter ratio of
3, the group settlement ratio Rs increases with increasing Q /Qf,
but at spacings of 4.5- and 6-pile diameters, the trend is reversed,
and Rs decreases as Q /Qf increases. The reversal in Rs−Q /Qf
relationship with increasing pile spacing shown in Fig. 4 suggests
Fig. 1. Load settlement curves for single pile and groups of piles
with s /d=3
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that pile group settlement ratio depends not only on the average
load per pile at which Rs is determined but also on the geometry
of the pile group.
Theoretical values of settlement ratio are given by Poulos and
Davis �1980� for a range of pile group sizes, spacing/diameter
ratios, pile length/diameter ratios and pile stiffness factor. The
pile stiffness factor K is given by
K =
EPRA
ES
�1�
where EP�elastic modulus of pile material; ES�elastic modulus
of the soil; and RA�area ratio which is equal to 1 for a solid pile.
Fig. 2. Relationships between Rs and normalized load for groups of
piles with s /d=3
Fig. 3. Load settlement curves for nine-pile groups with various s/d
ratios
Fig. 4. Relationships between Rs and normalized load for nine-pile
groups with various s/d ratios
1418 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGIN
 J. Geotech. Geoenviron. Eng., 2
Poulos and Davis �1980� quote average values of K for solid
concrete piles of 1,500 if installed in loose sand, and 500 if in-
stalled in dense sand. A value of K of 1,000 is used here for
analysis of the Berezantsev et al. �1961� tests. The length/
diameter �L /d� ratio for the piles is 20.5.
The method of Poulos and Davis �1980� was used to compute
theoretical values of settlement ratio for groups comprising up to
25 piles at a spacing/diameter �s/d� ratio of 3. These Rs values are
indicated by the solid line in Fig. 5. The linear Rs−Q /Qf relation-
ships presented in Fig. 2 have been used to determine the values
of Rs for several values of normalized load Q /Qf for each of the
four pile groups. The corresponding Rs values are indicated in
Fig. 5 for Q /Qf values of 0.2, 0.33, and 0.4, corresponding to
factorsof safety of 5, 3, and 2.5 respectively. It is encouraging to
note that the trend of the theoretical values is consistent with
those derived from the field loading tests. Theoretical and mea-
sured Rs values diverge slightly as Q /Qf increases. This diver-
gence may be because at higher normalized loads corresponding
to lower factors of safety, pile response is increasingly nonlinear,
and the influence of point resistance is more pronounced than at
low normalized loads or high factors of safety. With more field
data, it may be possible to empirically calibrate the theoretical
results to account for pile type, method of installation, and other
factors.
The linear relationships between settlement ratio Rs and nor-
malized load Q /Qf for nine-pile groups at three different spacings
presented in Fig. 4 have been used to determine settlement ratios
at Q /Qf values of 0.2, 0.33, and 0.5, corresponding to single-pile
Fig. 6. Comparison of measured and predicted Rs values for nine-
pile groups with various s/d ratios
Fig. 5. Comparison of measured and predicted Rs values for pile
groups with s /d=3
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factors of safety of 5, 3, and 2. These values are plotted in Fig. 6
along with the theoretical values from Poulos and Davis �1980�.
Again, the theory is successful in predicting the trend of decreas-
ing settlement ratio with increasing pile spacing derived from the
field loading tests. There is also good agreement between theoret-
ical and measured Rs values for the nine-pile groups over a wide
range of pile spacings and single-pile factors of safety.
Although more full-scale field loading tests on pile groups,
such as those reported by Berezantsev et al. �1961�, are needed to
examine the effects of soil densification caused by pile driving
and the sharing of load between shaft friction and point resis-
tances and other factors on pile group settlement behavior, it is
encouraging to find that the method of Poulos and Davis �1980�
provides a reliable means for determining the influence of many
other important parameters, such as group size, pile spacing, pile
stiffness factor, and pile length, on pile group settlement ratio.
References
Berezantsev, V. G., Khrisoforov, V. S., and Golubkov, V. N. �1961�.
“Load bearing capacity and deformation of piled foundations.” Proc.
5th Int. Conf. Soil Mech. & Found. Engrg., Vol 2, Paris, 11–15.
Davisson, M. T. �1972�. High Capacity Piles. Proc. Soil Mech. Lecture
Series, Innovations in Foundation Construction, ASCE-Illinois Sec-
tion, Chicago, 81–112.
Poulos, H. G., and Davis, E. H. �1980�. Pile foundation analysis and
design, Wiley, New York.
Closure to “Settlement Ratio of Pile
Groups in Sandy Soils from Field Load
Tests” by Y. Xu and L. M. Zhang
August 2007, Vol. 133, No. 8, pp. 1048–1054.
DOI: 10.1061/�ASCE�1090-0241�2007�133:8�1048�
Y. Xu1 and L. M. Zhang, M.ASCE2
1Res. Asst., Dept. of Civ. Engrg., The Hong Kong Univ. of Sci. and
Technol., Clear Water Bay, Hong Kong. E-mail: xuyao@ust.hk
2Assoc. Prof., Dept. of Civ. Engrg., The Hong Kong Univ. of Sci. and
Technol., Clear Water Bay, Hong Kong. E-mail: cezhangl@ust.hk
The authors are grateful to Dr. William J. Neely for his discussion
on the note. We do agree with the discusser that a theoretical
method can be used to calculate values of settlement ratio by
better considering the effect of group size, spacing/diameter ratio,
pile length/diameter ratio, and pile stiffness factor. Under working
load conditions, pile group interactions depend on two sets of
dimensionless parameters: those related to the soil and pile char-
acteristics, and those related to the geometry of the piles and the
pile group �Poulos 1989�. The method of Poulos and Davis �1980�
provides a good prediction of settlement ratio for groups of up to
25 piles considering the influence of the above factors. However,
conventional theoretical analyses, in which the modulus of the
soil between the piles is assumed to be the same as that adjacent
to the piles, may considerably overestimate group settlement
ratio, particularly in sands, even though such inaccuracy tends to
be on the conservative side �Poulos 1989, 1994�. In that sense, the
statistical results in Fig. 5 of the note provide a general guide on
possible values of settlement ratio for three-diameter-spaced nine-
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pile groups in sand, which has been indicated in the note.
The authors were aware of the load tests in silty fine sand
reported by Berezantsev et al. �1961�. The test results were not
used in the note because detailed test conditions were not re-
ported, and the test results could not be verified. The load-
settlement curves of a single pile and corresponding pile groups
in Figs. 1 and 3 in the discussion are transcribed from the load
test results shown in Fig. 1 as originally reported by Berezantsev
et al. �1961�. From Fig. 1, the field tests appeared to be conducted
on pile groups with cap-ground contact. The bearing behaviors of
the six-diameter-spaced 2�2 group, the 4.5-diameter-spaced
five-pile group, and the three-diameter-spaced 3�3 pile group
were treated the same probably because they had the same “load
transmitting area” �i.e., the same A value in Fig. 1�. So were the
behaviors of the three-diameter-spaced 4�4 pile group and the
4.5-diameter-spaced 3�3 pile group, and the behaviors of the
three-diameter-spaced 5�5 pile group and six-diameter-spaced
3�3 pile group. The authors could not verify such treatment
because no detailed information on soil variability, pile installa-
tion, load test arrangements and procedure, and load-settlement
curves for the single pile and the pile groups was reported in
Berezantsev et al. �1961�. The effects of pile installation on char-
acteristics of the soil within a pile group can be very important,
particularly for piles in sand �Poulos 1994�.
Based on the transcribed load-settlement curves of the single
pile and the three-diameter-spaced 2�2, 3�3, 4�4, and 5�5
pile groups shown in Fig. 1 of the discussion, the bearing capacity
group efficiency factors of these pile groups, defined at the settle-
ment at the single pile failure, range from 0.3 to 0.5. This result is
contradictory to the common understanding that the group effi-
ciency factor of a driven pile group with cap-ground contact in
sandy soils is usually larger than unity due to soil densification
and additional contribution from the cap-ground contact �Vesic
1969; Poulos and Davis 1980; O’Neill 1983; Zhang et al. 2001�.
For instance, the average group efficiency factor of three-
Fig. 1. Results of field tests on piled foundations with piles driven to
a depth of 5.6 m in silty fine sand. A�load transmitting area from the
piled foundation to the soil; ��load transmitting angle; B�pile di-
ameter; and p�total load/A �after Berezantsev et al. 1961�.
diameter-spaced pile groups with cap-ground contact in sandy
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soils are 1.40 based on results of 17 field and model load tests
�Zhang et al. 2001�. In addition, the settlement ratio values from
this series of tests �Fig. 2 in the discussion� are in the range of 3
to 10, which are rather high. O’Neill �1983� summarized pub-
lished data of settlement ratios of pile groups with cap-ground
contact. Most settlement ratio values were in the range of 1 to 4
for pile groups with a group width/pile diameter ratio of less than
10, which is consistent with the range of the settlement ratios for
the three-diameter-spacedpile groups at the working load level
�i.e., Q /Qf=0.5 in Fig. 4 in the note.�
References
Berezantsev, V. G., Khrisoforov, V. S., and Golubkov, V. N. �1961�.
“Load bearing capacity and deformation of piled foundations.” Proc.
5th Int. Conf. Soil Mech. and Found. Eng., Vol. 2, Paris, 11–15.
O’Neill, M. W. �1983�. “Group action in offshore piles.” Geotechnical
Practice in Offshore Engineering, S. G. Wright, ed., ASCE, New
York, 25–64.
Poulos, H. G. �1989�. “Pile behavior: Theory and application.” Geotech-
nique, 39�3�, 365–415.
Poulos, H. G. �1994�. “Settlement prediction for driven piles and pile
groups.” Geotech. Special Publication No. 40, A. T. Yeung and G. Y.
Félio, eds., Vol. 2, ASCE, New York, 1629–1649.
Poulos, H. G., and Davis, E. H. �1980�. Pile foundation analysis and
design, Wiley, New York.
Vesic, A. S. �1969�. “Experiments with instrumented pile groups in sand.”
Performance of deep foundations, ASTM STP 444, ASTM, West Con-
shohocken, Pa., 177–222.
Zhang, L. M., Tang, W. H., and Ng, C. W. W. �2001�. “Reliability of
axially loaded driven pile groups.” J. Geotech. Geoenviron. Eng.,
127�12�, 1051–1060.
Discussion of “Development of Residual
Forces in Long Driven Piles in Weathered
Soils” by L. M. Zhang and Hao Wang
October 2007, Vol. 133, No. 10, pp. 1216–1228.
DOI: 10.1061/�ASCE�1090-0241�2007�133:10�1216�
Bernadete Ragoni Danziger1 and
Francisco de Rezende Lopes2
1Associate Professor, Rio de Janeiro State Univ., Rio de Janeiro, RJ,
Brazil. E-mail: brdanzig@uerj.br
2Professor, COPPE, Federal Univ. of Rio de Janeiro, Rio de Janeiro, RJ,
Brazil. E-mail: flopes@coc.ufrj.br
The discussers found the authors’ paper of great value because it
is very seldom that a large-scale field-monitoring program includ-
ing the measurement of residual forces in long driven steel H
piles is performed. The information provided can be used to infer
previous conclusions supported by a paper related to prediction of
residual driving stresses in piles �Costa et al. 2001�. The authors
also found in the paper important residual stresses caused by re-
covery of the soil after the disturbance of the installation, but the
discussers will focus their attention only on toe residual loads
occurring just after driving.
In the paper by Costa et al. �2001�, a parametric analysis of
residual driving stresses was carried out by means of a driveabil-
ity program. This analysis presents a more rigorous approach to
residual stress prediction because it allows an important feature to
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be taken into account. Instead of using the common procedure of
most driveability programs, in which the final set is determined
indirectly, subtracting the quake from the maximum displacement
of the pile toe, in this analysis the set was determined directly, as
a final time was previously chosen and the pile toe displacement
was calculated up to this time. This maximum time period can be
adjusted, if necessary, until a stabilized pile toe displacement is
achieved. The procedure usually leads to a final stabilized toe
displacement higher than the usual set determined by most pro-
grams. This difference is attributed to residual stresses locked in
the pile toe during unloading. The residual stresses can be calcu-
lated from the stabilized stresses obtained in each pile element.
For the parametric study, a steel pile driven in a sandy soil was
chosen by Costa et al. �2001�. The pile was a 20 m-long double I
pile with a cross section of 154.6�10−4 m2. The toe resistance
percentage of the overall pile bearing capacity, among other fac-
tors, was analyzed. Table 1 from Costa et al. �2001� presents the
ratio of toe residual load to total bearing capacity for different
simulated cases. P2, P5, and P8 denote the toe resistance percent-
age to total capacity of 20, 50, and 80%, respectively. DL and DU
denote linear and uniform shaft friction resistance distribution.
Table 1 indicates that the ratio of toe residual load after driving
to total bearing capacity first increases and then decreases with
increasing end bearing capacity. In fact, residual shaft friction
along the pile arises as a means to equilibrate residual toe resis-
tance. When the percentage of toe resistance is low, the residual
toe resistance is obviously low and the ratio of residual load at the
pile toe �Qp,res driv� to total bearing capacity �QR� is small. On the
other hand, when the percentage of toe resistance is high, the
residual toe resistance is low because there is not enough residual
shaft friction to equilibrate it. When toe resistance approximates
shaft friction resistance, that is, when end bearing is close to 50%
of the total bearing capacity, the residual load can be high, as
residual shaft friction can be sufficient to restrain complete un-
loading of the pile toe.
Results from Table 1 also support some findings from the au-
thors that obtained increased residual forces with penetration
depth. In fact, for piles with short penetration depth the lateral
resistance is usually small and the toe resistance percentage to
total bearing capacity is usually high. From Table 1 one can ob-
serve that for high toe resistance percentage the toe residual load
after driving is relatively low because there is not enough lateral
resistance to prevent complete pile rebound. On the other hand,
for long piles with pile toe embedded in completely or moderately
decomposed granite with very high SPT blow count, as in the
case analyzed by the authors, lateral resistance is sufficiently high
to restrain the complete pile rebound. In the case of long floating
piles the toe resistance percentage to total bearing capacity is so
low that residual toe resistance will not be present at all. Even
though it is not common to have such a large-scale field-
monitoring program including the measurement of residual
forces, the authors’ tested piles are long driven steel H piles with
high toe resistance, with the boundary conditions favorable to the
development of high toe residual loads.
Figs. 5–8 from the authors’ paper show the development of
residual forces with depth shortly after driving each pile section.
From these curves the discussers can obtain the toe residual load
Table 1. Ratio of Toe Residual Load after Driving �Qp,res driv� to Total
Bearing Capacity �QR� �Adapted from Costa et al. �2001��
Case P2-DL P5-DL P8-DL P2-DU P5-DU P8-DU
Qp,res driv /QR 0.07 0.15 0.10 0.09 0.16 0.09
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 at each tested pile at the end of driving. Those values are repro-
duced in Table 2 at the column related to Qp res driv. The authors
reported that, at the end of driving, the residual forces were moni-
tored continuously until kentledges were installed for static load-
ing tests. The residual forces continued to develop over time in
the absence of fill or excavation activities. The authors found that
the distribution pattern of the residual forces along a pile does not
change over time but the toe residual forces increase continuously
over time. After the substantial increase in the toe residual force
after pile installation, the authors verified that the toe residual
force can consist of an even larger fraction of the ultimate toe
resistance. The authors explain that recompression may govern
the continued development of residual forces following pile
installation.
In Figs. 17, 18, and 19 of the authors’ paper the distribution of
the residual forces, apparent forces, and true forces for three dif-
ferent applied loads in the static loading tests are presented. The
authors considered the ultimate load as that obtained by the
Davisson failure criterion �Davisson 1972�. Based on the same
premise, the discussers extracted the toe resistance, thetotal ca-
pacity, and the toe residual loads from the aforementioned figures.
The discussers considered those values from the load transfer
curves for the higher load corresponding to the Davisson failure
criterion. The residual loads at pile toe at the end of driving for
the tested piles, as obtained from Figs. 5–8, are also presented in
Table 2. It should be emphasized that the approximate values in
Table 2 were obtained directly from each one of the aforemen-
tioned figures. In Table 2 the piles were listed according to in-
creasing toe bearing to total capacity percentage �Qp /QR�.
In Table 2 Qp denotes the toe resistance, QR the total capacity
�Davisson failure criterion�, and Qp,residual the residual load at pile
toe, observed at the static load test. Those are the “true values,”
as defined by the authors, referring to the force in the pile when
the residual force is included. Following the other columns of
Table 2, Qp res driv is the toe resistance developed at the end of
driving. The remaining columns express the ratios of the previous
ones.
Piles 1B1-1 and 1B2-2 were not included in Table 2 for the
following reasons: In pile 1B1-1 the authors reported that enor-
mous residual forces were developed when the pile was driven
into the bedrock. The true toe loads are significantly larger than
the applied loads over the entire applied load range. The static
loading test information of pile 1B2-2 is not complete and thus
was not included in the authors’ paper.
The general view observed by Costa et al. �2001� in the para-
metric study can also be inferred from the authors’ data included
in Table 2. It can also be verified that for all piles, except the first
two piles in Table 2, the ratio of toe residual load after driving to
Table 2. Summary of Tested Piles Information
Pile
Qp
�kN�
QR
�kN�
Qp /QR
�%�
Qp,residual
�kN�
Qp,res driv
�kN�
Qp,residual /
Qp res driv
Qp res driv /
QR
1A13 3940 11354 35 2062 1645 1.25 0.14
1B3-2 4313 10644 41 1125 1142 0.98 0.11
1B1-2 4125 9934 42 2812 656 4.28 0.07
1A22 5250 11708 45 4125 790 5.22 0.07
1B3-1 5625 11708 48 3375 1190 2.83 0.10
1B2-1 4125 8496 49 1500 903 1.66 0.11
1B1-4 5620 9580 59 380 444 0.85 0.05
1B3-3 6750 9580 70 750 681 1.10 0.07
1B1-3 8190 10999 74 380 217 1.75 0.02
total bearing capacity first increases and then decreases with in-
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 J. Geotech. Geoenviron. Eng., 2
creasing end bearing percentage, the higher ratio occurring for an
end bearing percentage close to 50%. One can then conclude that
residual loads after driving can be properly predicted by drive-
ability programs if the pile set is determined directly, as the usual
procedure adopted by the discussers. Table 2 also suggests that
for toe resistance percentage to total pile bearing capacity close to
50%, in addition to the greater ratio of toe residual load after
driving to total bearing capacity, the ratio of the long-term toe
residual load to the toe residual load after driving seems also to
assume the greater values. The discussers consider this aspect
worth analyzing in the future.
The authors have compared the tested piles’ behavior with the
SPT blow count, N, a number of times. However, no information
was given about the SPT system used and the corresponding en-
ergy delivered to the rod stem. In fact, Schmertmann and Palacios
�1979� have shown that the number of blows N varies inversely to
the energy delivered to the rod stem, at least for N values up to
50. Moreover, the International Society of Soil Mechanics and
Foundation Engineering �ISSMFE 1989� has recommended that
N values be adjusted by calibration to a reference energy of 60%
of the nominal energy of 474 J. Finally, it would be useful if
the authors could provide details about the procedure used to
measure such high N values, some greater than 200, including the
rate of blows.
References
Costa, L. M., Danziger, B. R., and Lopes, F. R. �2001�. “Prediction of
residual driving stresses in piles.” Can. Geotech. J., 38�2�, 410–421.
Davisson, M. T. �1972�. “High capacity piles.” Proc., Soil Mechanics
Lecture Series on Innovations in Foundation Construction, ASCE,
Illinois Section, Chicago, 81–112.
ISSMFE. �1989�. “Int. reference test procedure for the standard penetra-
tion test �SPT�.” Rep. of the ISSMFE—Technical Committee on Pen-
etration Testing of Soils—TC 16, 17–19.
Schmertmann, J. H., and Palacios, A. �1979�. “Energy dynamics of SPT.”
J. Geotech. Engrg. Div., 105�8�, 909–926.
Closure to “Development of Residual
Forces in Long Driven Piles in Weathered
Soils” by L. M. Zhang and Hao Wang
October 2007, Vol. 133, No. 10, pp. 1216–1228.
DOI: 10.1061/�ASCE�1090-0241�2007�133:10�1216�
L. M. Zhang1 and Hao Wang2
1Associate Professor, Dept. of Civil Engineering, The Hong Kong Univ.
of Science and Technology, Clear Water Bay, Hong Kong. E-mail:
cezhangl@ust.hk
2Postdoctoral Research Associate, Dept. of Civil Engineering, The Hong
Kong Univ. of Science and Technology, Clear Water Bay, Hong Kong.
The authors are grateful to Professor B. R. Danziger and Profes-
sor F. R. Lopes for their discussion of the paper and their inter-
pretation of the toe residual loads just after driving based on
outcomes from a residual stress analysis presented by Costa et al.
�2001�. Table 1 presented by Costa et al. �2001� indicates that the
ratio of toe residual load after driving to total bearing capacity
first increases and then decreases with increasing end bearing
capacity. This tendency is confirmed by the field test information
VIRONMENTAL ENGINEERING © ASCE / SEPTEMBER 2008 / 1421
008, 134(9): 1420-1421 
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reported in the paper, as summarized in Table 2 in the discussion.
As pointed out by the discussers, when the percentage of toe
resistance is low, the residual toe resistance is also low and the
ratio of residual load at the pile toe to total bearing capacity is
small. When the percentage of toe resistance is high, however, the
residual toe load is low because there is not enough residual shaft
friction to equilibrate it. The residual load can be high when the
toe resistance is close to 50% of the total bearing capacity be-
cause the residual shaft friction can be sufficient to restrain com-
plete unloading of the pile toe.
The gradual changes of residual forces in the piles might be
closely related to the setup of pile capacity over time �Shek et al.
2006�. In addition, as shown in the paper, large residual forces in
the piles affect the interpretation of load transfer along the piles
and, in turn, the settlement behavior of the piles �Zhang et al.
2008�. Residual forces in piles and setup in pile capacity rely not
only on pile and soil conditions but also on construction methods,
which introduce a source of uncertainty to the prediction of pile
capacity �Zhang et al. 2004�.
In the field study, the standard penetration test �SPT� blow
count values are generally greater than 200 blows/0.3 m at depths
greater than 50 m below the ground surface; this shows that the
weathered soils are rather strong. The SPT procedure in Hong
Kong has been well described by the Geotechnical Engineering
Office �1987, 1997� and conforms with the recommendations of
the International Society of Soil Mechanics and Foundation En-
gineering �ISSMFE 1989�. The blow count values reported in this
paper are not corrected for energy delivered to the rod stem and
embedment depth. A limited calibration study by the Geotechnical
Engineering Office �1997� showed that the percentage of energy
transferred to the rod stem ranged from 29 to 43% and that the
percentage did not vary much along depth up to 27 m. These
measured energy transfer ratio values are smaller than the refer-
ence energy ratio of 60%. When the SPT blow count approaches
200 blows/0.3 m, the energy used to penetrate soil is expectedto
1422 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGIN
 J. Geotech. Geoenviron. Eng., 2
be even smaller. Take the energy transfer during pile driving as an
example. The driving energy values from hydraulic hammers
were measured and studied during the pile driving at the same site
�Zhang 2005�. The records showed high energy transfer ratios to
the piles, 0.67-0.79 for DKH1523 hydraulic hammers and higher
ratios for Juttan20 hydraulic hammers. However, the effective
energy directed to cause permanent pile penetration was very
small �0.5-7.0%� toward final driving. The majority of energy
transferred to the pile was to cause elastic deformations of the
helmet-pile-soil system.
References
Costa, L. M., Danziger, B. R., and Lopes, F. R. �2001�. “Prediction of
residual driving stresses in piles.” Can. Geotech. J., 38�2�, 410–421.
Geotechnical Engineering Office. �1987�. Geoguide 2: Guide to site in-
vestigation, Geotechnical Engineering Office, Government of the
Hong Kong Special Administrative Region, Hong Kong.
Geotechnical Engineering Office. �1997�. Interim review of the standard
penetration test procedures with reference to Hong Kong practice,
Geotechnical Engineering Office, Government of the Hong Kong
Special Administrative Region, Hong Kong.
ISSMFE. �1989�. “Int. reference test procedure for the standard penetra-
tion test �SPT�.” Report of the ISSMFE-Technical Committee on Pen-
etration Testing of Soils-TC, 16, 17–19.
Shek, M. P., Zhang, L. M., and Pang, H. W. �2006�. “Setup effect in long
piles in weathered soils.” Geotech. Eng., 159�3�, 145–152.
Zhang, L. M. �2005�. “Pile driving process monitoring based on field
energy measurements.” Soils Found., 45�6�, 31–41.
Zhang, L. M., Tang, W. H., Zhang, L. L., and Zheng, J. G. �2004�.
“Reducing uncertainty of prediction from empirical correlations.”
J. Geotech. Geoenviron. Eng., 130�5�, 526–534.
Zhang, L. M., Xu, Y., and Tang, W. H. �2008�. “Calibration of models
for pile settlement analysis using 64 field load tests.” Can. Geotech.
J., 45�1�, 59–73.
EERING © ASCE / SEPTEMBER 2008
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