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JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGINEERING / SEPTEMBER 1999 / 787
oper settlement analysis for bridges on deep foundations
ld be performed. If, instead, one relies on single pile load
, which indicate that, at working loads, a pile will settle
a few millimeters, the impression is that piles do not
. However, these pile load tests generally last a few hours,
pencer J. Buchanan Prof., Dept. of Civ. Engrg., Texas A&M Univ.,
ge Station, TX 77843-3136. E-mail: briaud@tamu.edu
eotech. Engr., Kleinfelder Inc., 9555 Chesapeake Dr., Ste. 101, San
, CA 92123-6300.
te. Discussion open until February 1, 2000. To extend the closing
ne month, a written request must be filed with the ASCE Manager
urnals. The manuscript for this paper was submitted for review and
ble publication on July 10, 1996. This paper is part of the Journal
otechnical and Geoenvironmental Engineering, Vol. 125, No. 9,
mber, 1999. qASCE, ISSN 1090-0241/99/0009-0787–0796/$8.00
0 per page. Paper No. 13680.
The third layer is a mixed unit with an increasing amount of
clay seams and gravel layers; it is also of Pleistocene age and
was deposited by a stream of fluctuating energy. Below these
200,000-year-old sand layers and about 10 m below the
ground surface is the 45-million-year-old Eocene bedrock; this
bedrock is a dark gray clay shale that was deposited in a series
of marine transgressions and regressions. Erosion of the Eo-
cene marine clay took place before the Pleistocene river sed-
iments were deposited.
TEST SETUP AND LOAD SETTLEMENT CURVES
The plan view of the footings arrangement is presented in
Fig. 4. Five reinforced concrete square footings were con-
structed at the site, with the as-built dimensions listed in Table
1. Reaction shafts were belled as shown in Fig. 5 to resist in
BEHAVIOR OF FIVE LARGE
By Jean-Louis Briaud,1 Fellow, ASCE
ABSTRACT: Five square spread footings ranging in
settlement. They were all embedded 0.75 m into a
settlement curves are presented, as well as creep curve
the soil mass was instrumented with telltales and inc
mass were obtained as a function of depth and latera
measure footing settlement, how to present load test re
effect of cyclic loading and preloading on creep rate,
of the soil mass, and volume change observations. T
and the WAK (wave activated stiffness) test are evalua
Many results of these large-scale instrumented tests c
INTRODUCTION
This article describes a series of load tests performed on
five square spread footings ranging in size from 1 3 1 m to
3 3 3 m. The load-settlement curves have already been pre-
sented by Briaud and Jeanjean (1994) as an aid to the devel-
opment of a new load-settlement curve method for spread foot-
ings and by Briaud and Gibbens (1994) as the backbone of an
international prediction symposium. This article focuses on the
presentation and analysis of three new items related to these
tests: creep deformation as a function of time, vertical dis-
placement of the soil mass as a function of depth below the
center of the footings, and horizontal displacement of the soil
mass as a function of depth and lateral extent near the edge
of the footings. The soil data, footing test setup, and load set-
tlement curves are summarized to place the new results in
perspective.
SPREAD FOOTINGS OR DEEP FOUNDATIONS?
There are approximately 600,000 bridges in the United
States. If these bridges had to be replaced today, it would cost
about $300 billion. Each year some 6,000 new bridges are
built, and the Federal Highway Administration is conducting
research to minimize the cost of this infrastructure while op-
timizing safety and reliability. One such effort consists of help-
ing engineers place more bridges on spread footings by im-
proving the confidence in the design predictions. Indeed,
spread footings are generally less expensive than deep foun-
dations, with savings up to 20% of the cost of the bridge
(Briaud 1993).
Thinking that spread footings are more prone to settlement
than deep foundations would be a misconception, as shown in
two separate studies by Moulton et al. (1985) and Hearn
(1995); both engineers showed that, on the average, the set-
tlement of bridges on spread footings is very similar to that
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SPREAD FOOTINGS IN SAND
, and Robert Gibbens,2 Member, ASCE
size from 1 to 3 m were load tested up to 150 mm of
medium dense, fairly uniform, silty silica sand. Load-
relating settlement and time under a constant load. Since
inometers, vertical and horizontal movements in the soil
l extent. Conclusions are reached regarding how best to
ults, new correlations for use in design, creep settlement,
one of influence under the footing, mode of deformation
elve settlement methods, six bearing capacity methods,
ted by comparing the predictions with the measurements.
nfirm findings at small scale of previous researchers.
and the load is often increased when the settlement under the
previous load step has become less than 0.25 mm/h (ASTM
D1143). If such a rate is maintained for the design life of the
bridge, say 50 years, the settlement at 50 years would be 110
m. While it is unreasonable to assume that the settlement rate
will remain constant, the point is made that observations on
short-term tests and on single piles can be misleading for the
long-term behavior of pile groups.
The best foundation is the one that meets the design re-
quirements while minimizing cost and optimizing safety. Ex-
cept for obvious cases, spread footings should always be con-
sidered as a foundation alternative and eliminated only on the
basis of calculations.
SOIL
The spread footing tests were performed at the National
Geotechnical Experimentation Site on the Texas A&M Uni-
versity Riverside Campus near College Station, Tex. The soil
at the site is a medium dense, fairly uniform, silty fine silica
sand with the following average properties near the footings
and within the top 5 meters: mean grain size D50 = 0.2 mm,
SPT (standard penetration test) blow count 18 blows per 0.3
m, CPT (cone penetrometer test) point resistance 6 MPa, PMT
(pressuremeter test) limit pressure 800 kPa, PMT modulus 8.5
MPa, DMT (dilatometer test) modulus 30 MPa, borehole shear
test friction angle 327, estimated total unit weight 15.5 kN/m3,
and crosshole shear wave velocity 240 m/s. The water table is
4.9 m deep. Summary profiles of soil tests performed near the
footings are shown in Figs. 1–3. The exact location of the
soundings is shown in Fig. 4. Additional data can be found in
Briaud and Gibbens (1994).
Geologically, the top layer of sand is a flood plain deposit
of Pleistocene age (Jennings et al. 1996) about 3 m thick with
a high fine content. The next layer of sand is a river channel
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I
788 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENG
FIG. 1. SPT Profile
FIG. 2. CPT Profile
FIG. 3. PMT Profiles
tension the 12-MN maximum downward load that could be
applied to the footings. The dywidag bars went from the re-
action beam down to the bottom of the shaft, but only the
portion of the shaft in the shale was filled with concrete. The
portion of the shaft in the sand was filled with sand to mini-
mize the influence of the reaction shaft on the footing and soil
behavior.
The load was measured with a 12-MN load cell resting be-
tween the jack and the load frame. The settlement was mea-
sured as the average reading of four LVDTs (linear variable
displacement transducers) placed at the corners of the footings
and tied to two reference beams, the length of which is givenin Table 2. As was learned from the five tests in this study,
reference beams are not well suited for such large footing tests;
indeed, long steel beams are sensitive to temperature changes
and long wooden beams can creep significantly. It is much
easier to measure the settlement of a large footing by a deep
telltale placed, for example, at 4B below the center of the
footing, where B is the footing width. An LVDT is tied to the
telltale rod and the stem rests on the footing to measure the
difference in movement between the footing and the 4B-deep
point in the soil. The value of 4B is an educated guess. Three
such telltales were placed under each footing at 2B, 1B, and
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NEERING / SEPTEMBER 1999
FIG. 4. Plan View Showing Footings and Borings
TABLE 1. As-Built Dimensions of Footings
Footing
number
(1)
Length 3
width
(m)
(2)
Thickness
(m)
(3)
Embedment
depth
(m)
(4)
In text,
referred to as
(5)
1 3.004 3 3.004 1.219 0.762 3 m north footing
2 1.505 3 1.492 1.219 0.762 1.5 m footing
3 3.023 3 3.016 1.346 0.889 3 m south footing
4 2.489 3 2.496 1.219 0.762 2.5 m footing
5 0.991 3 0.991 1.168 0.711 1 m footing
FIG. 5. Cross Section of Footing Setup
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JOURNAL OF GEOTECHNICA
TABLE 2. Length of Settlement Beams
Footings
(1)
Distance between settlement
beam supports
(2)
1 m 7.625 m (7.625B)
1.5 m 9.3 m (6.2B)
2.5 m 10.43 m (4.17B)
3 m (south) 13.98 m (4.66B)
3 m (north) 13.98 m (4.66B)
FIG. 6. Load-Settlement Curve for 1 m Footing
0.5B deep below the footing embedment depth (Fig. 5) to
monitor the vertical movement of the soil versus depth. In the
plan view, they were located at 0.4 m from the center of the
footing. In addition, inclinometer casings were installed near
the edge of the footings as shown in Figs. 4 and 5 to monitor
the lateral movement of the soil versus depth. The location of
the casings in the plan view corresponds to the holes labeled
SPT1 through 5 in Fig. 4. Their depth is in Fig. 5.
The testing procedure consisted of applying the load in in-
crements equal to one-tenth of the estimated footing capacity
as determined by commonly used bearing capacity calculation
methods. Each load step lasted 30 min, with settlement read-
ings at 1, 3, 5, 7, 10, 20, and 30 min. This 30-min period was
considered sufficiently long to bring the settlement rate at the
end of each load step to a very small value and to calibrate
the creep settlement model while deeping the load-test dura-
tion reasonable. The load-settlement curves are shown in Figs.
6–10. These curves show the complete history of load and
settlement during the load test, including all the settlement
readings taken during 30 min under each load. As a settlement
of about 25 mm, the load was held for 24 h while recording
the settlement. Unload-reload cycles were necessary in order
to install shims above the jack. Additional data can be found
in Briaud and Gibbens (1997).
CREEP SETTLEMENT
During each load step, the evolution of the settlement is
modeled as follows. The settlement S at a time t after the
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L AND GEOENVIRONMENTAL ENGINEERING / SEPTEMBER 1999 / 789
FIG. 7. Load-Settlement Curve for 1.5 m Footing
FIG. 8. Load-Settlement Curve for 2.5 m Footing
beginning of a load step is the total settlement of the footing
since the beginning of the load test. The settlement S1 is the
value of S for a time t1 = 1 min after the beginning of a load
step. The rate effect model proposed by Briaud and Garland
(1985) has been used to predict the time-dependent behavior
of soils. This model is
n
S t
= (1)S DS t1 1
where n is the creep exponent, a property of the soil. In order
to find n, the data are plotted as log S/S1 versus log t/t1, and
the slope of the straight line regression gives n. Typical n
values range from 0.005 to 0.03 for sands and 0.02 to 0.08
for clays.
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INEERING / SEPTEMBER 1999
790 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENG
FIG. 9. Load-Settlement Curve for 3 m North Footing
FIG. 10. Load-Settlement Curve for 3 m South Footing
Creep curves are presented in Figs. 11–14 for the 30-min-
long steps and the 24-h-long steps. As can be seen, the curves
are quite close to being linear, and therefore (1) is a reasonable
model to describe these data. Toward the end of the 24-h creep
test on the 3.0-m north footing, problems were encountered
with the long settlement beam (Fig. 14). If it is assumed that
the model is applicable from 30 min to 50 years, the settlement
at 50 years can be compared to the settlement at 30 min as
follows:
n
S 50 years50 years
= (2)S DS 30 min30 min
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FIG. 11. Creep Curve for 30-Minute Load Steps and for 1.5 m
Footing
FIG. 12. Creep Curve for 24-Hour Load Step and for 1.5 m
Footing
For a value of n equal to 0.03, the ratio of S50 years over S30 min
would be 1.51. For sands such as the one tested, this gives an
idea of the increase in settlement due to creep between the 50-
year life of a structure and the 30-min load steps of a load
test. Note that (2) represents an extrapolation of the model
much beyond the range of verification, which is 24 h.
Schmertmann (1970) proposed a creep coefficient for sands
(C = 1 1 0.2 log tyears/0.1). This coefficient C is the ratio of
the settlement at t years over the settlement at 0.1 year; for t
= 50 years, the coefficient C is 1.54. In order to get the same
settlement ratio with (2), the creep exponent n must be 0.07.
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JOURNAL OF GEOTECHNICA
FIG. 13. Creep Curve for 30-Minute Load Steps and for 3 m
North Footing
FIG. 14. Creep Curve for 24-Hour Load Step and for 3 m North
Footing
In our experience, this is a very high value of n for sands and
leads to an upper-bound estimate of creep. It does make sense
that Schmertmann chose an upper-bound value for C since this
coefficient had to be valid for all sands. In that sense, (2) offers
more flexibility because a site-specific estimate of creep is now
possible; indeed, site-specific values of n can be measured with
the pressuremeter (Briaud 1992).
Fig. 15 shows the variation of n with the stress level under
the footing. The n values are obtained from Fig. 13 for each
load step as described above, and the stress level is character-
ized by the ratio of the load Q for a load step over the ultimate
load Qu, defined as the load reached for a settlement equal to
0.1 times the footing width. As can be seen, the n values and
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AND GEOENVIRONMENTAL ENGINEERING / SEPTEMBER 1999 / 791
FIG. 15. Creep Exponent versus Relative Load Level
therefore the creep rates increase with Q/Qu during the first
monotonic loading along OA on Fig. 15. Then there is a drop
in the n value after the first unload-reload cycle (Figs. 9 and
15) followed by an increase of n with Q/Qu. This second in-
crease, along line OB from the point of origin, is different
from OA (the first monotonic loading). This phenomenon is
observed again after the second and third cycles. Therefore, n
decreases with cycles, and cyclic loading decreases the creep
rate of settlement; the larger the number of cycles, the larger
the decrease.
It was also observed that if a load Q1 is held for 24 hours
and then decreased to Q2, where it is held for a new 24 hours,
the creep under Q2 is very small compared to the creep under
Q1. This supports the idea that preloading decreases creep set-
tlement.
STRAIN VERSUS DEPTH
Figs. 16 and 17 show some of the results from the telltale
measurements as plots ofthe normalized settlement S/Stop
against the normalized depth Z/B. The parameter S is the
downward movement of a point at a depth Z in the soil mass.
The parameter Stop is the settlement of the footing, and B is
the footing width. Fig. 16 is for Stop values equal to 1% of B,
while Fig. 17 is for Stop values equal to 5% of B.
The first observation from those plots is that the settlement
of the soil at a depth of 2B below the footings ranges from 0
to 10% of the settlement at the surface and averages 3.2%.
Therefore, 97% of the settlement takes place within a depth
of 2B below the footing, and it seems appropriate to use a 2B
depth of influence for calculating the settlement of square foot-
ing on sand. This confirms previous findings at smaller scales,
including Schmertmann (1970). Furthermore, the settlement at
a depth of 1B below the footing ranged from 11 to 30% of
the settlement at the surface and averaged 22%. Therefore,
78% of the settlement occurs within a depth of 1B below the
footing, and it is very important to obtain the compressibility
properties of the soil within that zone.
The second observation relates to the use of the parameter
S/B instead of S in the presentation of load test results. Since
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INEERING / SEPTEMBER 1999
792 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENG
FIG. 16. Normalized Settlement versus Depth at Footing Set-
tlement of 0.01B
FIG. 17. Normalized Settlement versus Depth at Footing Set-
tlement of 0.05B
97% of the settlement occurs within 2B below the footing, the
average vertical strain in the soil within that zone is approxi-
mately S/2B. This extends to sand the theoretical findings for
clay published by Skempton (1951) and confirms the results
of Schmertmann (1970). It also shows that S/B is related to
the average strain under the footing. Since the pressure P un-
der the footing is related to the average vertical normal stress
under the footing, the plot of P versus S/B is directly related
to the stress-strain curve of the soil mass under the footing.
Briaud and Jeanjean (1994) showed that the P versus S/B
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FIG. 18. Pressure versus Settlement over Width Plots for All
Footings
curve was unique for five footings of different sizes tested near
the surface of the same sand deposit (Fig. 18). Therefore, this
P versus S/B presentation of the results of spread footing load
tests is preferred because it would be independent of size,
whereas a load-settlement curve is directly influenced by the
footing size. Previous data obtained by Osterberg (1947), Pal-
mer (1947), Skempton (1951), and Ismael (1985) was plotted
as P versus S/B and also showed the uniqueness of the P
versus S/B curve. These data refute the idea of a scale effect
for footings on sand up to an S/B of about 0.1 and have been
discussed in detail by Briaud and Jeanjean (1994). If P0.1B and
P0.01B are the pressures at a settlement equal to 0.1B and 0.01B,
respectively, where B is the footing width, the following re-
lationships are true for the five footings tested in this study:
P = 1.7P (3)0.1B L
P = 0.23q (4)0.1B c
P = 0.075N with N in blows/0.3 m and P in MPa (5)0.1B 0.1B
P = 0.7P (6)0.01B L
P = 0.09q (7)0.01B c
P = 0.03N with N in blows/0.3 m and P in MPa (8)0.01B 0.01B
where PL, qc, and N are the pressuremeter limit pressure, the
cone penetrometer point resistance, and the standard penetra-
tion test blow count, respectively, averaged within one footing
width below the footing.
The third observation deals with the evolution of the settle-
ment-versus-depth profile as the load increases during the load
test or as the footing size increases from one load test to the
next. The data show that, for all practical purposes, the depth
of influence and the profiles remain nearly constant as the load
increases and as the width of the footing varies.
The S/Stop versus Z/B profiles can be used to prepare an
average strain-versus-depth profile. The strain in a layer is
S(at Z ) 2 S(at Z 1 DZ )
e (at Z 1 DZ/2) = (9)
DZ
where Z is the distance from the bottom of the footing to the
top of the layer; DZ is the thickness of the layer; and S(at Z)
is the settlement at the depth Z. If
S(at Z ) = a S (10)Z top
and
DZ = b B (11)Z
where Stop is the settlement of the footing, then
(a 2 a )SZ (Z1DZ) top
e (at Z 1 DZ/2) = (12)
b BZ
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JOURNAL OF GEOTECHNIC
FIG. 19. Strain versus Relative Depth
FIG. 20. Horizontal Movement Profile at Large Deformation for
1 m Footing
Fig. 19 is a plot of e (at Z 1 DZ/2)(B/Stop) versus Z/B. The
profile shows the natural decrease in strain with depth except
close to the bottom of the footing, where the strain decreases
due to the lateral confinement brought about by the roughness
of the footing base. This is consistent with the theory of elas-
ticity profiles generated by Schmertmann (1970).
LATERAL DEFORMATION
Figs. 20 and 21 show examples of horizontal movement
profiles obtained in the inclinometer casings next to the 1- and
3-m footings at high loads; the exact loads are indicated on
the figures.
The first observation is the shape of the profile, which is
the same for the small and the large footing. The general bulg-
ing points out that the deformation pattern in the soil mass
under the footing corresponds quite well to an expansion pro-
cess, even at large deformations. This observation supports the
idea of a pressuremeter-based method for footing design
(Briaud and Jeanjean 1994) and is not consistent with the as-
sumption of a single shear slip surface analysis, as in the clas-
sical bearing capacity equation.
The second observation is related to the bottom of the bulb,
which is at a depth of 1.75B for the 3-m footing, at 3.8B for
the 1-m footing, and averaged 2.33B for all the footings. While
the 2-B depth of influence observed on vertical movement pro-
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AL AND GEOENVIRONMENTAL ENGINEERING / SEPTEMBER 1999 / 793
FIG. 21. Horizontal Movement Profile at Large Deformation for
3 m North Footing
FIG. 22. Depth of Influence as Function of Footing Size
files is within the above values, the influence of size appears
different from the vertical movement observation. This influ-
ence of footing size on the relative depth of influence is shown
in Fig. 22 for all the data collected, is consistent with results
from a finite element method simulation (Hossain and Briaud
1996), but remains unexplained at this time.
The third observation deals with the magnitude of the hor-
izontal displacement at the edge of the footing compared to
the magnitude of the vertical displacement of the footing. Fig.
23 shows the relationship between the ratio dh(max)/Stop as a
function of H/B. The parameter dh(max) is the maximum hori-
zontal movement measured with the inclinometer under the
load Q, which corresponds to Stop, H is the distance between
the footing edge and the casing, and B is the footing width.
Fig. 23 indicates that at the footing edge the maximum hori-
zontal movement is of the order of 15% of the footing settle-
ment, and the horizontal zone of influence extends to about
1.8B on each side beyond the edge of the footing. This means
that settlement beams should be at least 5B long to provide a
good measurement of the absolute settlement.
The fourth observation deals with the volume change in the
soil. Indeed, the inclinometer results make it possible to esti-
mate the volume change of the soil mass below the footing.
Fig. 24 is a schematic of the soil mass influenced by the foot-
ing. The change in volume imposed by the penetration of the
footing is (ACDB in Fig. 24)
2DV = S B (13)footing top
ct to ASCE license or copyright; see http://pubs.asce.org/copyrightd
= = 3 (19)
zontal Distance from Footing Edge
FIG. 24. Schematic of Soil Mass Influenced by Footing
TABLE 3. Comparison between Bearing Capacity Pre
Methods for predicted bearing
capacity
(1)
1.0 m footing
(MPa)
(2)
1.5 m footing
(MPa)
(3)
Briaud-CPT (1993) 1.743 1.608
Briaud-PMT (1992) 0.872 0.779
Hansen (1970) 0.772 0.814
Meyerhof (1951, 1963) 0.832 0.991
Terzaghi (1943) 0.619 0.740
Vesic (1973, 1974) 0.825 0.896
Measured pressure @ s = 150 mm
after 30 min of load application 1.740 1.511
The change in volume of the soil mass corresponding to the
B 2 area and to the depth of influence riB can be estimated as
follows. The initial volume V0 is (AEFB in Fig. 24)
3V = r B (14)0 i
The deformed volume is approximated by
2
dh(max)V = r B 3 2 1 B (15)i S D2
This assumes that the deformed volume CIEFND in Fig. 24
can be approximated by GOPH, where JK is equal to 1/2 IK
and IK = dh(max). Note also that the profile DNF is an approx-
imation of the shape of the inclinometer profiles, such as the
ones of Figs. 20 and 21. From Fig. 21 the value of dh(max) at
the edge of the footing is taken as 0.15Stop. Therefore, the
volume change of the soil mass AEFB is
794 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGI
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ictions and Measured Pressure at 150 mm of Settlement
2.5 m footing
(MPa)
(4)
3.0 m (north) footing
(MPa)
(5)
3.0 m (south) footing
(MPa)
(6)
1.737 1.892 1.892
0.781 0.783 0.783
0.769 0.730 0.730
1.058 1.034 1.034
0.829 0.826 0.826
0.885 0.885 0.855
1.136 1.000 1.139
V V B r0 i
This shows that, for an average ri value of 2.33, the soil mass
compresses slightly under the footing.
EVALUATION OF EXISTING METHODS
There are a large number of methods to calculate the bearing
capacity of footings on sand and the settlement at working
loads. Six bearing capacity methods and twelve settlement
methods were selected (Tables 3 and 4). The calculated bear-
ing capacities are shown in Table 3 for all the footings. The
details of the calculations can be found in Gibbens and Briaud
(1995). The calculated values can be compared with the mea-
sured pressures at 150 mm of settlement. One can argue that
150 mm of settlement is not sufficient to generate the bearing
capacity; nevertheless, the measured pressures at 150 mm of
settlement are larger than almost all calculated bearing capac-
ities.
For the settlement methods, a reverse scheme was adopted.
Instead of comparing the calculated settlement for a chosen
load, the load that would generate 25 mm of settlement was
calculated for all methods. These loads are shown in Table 4,
and the details of the calculations can be found in Gibbens
and Briaud (1995). They can be compared with the measured
loads at 25 mm, which are also listed in Table 4. The measured
values in Table 4 were obtained by first preparing the load
settlement curve corresponding to the envelope of the 30-min
readings (Briaud and Gibbens 1994) and then reading the load
corresponding to 25 mm of settlement on that envelope. There-
fore, 25 mm is the settlement after 30 minutes of load appli-
cation, and not some 50 years, as would be the case for a real
structure. As calculated in (2), the settlement at 50 years could
be 50% larger than the settlement at 30 min. This must be kept
in mind when evaluating the results in Table 4 for design pur-
poses. While some methods do better than others, one should
be careful not to form a definite opinion on a method based
only on the comparison with five load tests. On the other hand,
such a comparison cannot be ignored.
These large footing tests also represented an opportunity to
evaluate the WAK test (Briaud and Lepert 1990). The WAK
test consists of hitting the spread footing with an instrumented
sledge hammer and recording the footing response through
FIG. 23. Maximum Horizontal Movement as Function of Hori-
2 2DV = r B((0.15S 1 B) 2 B ) (16)soil i top
Since 0.15Stop is very small compared to B, then
2DV = 0.3r S B (17)soil i top
and
DVsoil
= 0.3r (18)i
DVfooting
The net volume change of the soil mass is the difference be-
tween DVfootings and DVsoil; the volumetric strain is
DV DV 2 DV S 1 2 0.3rfooting soil top i
NEERING / SEPTEMBER 1999
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P = 0.7P (23)0.01B L
P = 0.09q (24)0.01B c
P = 0.03N with N in blows/0.3 m and P in MPa (25)0.01B 0.01B
where P0.1B and P0.01B are the pressures corresponding
to a settlement of 0.1B and 0.01B, respectively; B is the
footing width; PL is the pressuremeter limit pressure; qc
is the cone penetrometer point resistance; and N is the
SPT blowcount.
4. The Briaud-Garland (1985) model is well suited to de-
scribe the creep behavior of this sand as measured in
the load tests (up to 24 h). If this model is extrapolated
to 50 years, it shows that the ratio of the settlement at
50 years over the settlement obtained in the load tests
(30-min load steps) is about 1.5.
5. The creep rate is sensitive to previous unload-reload
cycles, to preloading, and more generally, to previous
loading history.
This project was sponsored by the Federal Highway Administration.
Albert DiMillio of the FHWA is thanked for his vision, constructive crit-
icism, and contagious positive attitude. Michael Adams of the FHWA is
thanked for his tremendous energy, technical insight, and field-testing
knowledge. The main contractor was Geotest Engineering in Houston,
where Dr. Vijay Vijayvergiya was leading the team. Many individuals
contributed to this large-scale project, among them Dr. Silvano Marchetti
and Maurizio Calabrese (University of L’Aquila), Dr. Philip Buchanan
(Buchanan Soil Mechanics), Dr. George Goble (Goble-Rausche-Likius),
Recep Yilmaz (Fugro McClelland), Dr. Alan Lutenegger and Dr. Don
Degroot (University of Massachusetts), Dr. Derek Morris and Tony Yen
(Texas A&M University), Glyen Farmer (Drillers, Inc.), Dr. Marc Bal-
louz, Jim Maxwell, Dr. George Nasr, and Dr. Philippe Jeanjean (Texas
A&M University). They are all thanked for their help.
APPENDIX. REFERENCES
Ballouz, M., Maxwell, J., and Briaud, J. L. (1995). ‘‘WAK tests on five
full scale footings in sand.’’ Rep. to the Federal Highway Administra-
tion, Dept. of Civ. Engrg., Texas A&M University, College Station,
Tex.
TABLE 4. Comparison between Predicted
Methods for predicted load
@ s = 25 mm
(1)
1.0 m footing
(MN)
(2)
1.5 m footing
(MN)
(3)
Briaud (1992) 0.904 1.314
Burland and Burbidge (1985) 0.733 1.148
DeBeer (1965) 1.14 0.803
Menard and Rousseau (1962) 0.247 0.394
Meyerhof-CPT (1965) 0.288 0.446
Meyerhof-SPT (1965) 0.195 0.416
Peck and Bazarra (1967) 1.042 1.899
Peck et al. (1974) 0.319 0.718
Schmertmann-CPT (1970) 0.455 0.734
Schmertmann-DMT (1986) 1.300 2.165
Schultze and Sherif (1973) 1.465 2.615
Terzaghi and Peck (1967) 0.287 0.529
Measured load @ s = 25 mm
after 30 min of load applica-
tion 0.850 1.500
geophones. Analysis of the data using Fast Fourier Transforms
leads to an estimate of the static stiffness of the footing-soil
assembly, that is to say, the slope of the load-settlement curve
obtained from a load test. The WAK tests predicted a secant
that corresponds to settlement of about 10 mm for all footings
(Ballouz et al. 1995).
CONCLUSIONS
The following conclusions are based on the load testing of
five square spread footings from 1 to 3 m in size embedded
0.75 m in a medium dense fine silty silica sand.
1. Settlement beams should be at least 5B long where B
is the footing width as indicated by the inclinometer
readings away from the footings. A better way to mea-
sure the footing settlement, however, is through a tell-
tale anchored 4B below the center of the footing.
2. When the load settlement curves are plotted as pressure
P versus settlement over width S/B, thefive curves col-
lapse into one and the apparent scale effect disappears.
3. For these footings, the pressure at large deformation
(bearing capacity) and the pressure at small deformation
(working loads) can be related to PL, qc, and N as fol-
lows:
P = 1.7P (20)0.1B L
P = 0.23q (21)0.1B c
P = 0.075N with N in blows/0.3 m and P in MPa (22)0.1B 0.1B
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and Measured Loads at 25 mm of Settlement
2.5 m footing
(MN)
(4)
3.0 m (north) footing
(MN)
(5)
3.0 m (south) footing
(MN)
(6)
2.413 2.817 2.817
2.175 2.799 2.799
0.617 0.597 0.597
0.644 1.017 1.017
0.738 0.918 0.918
1.000 1.413 1.413
4.144 5.679 5.679
1.981 2.952 2.952
1.475 1.953 1.953
4.114 5.256 5.256
4.750 5.850 5.850
1.244 1.476 1.476
3.600 4.500 4.500
6. Most of the settlement S is concentrated within one
footing width (B) below the footing, with 78% of the
settlement S occurring within 1B and 97% occurring
within 2B. The ratio S/2B is an approximation of the
average vertical strain within 2B below the footing. The
results confirm, at the large scale of these footings, ear-
lier findings on smaller footing tests.
7. The lateral deformation profiles show a general bulging
of the soil mass under the footing indicative of a cavity
expansion process much like that measured by the pres-
suremeter; this observed cavity expansion process is not
consistent with the assumption of a single shear slip
surface analysis, as in the classical bearing capacity
equation.
8. The maximum measured lateral movement at the edge
of the footing was 15% of the vertical movement, and
the bottom of the lateral bulge occurred at a depth of
1.75B for the 3-m footing and the 3.8B for the 1-m
footing. This large variation in the depth of influence
and its size dependency is in contrast with the 2B depth
of influence found from the vertical displacement mea-
surements, and remains unexplained.
9. Estimates of volume change in the soil mass under the
footings at large deformations indicate that the sand
contracts on the average.
10. Twelve settlement methods, six bearing capacity meth-
ods, and the WAK test are evaluated by comparing the
predictions with the measurements.
ACKNOWLEDGMENTS
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