-
EFFECTS OF VEHICLES ON BURIED, HIGH-PRESSURE PIPE
By John C. Potter,1 M. ASCE (Reviewed by the Pipeline
Division)
ABSTRACT: The mechanical effects of selected types of traffic on
buried, high-pressure steel pipe are examined. A section of the
Colorado Interstate Gas pipe-line across Fort Carson's Pinon Canyon
Maneuver Site, which should be most susceptible to damage, is
studied. The relationship between depth of cover and steel pipe
deflection is clarified using field test results, with special
emphasis on dynamic effects (impact factor). The results presented
provide the technical basis for evaluating the feasibility of
pipeline protection from anticipated traffic by earth cover.
INTRODUCTION
Objective.The mechanical effects of selected types of military
traffic on buried, high-pressure steel pipe were examined in this
study. Par-ticularly considered was the section of the Colorado
Interstate Gas (CIG) pipeline across Fort Carson's Pinon Canyon
Maneuver Site (PCMS), which should be most susceptible to damage
(see Fig. 1). Preliminary analysis, as detailed in the Analysis
section, identified the 10-in. (25-cm) pipe as being the most
susceptible to damage from traffic. Therefore, the study focused on
the behavior of the 10-in. (25-cm) pipe under wheeled- and
tracked-vehicle loading. The relationship between depth of cover
and pipe stresses/deflections was clarified experimentally, with
special em-phasis on dynamic effects (impact factor). The results
of this study pro-vided the technical basis for evaluating the
feasibility of CIG pipeline protection from anticipated PCMS
traffic by earth cover.
Background.The complexity of the behavior of buried pipes goes
beyond the fact that a wide variety of pipe types, loads, soils,
and in-stallation conditions are encompassed. Buried pipe behavior
is a soil-structure interaction problem. Not only does pipe
response depend on load conditions, but the pipe response affects
load conditions. Pipe re-sponse to load conditions is primarily a
function of pipe rigidity. In this respect, three classes of pipe
are generally recognized (2). These three classes are rigid,
semirigid, and flexible. Rigid pipes exhibit a brittle-type
failure, and their cross-sectional shapes cannot be distorted
enough to change their vertical or horizontal dimensions more than
0.1% without causing damage. Semirigid pipes can withstand changes
in their vertical or horizontal dimensions of up to 3.0% without
damage. Flexible pipes can withstand changes in their vertical or
horizontal dimensions of more than 3.0% before damage occurs.
In its manual, Steel Pipe Design and Installation (AWWA Manual
No. 11), the American Waterworks Association notes that steel pipe
can al-
1 Research Civ. Engr., USAE Waterways Experiment Station,
Vicksburg, Miss.
Note.Discussion open until October 1, 1985. To extend the
closing date one month, a written request must be filed with the
ASCE Manager of Journals. The manuscript for this paper was
submitted for review and possible publication on July 24, 1984.
This paper is part of the Journal of Transportation Engineering,
Vol. Ill , No. 3, May, 1985. ASCE, ISSN
0733-947X/85/0003-0224/$01.00. Paper No. 19734.
224
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
VALVE BLOCK NO,
OBSR PIT-10"
VALVE BLOCK NO. 2
OBSR PIT 198-B"
VALVE BLOCK NO. 3
L_COL.O_RADO__ ~~K SPRINGS ( )\ FT, CARSON
PUEBLOO - ^ *
Nr-' ^ph* P\^--ITRINIDAO0V^~i.X-^
FIG. 1.Pinon Canyon Maneuver Site (PCMS)
ways function as a flexible conduit (2). The maximum
load-carrying ca-pacity of a flexible pipe is affected by its wall
thickness and section mod-ulus, but the flexible pipe is also able
to take advantage of the load-carrying ability of the soil around
it. Since it can withstand deflections as high as 20% without
collapsing, it can transfer part of the load into a radial thrust,
which is resisted by the soil surrounding the pipe. Rigid pipe
cannot withstand large deflections, and, therefore, must carry the
entire vertical load itself. This difference between rigid and
flexible pipe behavior is why the classic bending moment formulas
apply to rigid pipe, but not to flexible pipe.
Spangler also notes that flexible conduits fail by excessive
deflection, rather than by rupture of the pipe wall (11). He
recommends the in-vestigation of flexible conduit deflections and
the establishment of al-lowable deflection limits.
As suggested by the American Water Works Association, "failure"
for steel pipe means falling short of satisfactory performance
(e.g., hy-draulic capacity), rather than a state in which localized
stresses, calcu-lated using bending-moment formulas, appear to
exceed the yield stress of the material (2). Also, the compressive
stresses in the pipe due to soil loads must be overcome by the
tensile force due to internal pressure before the pipe wall can go
into tension. Therefore, the practice of sizing pipe walls to
resist hoop stress due to internal pressure is sound and
conservative. The theoretical bending stresses and the hoop stress
should not be totaled for steel (flexible) pipe as for rigid
pipe.
Spangler's equation for deflection (change in diameter) of a
flexible pipe under an earth fill is
Ax = kDr3vv
EI + 0.061 E'r3 (1)
225
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
in which Ax = deflection of the pipe, in.; k = bedding constant
(0.108 for open trench); D = outside diameter of pipe, in.; r =
mean radius of pipe, in.; uv = net vertical load on pipe, psi; =
modulus of elasticity of pipe metal, psi; I = moment of inertia of
cross section of pipe wall, per unit length, in.4 per in.; and E' =
modulus of soil reaction, psi (300 psi for untamped backfill).
A maximum design deflection of 5% of the nominal pipe diameter
is generally considered reasonable (1,11). Since the area and
hydraulic ra-dius of a pipe deflected by 5% are 99.75% of the
values for a perfect circle, the effect of a 5% deflection on
hydraulic efficiency is negligible.
It is important to note that the vertical load used in Eq. 1 is
the net vertical load. The net vertical load is equal to the total
external vertical load only in the case of an unpressurized pipe.
When the pipe is pres-surized, the resultant of the vertical
components of the internal pressure is greater than the resultant
of the horizontal components because of the elliptical shape of the
deflected pipe. The excess vertical internal pres-sure against the
upper half of the pipe counteracts the external vertical load. The
result is a decrease in the net vertical load and a decrease in
deflection in the case of a pressurized pipe (Fig. 2).
The relationship between deflection and excess vertical pressure
re-sulting from internal pressure is easily derived from simple
geometric relationships. The excess vertical pressure, A EXTERNAL
LOAD ONLY ( c | EXTERNAL LOAD PLUS
NO INTERNAL PRESSURE NO INTERNAL PRESSURE INTERNAL PRESSURE,
P
FIG. 2.Effects of Interna! Pressure
226
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
The dead load acting on the pipe may be calculated using
Marston's theory of loads (11). Evaluation of the required
parameters depends upon extensive knowledge or assumptions about
ditch geometry, soil type, and settlement ratio, or both. However,
the upper bound for a flexible conduit is the weight of the soil
prism directly above the pipe. This up-per limit has also been
confirmed by experiment (3,4). For design, the American Iron and
Steel Institute (1) and others recommend using
hy
DL = T^s ' (5)
in which h = height of soil over pipe, in.; and 7 = unit weight
of soil over pipe, pcf. Little error results from making this
conservative sim-plification, especially for shallow
installations.
A lag factor is sometimes applied to the dead load to account
for long-term deformation of the backfill at the sides of the pipe.
However, when the pressure in the pipe equals or exceeds the
dead-load vertical pres-sure, the lag factor is 1.00 (3). This is
the case for the pipeline at the PCMS. Also, the test section is
not subject to long-term effects; thus a lag factor does not appear
in these equations.
Live loads, such as static-vehicle wheel loads, are transferred
through the soil to buried conduits, essentially in accordance with
the Boussinesq distribution (11). The vertical stress due to live
load can be calculated directly from the Boussinesq equation if the
depth is greater than about twice the width of the load contact
area. For tank tracks and dual truck wheels, this minimum depth is
about 4 ft (1.2 m). For depths less than 4 ft (1.2 m), the contact
pressure must be integrated over the contact area to give the
stress increase at depth (10).
The impact factor depends upon characteristics such as vehicle
type, vehicle velocity, and ground surface roughness. Precisely how
these characteristics affect the impact factor is not well
understood. Few em-pirical data are available for estimating the
impact factors for military vehicles operating in a cross-country
environment. Thus, the impact fac-tors appropriate to this special
case were measured as a part of this study. Specific procedures and
results are presented in the Testing and Anal-ysis sections.
TESTING
Field Conditions.A subsoil survey for the proposed cantonment
area at the PCMS provided a general description of PCMS soil
conditions. The principal investigator also conducted subsoil
investigations along the PCMS pipeline right-of-way to provide
additional information.
The cantonment area investigation consisted of 40 borings,
advanced 13.4-20 ft (4-6 m) below the ground surface. The results
pertaining to the top stratum are of interest to this
investigation. The upper stratum, encountered in all borings, is a
residual lean and sandy clay (CL) with an average thickness of
about 10 ft (3 m). The natural moisture content was from 4-15%.
Temporary saturation pjoducing a moisture content of 27-37% can
result from heavy precipitation. The liquid limit was found
227
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
to be 23-44%; the plasticity index, 9-27%. Dry density ranged
from 75-110 pcf (1,200-1,760 kg/m3).
For the principal investigator's study of the pipe installation
and soil conditions, an observation pit was excavated at one
location along each of the three segments of pipeline crossing the
PCMS. The locations of these observation pits are shown in Fig. 1.
The lean and sandy clay (CL) identified in the cantonment area
survey was observed in all three pits. The backfill material around
and above the pipe was the spoil material from the trenching
operation. The backfill was found to be firm, but less dense and
more weathered than the virgin soil at pipe depth. The av-erage dry
density of the backfill was 81 pcf. The moisture content of the
backfill averaged 9%. The virgin material had a dry density of 110
pcf (1,760 kg/m3) and a moisture content of 9%.
Study of the two pits along the 8-in. (20-cm) lines indicated a
uniform trench width of 22 in. (56 cm). An inch or less of backfill
material was observed between the pipe and trench bottom. The
trench for the 10-in. (25-cm) pipe, however, was very irregular.
This irregularity appar-ently resulted from the necessity to remove
a 12-in. (30-cm) layer of sandstone located just below the ground
surface before proceeding with pipe trench construction. The
sandstone layer was removed in a 5-6-ft-wide (1.5-2.0-m) swath. The
remainder of the trench was then exca-vated. This sub-ditch was
observed to be approximately 37-in. (94-cm) wide at the top of the
pipe. Backfill between the pipe and the trench bottom was 5-6 in.
(13-15 cm) deep.
An ultrasonic thickness gage was used to determine pipe-wall
thick-ness at each pit location. In each case, the measured wall
thickness was slightly greater than the wall thickness specified
for the particular sec-tion of pipe.
Test Section.The test section was designed to emulate that
section of the PCMS pipeline which is judged most susceptible to
damage from military traffic. Instrumentation was installed to
measure pipe deflec-tions. When loaded, the pipe functions as a
proving ring, with deflection being directly proportional to load.
Thus, actual loads can be calculated from pipe deflections. These
loads were used to verify theoretical loads and determine impact
factors. As noted in the Background section, the effects of
internal pressure are easily analyzed. Pressurized pipelines
ex-hibit smaller deflections than unpressurized pipelines carrying
the same load. By not pressurizing the pipe, safety and speed of
construction were increased, while costs were decreased without
reducing the validity of the test results.
Because of its larger span, the 10-in. (25-cm) pipe is the most
damage-susceptible type of pipe in the PCMS. A 31-ft (9-m) section
by 10- by 0.188-in.- (25- by 0.5-cm-) wall thickness X42 Grade EW
pipe was used in the test section.
Pipe deflection was measured by four spring-loaded Collins
Direct Current Differential Transformers (DCDT's). The DCDT's were
mounted inside the pipe under the wheel path of the test
vehicles.
The test section was constructed in an earth-floored shelter to
facilitate maintaining the desired moisture content of 6-10%. The
instrumented pipe was placed directly on a leveled section of
shelter floor. This place-ment duplicated the bedding condition
corresponding to a flattrench
228
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
*^y^_ _ ~T"'T ~^C "^ ^ ^ ^ , , - tMi2*S= DEFLECTION
^ ^ PIPE - _ ^ X / ^ GAUGES V -INSTRUMENTATION
FIG. 3.Test Section Layout
bottom, where special bedding materials or methods are not used.
This is conservative since some ditch spoil was found between the
pipe and trench bottom in the PCMS observation pits. However,
insufficient evi-dence of special bedding treatment was detected to
justify a less strin-gent bedding condition. Actual trench width at
the PCMS was observed to be variable. This variability and the
possibility of future, additional protective cover over the
pipeline led to selection of embankment (pos-itive projection)
construction for the test section. This insures the worst-case
behavior of the test section. Test section layout is shown in Fig.
3.
The test section was constructed with a processed sandy clay
(CL) blend similar to the sandy clay of the PCMS. The blended
material had a liquid limit of 23 and a plasticity index of 6. The
material was dried to an average moisture content of less than 10%
before placement. In-place moisture contents of 5-9% were recorded
during testing. The em-bankment was constructed without deliberate
compaction to simulate the untamped backfill around the pipeline at
the PCMS. However, some compaction did result from earthmoving and
test vehicle traffic. In-place dry densities of 110-120 pcf
(1,760-1,920 kg/m3) were recorded. These densities correspond to
expected densities in areas of concentrated ma-neuver traffic. For
most areas of the PCMS, however, densities are less than those of
the test section. Thus, the expected soil-load deflections along
the PCMS pipeline should be less than those observed in the test
section.
Traffic loads were applied with both wheeled and tracked
vehicles. Vehicles and payloads were selected to simulate the
severest vehicle-loading configurations anticipated at the PCMS.
The principal wheeled test vehicle was an M51 5-ton dump truck with
a maximum highway payload of 20,000 lb (10,870 kg) and a gross
weight of 42,000 lb (22,830 kg). Since the ground contact pressure
is approximately equal to tire pressure, a maximum inflation
pressure was chosen. All tires were in-flated to 70 psi (483 kPa).
The principal tracked test vehicle was a mod-ified M48 tank
chassis. In lieu of the turret assembly, special ballast weights
were used to achieve a gross weight of 140,000 lb (76,100 kg). The
hard-surface ground contact pressure of this vehicle exceeds 80 psi
(552 kPa). The soft-surface ground contact pressure is about 13 psi
(90 ka).
For each depth of soil cover over the test section pipe, three
types of deflections were recorded: Dead-load deflection,
static-vehicle deflec-tion, and dynamic-vehicle deflection.
Dead-load deflections were de-flections due solely to the weight of
the soil covering the pipe. Static-vehicle or live-load deflections
were the maximum deflections resulting from the weight of a vehicle
either parked or rolled slowly across the top of the test section.
Dynamic-vehicle deflections were the maximum
S>
229
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
deflections due to moving vehicles. Dynamic deflections were
recorded while operating at high speeds, moving over obstructions
or humps, and while stopping or accelerating.
Two techniques (multiple passes and gage-reading interpolation)
were used to ensure that maximum deflections were recorded.
Multiple passes increased the probability of maximum dynamic
vehicle loads being gen-erated over the axis of the pipe. For these
maximum loads, the maxi-mum pipe deflections were determined by
interpolation between de-flections recorded by the gages spaced
along the pipe axis. Thus, large deflections, not occurring exactly
at a gage location could be estimated. Pipe deflection along the
axis of the pipe and in the neighborhood of gages was modeled by a
second-degree polynomial. The coefficients of this polynomial were
determined using least squares regression through the points
determined from the deflection readings and from the loca-tions of
the gages along the pipe axis. The maximum deflection of the pipe
was then taken as the maximum of this polynomial. This technique
also compensates for small errors in deflection readings.
Deflections caused by loads occurring outside of the neighborhood
of the gages can also be rapidly identified.
ANALYSIS
Test Section.Test data were used to determine the relationship
be-tween pipe deflection and cover depth. As shown in Fig. 4, the
average deflection of the pipe was directly proportional to cover
depth.
The relationships between maximum static live load and cover
depth for both wheeled and tracked test vehicles are shown in Fig.
5. The shape of these curves is similar to that of the Boussinesq
stress distribution. Static deflections caused by the tracked test
vehicle are larger than those caused by the wheeled test vehicle
for all cover depths [0-30 in. (0-76 cm)].
Pipe deflection is directly proportional to load (Eq. 1).
Therefore, test section impact factors can be computed from the
ratio of the dynamic vehicle deflections to the static vehicle
deflections.
For normal modes of vehicle operation (e.g., high-speed travel,
stop-ping, and accelerating), the impact factors are less than or
equal to 1.3. In fact, impact factors less than 1.0 were common.
For moving vehicles, load durations are too short to allow
mobilization of full static deflec-tions. This phenomenon is well
documented in pavement-design liter-ature, and is reflected in
current Corps of Engineers pavement design procedures (6,12,13).
The impact factors generated by the tracked test vehicle were
larger than those generated by the wheeled test vehicle. Under
special conditions such as obstacle crossing, impact factors as
high as 4.1 were observed.
The test data suggest an upper limit of 5 for the impact factor.
This agrees with recommendations made by Gemperline, who measured
im-pact factors generated by loaded scrapers operating on a very
rough haul road (5). Recorded impact factors were as high as 3.0,
and he estimated the maximum impact factor to be near 5. This upper
bound is also con-sistent with the results of ride tests and shock
tests conducted with var-ious vehicles by the Mobility Systems
Division of the WES Geotechnical
230
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
PIPE DEFLECTION, INCHES -0.02 o.O 0.02 0.04 0.06 0.08
i 1 r - * -
FIG. 4.Average Soil-Load Deflection versus Cover Depth
Laboratory (7,8,9). They recorded total chassis accelerations up
to 5 g's (gravity) while maneuvering vehicles over abrupt
obstacles. They also observed that vehicle operators'physical
tolerance levels dictate restrict-ing chassis accelerations to
under 5 g's. In uncontrolled environments, operators usually choose
vehicle speeds which restrict chassis accelera-tions to 3.5 g's or
less. A maximum chassis acceleration of 3.5 g's is used as the
speed-limiting criterion for cross-country traffic in the Army
Mo-bility Model (8). All of these maximum impact factors were
recorded at vehicle speeds of 5-15 mph (8-24 kph). Lower or higher
speeds resulted in reduced impact factors.
The live-load deflections for the tracked test vehicle will be
used for the analysis of pipe performance, since the tracked test
vehicle produced both the largest static live-load deflections and
the largest impact factors. The test section pipe is subject to a
maximum dynamic live-load deflec-tion up to five times the static
live-load deflection produced by the tracked test vehicle. The
maximum total deflection is found by adding the max-imum dynamic
live-load deflection to the.dead-load deflection for each cover
depth. The relationship of maximum deflection to cover depth is
> o
20
25
30
231
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
PIPE DEFLECTION, IN.
FIG. 5.Maximum Static Vehicle Deflections versus Cover Depth
shown in Fig. 6. This curve was derived by adding the average
soil-load deflection (from Fig. 4) to five times the maximum static
vehicle deflec-tions for the tracked test vehicle (from Fig.
5).
For cover depths in the range of 0-30 in. (0-76 cm), the maximum
total deflection occurs at zero cover depth, and is less than 5% of
the nominal pipe diameter. For cover depths greater than 30 in. (76
cm), the maximum live load decreases asymptotically to zero, while
the dead load increases to some limiting value determined by the
amount of soil arch-ing above the pipe.
Pinyon Canyon Maneuver Site.From Eq. 3, it is seen that
pressur-izing the pipe will result in smaller deflections for a
given external ver-tical load. Thus, deflections in the 10-in.
(25-cm) pressurized pipeline crossing the PCMS should be smaller
than those of the test section pipe. Negative projection ditch
conditions, pipe bedding, and lighter over-burden along the PCMS
pipeline will also contribute to smaller deflec-tions in the 10-in.
(25-cm) pipeline crossing the PCMS.
The net vertical pressure is proportional to the deflection. It
can be computed from the deflection using Spangler's deflection
formula (Eq.
232
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
10
> o o
20
25
PIPE DEFLECTION, IN, 0.2 0.3 T
LEGEND
O WHEELED VEHICLE
0 TRACKED VEHICLE
-CALCULATED MAXIMUM TOTAL DEFLECTION
FIG. 6.Actual Total Deflection Readings from Test Section and
Calculated Max-imum Total Deflections versus Cover Depth
1). For the 10-in. (25-cm) pipe, one might estimate: k = 0.108;
D = 10.75 in.; r = 5.28 in.; E = 30 x 106 psi; I = 5.5 X 10~4
in.4/in.; and E' = 300 psi. The deflection of the 10-in. (25-cm)
pipe at the PCMS should then be
AX10 = 0.0089o-p (6) The deflection, expressed as a percentage
of nominal diameter, is
A%in = 0.0890%, (7) For the line designated as 19A-8 in. on the
site map (Fig. 1), one might use k = 0.108; D = 8.625 in.; r =
4.174 in.; E = 30 x 10* psi; I = 1.77 x 10-3 in.4/in.; and E' = 300
psi. Then AX19A.8 = 0.00124o- (8) and A%19A.8 = 0.0156o-o (9) For
the 19B-8 in. line (shown on the site map in Fig. 1), k = 0.108; D
=
233
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
8.625 in.; r = 4.2345 in.; E = 30 X 106 psi; 7 = 3.2 X 10"4
in.4/in.; and E' = 300 psi. Then AX19B.8 = 0.00644a,, (10) and
A%19B.8 = 0.0804a,, (11) Thus, the prediction that deflections in
both sections of 8-in. (20-cm) pipe at the PCMS will be less than
those of the 10-in. (25-cm) section for any given net vertical
load, aH. Percentage deflections for the 8-in. (20-cm) sections
will also be smaller.
The external vertical load, CT , acting on the test section pipe
can be estimated from the test section deflection data and Eq. 6.
From Fig. 5, the maximum static deflection = 0.095 in. (2.4 mm) at
zero cover depth. Using Eq. 6, the corresponding external vertical
load is computed to be 10.7 psi (74 kPa). This effective vertical
load is substantially less than the actual contact pressure because
of the small size of the vehicle footprint relative to the area of
the pipe carrying the load. From Fig. 6, the max-imum predicted
deflection is 0.475 in. (12.1 mm) at zero cover depth. Using Eq. 6,
the corresponding external vertical load is computed to be 53.4 psi
(368 kPa). Using this load and Eqs. 2, 3, and 6-11, the maximum
zero cover deflections of the pressurized pipes can be predicted.
These deflections for a typical operating pressure of 735 psi
(5,068 kPa) are shown as follows:
AX Pipe in. Percent
l O l n - 0^217 2.17 19A-8 in. 0.055 0.68 19B-8 in. 0.166
2.07
CONCLUSIONS
For cover depths of 0-30 in. (0-76 cm), the test section pipe
was found to behave in the manner predicted by flexible pipe
theory. Rigid pipe theory, including load prediction and failure
mode, is not appropriate. The deflection of an unpressurized pipe
was shown to be directly pro-portional to the external vertical
load. This external vertical load is com-posed of dead-load and
live-load components. The dead-load compo-nent was found to be
slightly more than the weight of the soil prism directly above the
pipe. The live-load was found to essentially follow the Boussinesq
stress distribution for maximum vertical stress with depth.
However, due to the small size of the vehicle footprints, the
effective vertical load for deflection calculations was much less
than that com-puted from actual contact pressures.
The impact factor was found to be highly variable. The largest
impact factors were observed as vehicles crossed obstacles placed
on the em-bankment above the pipe. These impact factors were,
however, in no case as large as 5. This upper bound is consistent
with tests conducted by Gemperline (5), Schreiner and Green (8),
and others.
The largest total deflection observed in the test section was
0.273 in. (6.9 mm). Using an impact factor of 5, the maximum
anticipated deflec-
234
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
tion for the test section is 0.475 in. (12.1 mm). This
deflection occurs at zero cover depth. A 0.475-in. (12.1-mm)
deflection is less than 5% of the nominal pipe diameter [10 in. (25
cm)], and, thus, within design limits.
The 10-in. (25-cm) pipeline in the PCMS can be expected to
experience smaller deflections than the test section pipe, and is,
thus , within de-flection design limits for cover depths of 0-30
in. (0-76 cm).
Analysis using the Spangler deflection formula indicates that
both ab-solute and percent deflections for both the 19A-8-in.
(20-cm) and 19B-8-in. (20-cm) pipelines crossing the PCMS will be
smaller than those of the 10-in. (25-cm) line.
Internal pressurization of the pipe has the effect of reducing
the ver-tical load on the deflected pipe. Expected deflections in
the operating (pressurized) pipeline crossing the PCMS are, thus,
even less than those predicted by trafficking over the
unpressurized test section pipe. The maximum total deflection at
zero cover depth is estimated to be less than 2.2%.
The preceding findings indicate that the pipeline at the PCMS
(pres-surized or unpressurized) is not susceptible to damage from
random crossings by anticipated military traffic with the maximum
expected im-pacts as long as there is sufficient cover over the
pipeline to prevent its protrusion (as an obstacle), and to prevent
metal-to-metal contact with any passing vehicles. Further, any
deflections which do occur, even at points of concentrated vehicle
crossings, will be less than the recom-mended design limit of 5% of
the nominal pipe diameter.
ACKNOWLEDGMENT
Thus study was sponsored by the Commander of Ft. Carson and the
4th Infantry Div. (Mech.), Ft. Carson, Colo. The writer is also
grateful to the U.S. Army Corps of Engineers Waterways Experiment
Station for administrative and technical support , as well as
permission to publish this paper.
APPENDIX I.REFERENCES
1. American Iron and Steel Inst., Handbook of Steel Drainage and
Construction Products, 2nd Ed., American Iron and Steel Inst., New
York, N.Y., 1971.
2. American Water Works Assoc, Steel Pipe Design and
Installation, American Water Works Assoc, Manual No. 11, 1964.
3. Barnard, R. E., "Design and Deflection Control of Buried
Steel Pipe Sup-porting Earth Loads and Live Loads," Proceedings,
American Society for Testing and Materials, Vol. 57, 1957, p.
1233.
4. Braune, G. M., Cain, W., and landa, H. F., "Earth Pressure
Experiments in Culvert Pipe," Public Roads, Vol. 10, No. 9, Nov.,
1929.
5. Gemperline, M. C , "Results of the Current Creek Pipeline
Impact Loading StudyCurrent Creek PipelineCentral Utah Project,
Utah," Geotechnical Branch Memorandum, Engineering and Research
Center, U.S. Bureau of Re-clamation, Geotechnical Branch Reference
No. 82-7, 1982.
6. Horonjeff, R., Planning and Design of Airports, McGraw-Hill
Book Co., New York, N.Y., 1962, p. 327.
7. Schreiner, B. G., "Ride and Shock Test Results and Mobility
Predictions of the Swedish Scania SBAT III Cargo Truck/' U.S. Army
Engineer Waterways Experiment Staiton, CE, Technical Report
GL-81-3, Mar., 1981.
235
J. Transp. Eng. 1985.111:224-236.
Dow
nloa
ded
from
asc
elib
rary
.org
by
Li. C
o.Sa
818
1901
/mi/1
5598
5 on
06/
06/1
5. C
opyr
ight
ASC
E. F
or p
erso
nal u
se o
nly;
all r
ight
s res
erve
d.
-
8. Schreiner, B. G., and Green, C. E., "Mobility Assessment of
the Roland Wheeled Vehicle System; Report 1, Results of Field
Tests," U.S. Army En-gineer Waterways Experiment Station, CE,
Technical Report GL-82-12, Nov., 1982.
9. Schreiner, B. G., and Randolph, D. D., "Ride and Shock Test
Results and Mobility Assessment of Selected 10-Ton Cargo Trucks,"
U.S. Army Engineer Waterways Experiment Station, CE, Technical
Report GL-79-5, May, 1979.
10. Sowers, G. B., and Sowers, G. F., Introductory Soil
Mechanics and Foundations, 3rd Ed., The Macmillan Co., London,
England, 1970.
11. Spangler, M. G., "Culverts and Conduits," Foundation
Engineering, Chapt. 11, G. A. Leonards, ed., McGraw-Hill Book Co.,
New York, N.Y., 1962.
12. Yoder, E. J., Principles of Pavement Design, John Wiley and
Sons, Inc., New York, N.Y., 1959, pp. 105-119.
13. Yoder, E. J., and Witczak, M. W., Principles of Pavement
Design, John Wiley and Sons, Inc., New York, N.Y., 1975, p.
447.
APPENDIX II.NOTATION
The following symbols are used in this paper: B D E
E' g h I
k L V
