longitudinal stresses in pressure pipeline design – a critical review

1

Longitudinal Stresses in Pressure Pipeline Design – A Critical Review

By

Sri K. Rajah, Ph.D., P.E., M.ASCE.1 and Souheil Nasr, M.S.C.E., P.E., M.ASCE2

Abstract

Current standards of practice for the design of steel water pipelines do not adequately andconsistently address the potential longitudinal stresses in a pipeline, especially for highpressure applications. For steel pipelines, the low-pressure applications are typicallydesigned using AWWA design manual M11 and the high pressure pipelines are typicallydesigned using ASCE manual of practice for steel penstocks. In handling the longitudinalstresses in the pipeline due to internal pressure, both documents consider the Poisson’seffect, but do not consider the complete pressure-axial strain relation (often known asBourdon effect). In contrast, most mechanical piping designers use this effect incalculating stresses using piping analysis software, particularly for high pressureapplications. Despite this shortcoming, current design practice for water pipelines hasresulted in numerous successful projects. However, there have also been several failuresdue to excessive longitudinal stresses and poor or inadequate design of welded joints andthrust supports.

This paper presents a comparison of current design standards by AWWA, ASCE andASME and a proposed approach for estimating the longitudinal stresses for the design ofburied pressure pipelines. Also, the paper presents a verification of this approach using apublished case history of a pipeline failure.

INTRODUCTION

In water/wastewater practice, pipeline design is primarily based on the two-dimensionalbehavior of the pipe sections. The longitudinal stresses in the pipeline are consideredsecondary. It is generally recognized that the longitudinal stresses could become primarydesign stresses for above ground pipelines and for buried pipelines with variable supportconditions along the alignment or welded joints near bends. Despite this recognition,current water pipeline design standards and guidelines do not adequately and consistently

1 Senior Structural Engineer, HDR Engineering, Inc., 500 108th Ave NE, #1200, Bellevue,Washington, USA 98004, PH: 1-425-450-6269, email: [email protected] Principal Engineer, City of Everett, Public Works, 3200 Cedar Street, Everett, Washington, USA98201, PH: 1-425-257-7210, email: [email protected]

Pipelines 2008 © 2008 ASCE

Copyright ASCE 2008 Pipelines Congress 2008 Pipelines 2008

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address the handling of potential longitudinal stresses in designing a pipeline (Jeyapalanand Abdel-Magid 1987; Jacob et al. 2007).

Longitudinal stresses in a pipeline could result from non-uniform transverse loads(changes in cover, traffic loads and other point/concentrated loads); variable soil support(soil bearing) along the pipeline due to variation of subsurface soils and installationconditions; settlement and permanent seismic ground displacement; thermal loadingcoupled with restraints for movement; and internal pressure induced axial strains both atbends and straight segments. For the general problems, it is customary to evaluate eachload separately and superimpose the structural response to obtain the composite stress-strain behavior of the pipe, which assumes elastic response of the pipe despite the non-linear elasto-plastic behavior of the soil-pipe system and welded joints.

The importance of considering longitudinal stresses and strains arising from non-uniformor concentrated transverse loads and variable support conditions in reinforced plasticmortar (RPM) pipeline design have been emphasized by Jeyapalan and Abdel-Magid(1987) among others (Carlstrom 1981; Spangler 1954; Kitching and Kirk 1980). Whilethe same concerns are applicable for the welded steel pipe (WSP) design, a review of casehistories of the past WSP failures (Phillips et al., 1972; Eberhardt 1980; Moncarz et al.1987; and Jacob et al. 2007) show that the fabrication of the bells, configuration of thewelded joints, and the resulting behavior of the joints when subjected to significantthermal, internal pressure and seismic loads, have been of primary concern.

Internal Pressure induced Longitudinal Strain

Internal pressure would induce radial and axial (longitudinal) stresses and strains in a pipe,either straight or curved. This internal pressure-strain behavior consists of two distinctmechanisms: Poisson effect and Bourdon effect. Of these, the Poisson effect is a wellunderstood behavior, in which the pipe contracts axially due to the circumferential andradial stress components. On the other hand, Bourdon effect is not well understood to thisdate, despite successful application of the concept by Bourdon (Van der Pyl 1953) inpressure gages. The Bourdon effect is generally considered to act on bends, elbows andcurved pipes by trying to straighten them when they are pressurized, which results inextensional and rotational strains on the pipe. Similar, internal pressure inducedextensional strain is present in a straight pipe as well, as evidenced in a pressurizedcylinder with end caps. For simplicity, internal pressure induced extensional strain instraight and curved pipe is referenced as ‘Bourdon effect’ in this paper.

In most pipeline analysis models, pipeline engineers represent the effect of internalpressure with equivalent thermal loads by equating the axial contraction due to internalpressure induced Poisson contraction and an equivalent thermal load. While this approachis based on a reasonable concept, the exclusion of extensional strains from the Bourdoneffect results in wrong equivalent thermal loads. Also, as shown in Table 1, the pressureinduced stress-strain behavior is not entirely mirrored by the thermal induced stress-strainbehavior. For example, if a pipe is free to expand or contract, thermal loads would resultonly in axial strains, while internal pressure loading would result in both axial stresses andstrains. Without consideration of the Bourdon strain, strain-stress compatibility is notensured for certain boundary conditions thus analysis results using these equivalent



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thermal loads are unreliable, particularly for high pressure applications for which theseeffects are significant.

Table 1: Stress and Strain in Straight Pipe under Internal Pressure and Thermal LoadsThermal Internal Pressure

LongitudinalStrain

LongitudinalStress

LongitudinalStrain

LongitudinalStress

Unrestrained Pipe, closedend -- (End Cap effect)

Yes No Yes Yes

Unrestrained Pipe, freeend -- (open end)

Yes No Yes No

Restrained Pipe –(plane strain)

No Yes No Yes

Pipeline Design Standards and Codes

The AWWA standards (C200 and C206) and Design Manual M11 (2005) are generallyintended towards low-pressure steel pipeline applications. Most designers often recognizethis fact and base high pressure steel pipeline designs on the guidelines presented in ASCEmanual of practice 79 for steel penstocks (1993). In handling the longitudinal stresses inthe pipeline due to internal pressure, both documents consider the Poisson’s effect, but donot consider the complete pressure-axial strain relation. Both ASCE and AWWAdocuments address the thermal stresses in above ground pipelines, but recommendinstallation practices such as, closure joints at night, to control thermal stresses in buriedpipelines (Section 4.2.6, AWWA C206-03; AWWA M11, and Section 5.4.2, MOP 79).Per AWWA C206, purchaser is responsible to evaluate the impact of thermal stresses.

Based on the maximum allowable joint efficiencies recommended by ASME Boiler andPressure Vessel (B&PV) Code (Table UW-12, Section VIII, Division 1), ASCEguidelines define weld joint reduction factors to reduce allowable longitudinal stresses atthe joints. These factors depend on the type of the welded joint and the type and frequencyof weld inspections. ASCE guidelines also caution the designer of longitudinal stresses atthe welded joints, especially with single or double-welded lap joints; these joints arelimited to thinner wall pipes (max thickness of 5/8-inch for circumferential joints) (ASCE1993). ASCE guidelines (1993) also note that single or double-welded lap joints are oftenused for internal pressures up to 250 psi. AWWA standards, however, do not considerjoint efficiencies for different weld types nor place any limitation with regard to internalpressure.

In addition to the welding joint efficiency factors, ASME codes (B&PV Code; B31.3; andB31.1) consider flexibility and stress concentration factors in the design of piping andcomponents. The flexibility and stress concentration factors are based on the analysis byRodabaugh and George (1957) and are incorporated in the codes to account for stressconcentration effects and to ensure adequate flexibility in the piping system.

Buried Pipelines and Soil-Pipe Interaction

While, the primary focus of the ASME codes is above ground piping, most pipelines builtusing AWWA and ASCE standards are buried pipelines. However, the interactionbetween the pipe and the soil behavior is complex and is difficult to obtain closed formsolutions due to the variability of loads and subsurface conditions along the alignment. For



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simplicity, the design of a buried pipe is treated as a two dimensional problem in the planeperpendicular to the pipe axis. This approximation is simplistic in that it does not modelthe soil-pipe interaction along its axis, which has a significant impact on the longitudinalstresses and strains on the pipeline. However, even with simplified assumptions, thisgeneral problem is too difficult to be solved by hand calculations and require numericalmodels (ALA 2001; Rajah et al. 2004). For most pipelines such a sophisticated numericalanalysis is often not warranted. On the other hand, complex loading conditions and/or theneed to optimize the design may warrant such a sophisticated analysis.

Most engineers ignore soil-pipe interaction effects, especially for the design of buriedstraight pipe segments. They estimate longitudinal tensile stresses from Poisson andtemperature effects, while some also add an additional tensile force to account for end-capeffects. However, simple force equilibrium equation would show that internal pressureinduced longitudinal stress in a straight pipe can not be greater than end-cap force inducedstress. While this approach may seem conservative, AWWA design manual (M11) doesnot consider the effects of stress concentrations and the reduced structural capacity of thebell section in the lap joints (Tawfik and O’Rourke 1985; Brockenbrough 1990), whichmakes it less conservative and possibly unsafe when thermal stresses are accounted for.This may explain why several WSP failures (Phillips et al., 1972; Eberhardt 1990; Jacobet al. 2007) were noted when the thermal loads were significant.

THEORY

Straight Cylindrical Pipe under Internal Pressure

The general stress-strain solution for thick, long, straight cylindrical pipes subjected tointernal and external pressures was originally developed by Lamé, which can be found inmost applied mechanics textbooks (Popov 1990; Harvey 1991). For pipelines subjected toonly internal pressure, the Lamé’s solution for thick and thin wall pipelines can besimplified as summarized in Tables 2 & 3 for the most common pipe end conditionsencountered in practice. Figure 1 shows the element stress diagrams of pipe sectionssubjected to internal pressure and the notations used in the equations presented in Table 2.

The commonly encountered end conditions in a pressure pipeline include: “closed end”,“open end” and “plane strain” conditions as illustrated in Figure 2. With closed endconditions, internal pressure in the pipeline would cause axial extensional strains in the

Figure 1: Stresses in a Thick Walled Cylinder Figure 2: Most common end conditions in apipe under internal pressure



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pipe, which is commonly referred to as “end cap effect”. Plane strain condition will bepresent in a fully constrained pipeline (e.g., buried) that results in no axial strain in thepipeline. Open end conditions are rare in a pressure pipeline, except where the fluid isbeing discharged to the atmosphere. Plane stress condition is not possible in a longpressure pipeline; therefore, the corresponding stress-strain equations are not included inTable 2. Where, tσ is tangential (hoop) stress (psi); rσ is radial stress (psi); lσ is

longitudinal stress (psi); ir is inner radius of the pipe (in); or is outer radius of the pipe(in); r is radial distance of a point in the pipe wall (in); ν is Poisson ratio; E is Modulus

of elasticity of the pipe material; t is pipe wall thickness; rε is the radial strain in the pipewall; tε is the tangential (hoop) strain in the pipe wall; and lε is the longitudinal (axial)strain in the pipe wall.

Table 2: Stresses and Strains in Thick Cylinder under Internal Pressure (After Lamé)End Conditions Stress StrainTangential & Radial Stress andStrains

−

−=

+

−=

2

2

2

22

2

2

22

2

1

1

r

r

rr

rp

r

r

rr

rp

o

io

ir

o

io

it

σ

σ

( )( )

( )( )

−

−+

−

+=

==

22

22

22

222

;

io

o

io

oi

rt

rr

rr

rr

rr

Er

pru

dr

du

r

u

ν

εε

Longitudinal Stress and Strain -Closed End (i.e., “End CapEffect”)

−=

22

2

io

il

rr

rpσ ( )

−

−=

22

221

io

il

rr

r

E

p νε

Longitudinal Stress and Strain -Open End

0=lσ

−−

=22

22

io

il

rr

r

E

pνε

Longitudinal Stress and Strain –Plane Strain

−=

22

2

2io

il

rr

rpνσ

0=lε

Table 3: Stresses and Strains in Thin Cylinder Pipe Internal Pressure (After Lamé)End Conditions Stress StrainTangential & Radial Stress/ Strains

or

ir

t

rr

rrpt

pD

===−=

=

,0

,2

σσ

σ

( )trEt

pru

dr

du

r

u

i

rt

ν

εε

+=

== ;

Longitudinal Stress/Strain - ClosedEnd (i.e., “End Cap Effect”) t

pDl 4=σ

( )t

pD

El 4

21 νε −=

Longitudinal Stress/Strain - Open End 0=lσt

pD

El 4

2νε −=

Longitudinal Stress/Strain – PlaneStrain t

pDl 2

νσ = 0=lε

Curved Cylindrical Pipe under Internal Pressure

The stress-strain relations of curved pipe, elbows and bends are considerably morecomplicated, and satisfactory stress-strain solutions are not available in the literature (Kingand Crocker 1973). The behavior of a curved pipe under in-plane bending was firstexplained by von-Karman in 1911 (Lubis and Boyle 2004), by introducing the notion of a



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Figure 3: Curved Cylindrical Pipe

“flexibility factor” and “stress intensification factor” to compare the stress-strain behaviorof the curved pipe with that of a straight pipe, and was later modified by Clark andReissner (1951). However, it was later realized by several authors (Kafka and Dunn1956; Crandall and Dahl 1956; Rodabaugh and George 1957; and Reissner 1959) that thebending and internal pressure response could not be additive, due to non-linear interactionof the responses. The effect of the internal pressure response, i.e., increased flexibility ofthe curved pipe, is well known and has been incorporated into piping design codes in theform of design (stress concentration and flexibility) factors.

Most common approximation to obtain the stress-strain distribution of a curved pipe is toapply membrane analysis, by neglecting discontinuity stresses at the boundaries betweencurved and straight pipe segments. With this approach, the hoop and longitudinal stressesin the curved pipe can be expressed as:

++

=θθσ

rSinR

rSinR

t

pDh

2

4(1a)

t

pDl 4=σ (1b)

Where, R is the centerline radius of the curved pipe and the remaining notations aredenoted on Figure 3. Using the membrane analysis approximations, the longitudinalstrain on the pipe could be derived as:

( ) ( )

−

+−=R

rSin

Et

pDl

θννε 121

4(2)

The first term in the above equation denotes the unit centerline elongation (axial strain),which is the same as the longitudinal strain value in a straight pipe. The second termdenotes the unit elongation along the pipe cross-section due to the unit angulardeformation of the curved pipe under internal pressure.

Straight Cylindrical Pipe under Thermal Load

Thermal stresses result in a pipeline when the thermal expansion or contraction of the pipedue to temperature changes is constrained. On the other hand, thermal stresses will notdevelop in the pipeline, if the pipe is free to expand in all directions. If the pipe is free toexpand in the radial direction, but is constrained in the axial direction, the resultinguniaxial thermal stress in the pipeline will be given by:

( )instol TTE −−= ασ (3)

Where, oT is the temperature on the outside surface of the pipe ( F° ); instT is theinstallation temperature of the pipeline ( F° ); E is the elastic Young’s modulus (psi); andα is the thermal expansion coefficient of the pipe material (in/in/ F° ).



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If the pipe is constrained both in radial and axial directions, the resulting axial thermalstress in the pipeline will be given by:

( ) ( )instol TTE

−−−

=νασ21

(4)

When the temperature on the inside surface of the pipe, iT ( F° ) differs from thetemperature on the outside surface, oT ( F° ), additional thermal stresses are induced due tothe temperature gradient across the thickness of the pipeline. Considering a symmetrictemperature distribution, T, with respect to pipe axis and constant along the axis, theprincipal stresses away from the end effects could be expressed as:

( )

−

−

−

−= ∫ ∫

o

i i

r

r

r

rio

ir TrdrTrdr

rr

rr

r

E22

22

21 νασ (5a)

( )

−+

−

+

−= ∫ ∫

o

i i

r

r

r

rio

it TrTrdrTrdr

rr

rr

r

E 222

22

21 νασ (5b)

( )

−

−−= ∫

o

i

r

riol TTrdr

rr

E22

2

1 νασ (5c)

The principal stresses in the pipeline, including the induced thermal longitudinal (axial)stresses due to the thermal gradient across the pipeline wall could be calculated, if thethermal gradient function is known. These stresses are calculated using the logarithmicand linear thermal gradient functions given below, and the results are summarized inTable 4.

Table 4: Thermal Stresses in a Straight Thin Cylinder PipeEnd Conditions Tangential & Longitudinal (axial) StressesAxially Constrained, but Radially Free ( )

0=−−=

t

instol TTE

σασ

Axially and Radially Constrained

( ) ( )instol TTE

−−

−=

νασ

1Logarithmic thermal gradient across the pipewall

( ) ( )( )

−=

ioe

oeoi rr

rrTTT

log

log

( ) ( ) ( )( )

( ) ( ) ( )( )ν

ασσ

να

σσ

−−

==

−−

−==

12

12

oiolot

oiilit

TTE

TTE

Linear thermal gradient across the pipe wall

( ) ( )( )

−−

−=io

ooi rr

rrTTT

( ) ( ) ( )( )

( ) ( ) ( )( )ν

ασσ

να

σσ

−−

==

−−

−==

12

12

oiolot

oiilit

TTE

TTE

PROPOSED APPROACH

In current practice, the longitudinal stresses from thermal and Poisson effect loadings areestimated by considering the strain-stress interdependencies. While there is no explicittreatment of Bourdon effect, the resulting stresses are tacitly considered as the “end-capeffect” in the analyses of straight pipes subjected to internal pressure. However,longitudinal strains from the Bourdon effect and soil-structure interaction effects are not



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considered by most engineers and design guidelines. This is a concern when thermalstresses and pressure-induced stresses are computed independently, without properconsideration of axial strains and soil-pipe interaction, and superimposed to design thewelded connections in a WSP.

Due to the differences in pipe temperatures at the time of installation ( instT ) and operation( operT ), thermal strains would develop if the pipe is allowed to expand. No longitudinal

stresses would develop if the pipe is allowed to expand. If the pipe, however, isconstrained axially and radially, then longitudinal tensile stresses would develop in thepipe, which could be expressed as,

( )T

El ∆

−

−=

βνασ

1(6)

Where, ( )instoper TTT −=∆ and β is a parameter representing the soil constraint against the

pipe expansion in the radial direction; its value could vary from 0 (for no radial constraint)to 1 (for complete radial constraint).

A review of the stress-strain relationships presented in Tables 2 and 3 above show that theinternal pressure induced axial strains from Bourdon and Poisson effects could be treatedas equivalent thermal strains. In other words, the internal pressure strain effects could berepresented with an equivalent thermal load, as follows:

( ) ( )t

pD

ETip 4

211

ανν β −−

−=∆ (7)

Therefore, the total equivalent temperature differential ( eqT∆ ) could be expressed as:

( ) ( )t

pD

ETTT instopereq 4

211

ανν β −−

−−=∆ (8)

Unlike the thermal load, the internal pressure will induce longitudinal stresses in the pipewhen the pipe is allowed to expand. The magnitude of this longitudinal stress (tensile) isgiven by,

( )t

pDfreeipl 4_ =σ (9)

When the pipe is constrained and not allowed to expand, as in the case of a plane strainproblem, the longitudinal stress induced by the internal pressure will be given by,

( )t

pDconstrainipl 2_

νσ = (10)

Using the above expressions, the longitudinal stress induced in a buried pipe due tothermal and internal pressure loadings could be expressed in terms of equivalent thermalload as:

( ) t

pDT

Eeql 21

ννασβ

+∆−

−= (11)

where, the first term represents the longitudinal stress corresponding to the equivalentthermal load (i.e., temperature differential) and the second term represents the internalpressure induced longitudinal stress corresponding to plane strain conditions.

Substituting for total equivalent temperature differential ( eqT∆ ) from equation (8), the

expression for longitudinal stress in straight pipe could be simplified as:



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( )( )

t

pDTT

Eoperinstl 41

+−−

=βν

ασ (12)

In reality, the temperature on the interior and exterior walls of the pipe would be different,and would closely reflect the temperature of the contents (water) and the surroundingground, respectively. Reflecting this, the longitudinal stresses on the inner and outerwalls, liσ and loσ could be expressed as:

( ) ( ) t

pDTTTTTE iooiinst

li 41212

2+

−−

+−

−−=

ννασ

β(13)

( ) ( ) t

pDTTTTTE iooiinst

lo 41212

2+

−−

−−

−−=

ννασ

β(14)

The above equations are valid for buried pipe segments subjected to no frictional forcesand away from restraints, supports, and welded joints. In reality however, as the frictionalforces on the pipe, the longitudinal stresses in a buried pipeline would also vary along thelength of the alignment. Accounting for this, the value of β would vary between 0 and 1along the pipe. Also, near the welded joints, longitudinal stresses would be subjected tostress intensification resulting from eccentrically transmitted axial loads and the resultingrotation of the joint. Accounting for this stress intensification, the longitudinal stresses onthe inner and outer walls, liσ and loσ could be expressed as:

( ) ( ) t

pDTTTTTE iooiinst

li 41212

2+

−−

+−

−−=

νναησ

β(15)

( ) ( ) t

pDTTTTTE iooiinst

lo 41212

2+

−−

−−

−−=

νναησ

β(16)

where, η is the stress intensification factor at the welded joint and is dependent on theconfiguration of the joint. Without considering beneficial contributions of the hoopconstraint, the stress intensification factor for a single welded lap joint can be derived as:

t

g34+=η (17)

where, g is the gap between the outer surface of the spigot and inner surface of the bellends as shown on Figure 4. For a zero-gap weld, this stress intensification factor wouldreduce to 4.0.

Figure 4: Bell and Spigot type Welded Lap Joint

As reported by Tawfik and O’Rourke (1985) and Brockenbrough (1990), the axialstrength of the bell of a single or double welded lap joint will be much lower than thestrength of the pipe barrel. Joint efficiency factors used by ASME BP&V code (2004)and ASCE MOP 79 (1993) for single or double welded lap joint in a buried pipe are 0.45and 0.55 respectively. In addition, the stress intensification factor used by B31.1 code(Appendix D) for fillet welded joints is 2.1, which results in overall axial strengthreduction factors of 4.6 and 3.8 for single or double welded lap joints, respectively.



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VERIFICATION USING A CASE HISTORY

To verify the validity of the proposed approach to estimate the combined longitudinalstresses on a straight cylindrical pipe from thermal and internal pressure loads, weconsider a recently published case history of welded steel pipeline (WSP) failure (Jacob etal. 2007).

City of Atlanta – 72-inch WSP Failure

The 72-inch buried WSP was installed in the early seventies as part of the City ofAtlanta’s water pipeline system to transmit pumped raw water from the ChattahoocheeRiver to the Hemphill Reservoirs in Northwest Atlanta. The pipe was ½-inch thick steelconforming to AWWA C202-64, Class B (specified minimum yield strength (SMYS) =35,000 psi; specified minimum tensile stress (SMTS) = 60,000 psi) and the fieldassembled joints were double welded bell and spigot type joints and butt straps at closurejoints. Designer’s specifications did not address the pipeline temperatures, but citedAWWA C206-62 as the welding standard. It is believed that the installation proceduresdid not consider installation temperature as a concern. The pipeline has a factory-appliedcoal-tar enamel and felt wrapped external coating system, and a field-applied ½” thickcement mortar interior lining. The bell ends were reported to be cold formed with anexpansion device and not rolled with an offset belling die.Since this pipeline was placed in service in 1975, it had numerous failures ((1980 x 2,1982, 1993, 1994, 1999, and 2003) consisting, primarily of full-circumferential brittlefractures through the bell at the toe of the internal fillet weld. The following listenumerates pertinent findings presented by Jacob et al. (2007) based on structuralassessments completed:

• All known failures have occurred in winter months (between December andFebruary) and cold water temperatures are well correlated with the incidence ofpipeline failure.

• Extrapolations from a USGS Chattahoochee River monitoring site downstream ofthe intake suggest median winter minimum daily water temperature isapproximately 48° F.

• Failure events are believed to have occurred when minimum daily watertemperature was 42°F or less.

• The steel is extremely brittle within the cold-formed bell (Charpy V-Notch < 6 ft-lbs at 60º F).

• Design internal pressure for the pipeline is 175 psig; post-installation hydrostatictest pressure of 160 psig; at the Hemphill Reservoirs, where the pipelineterminates approximately 5 miles from the pump station the static pressure is 100psig.

• The magnitude of a single pressure transient event resulting from a 40 MGDpump shutdown was measured, in 2006, approximately 2100 feet downstream ofthe pump station as: 20 psi increase in pressure above steady state and 40 psibelow steady state.

Based on the information summarized above, the longitudinal stresses on the pipe near thedouble welded lap joint were estimated with a stress intensification factor of 4.91



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corresponding to a gap measured from a sample tested (g=0.152”). Temperatures on theinterior and exterior surfaces of the pipe were assumed to be 40 and 42 F° , respectively.The calculated longitudinal stresses corresponding to different installation temperaturesare presented on Figure 5. The results show that for the conditions known and assumed tobe present, the pipeline could have been subject to tensile stresses higher than SMTS ofthe pipe material for installation temperatures above 80 F° .

SUMMARY

Reasonable estimation of longitudinal stresses is essential in the design of welded steeljoints and thrust restraint joints/supports to prevent poor or inadequate design of the WSPprojects. In water/wastewater industry, the current design standards and guidelines do notpresent guidelines to estimate longitudinal stresses present in welded steel pipelines. Thispaper presents an approach to estimate longitudinal stresses in a buried pipeline resultingfrom thermal and internal pressure loading. Also, this paper presents a verification of theproposed formulation using a published case history of a pipeline failure.

REFERENCES

American Lifelines Alliance (ALA) (2005). “Guidelines for the Design of BuriedSteel Pipe”, July 2001 with Addenda through 2005.

American Society of Civil Engineers (ASCE) (1993). “ASCE Manual and Reports onEngineering Practice No. 79: Steel Penstocks”, ASCE.

American Society of Mechanical Engineers (ASME). (2004). Boiler & PressureVessel Code, Section VIII, Division 1, Rules for Construction of PressureVessels, Parts UG & UW, New York, New York

American Water Works Association (AWWA). (1997). C200-97, Steel Water Pipe-6 In. and Larger, Denver, Colorado.

Figure 5: Longitudinal Stresses at Lap Joints as a Function ofInstallation Temperature (City of Atlanta 72-inch watertransmission line)



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American Welding Society (AWS). (2004). D1.1: 2004, Structural Welding Code –Steel, Miami, Florida.

Brockenbrough, R. L. (1990). “Strength of Bell and Spigot Joints.” J. of StructuralEngineering, ASCE, 116(7), 1983-1991.

Eberhardt, A. (1990). “108-in. Diameter Steel Water Conduit Failure and Assessmentof AWWA Practice.” J. of Performance of Constructed Facilities, ASCE,4(1), 30-49.

Harvey, J.F. (1991) “Theory and Design of Pressure Vessels”, Second edition, VanNostrand Reinhold, New York.

Hunt, L., Jacob, B., Williams, S., Marino, M., and Duppong, J. (2007). “ConditionAssessment and Rehabilitation Recommendations to Renew Raw WaterTransmission Pipelines for the City of Atlanta.” Conference Proceedings for2007 ASCE Pipelines Conference, Boston, Mass., July, 2007.

Jacob, B., Sundberg, C., Genculu, S. and Hunt, L. (2007). “Welded Lap Joint BrittleFailure: A Structural Assessment of an Atlanta 72-inch Welded Steel WaterPipe Demonstrated Need for Improvement in AWWA Standards”, ConferenceProceedings for 2007 ASCE Pipelines Conference, Boston, Mass., July, 2007.

King, R.C. and Crocker, S. (1973) “Piping Handbook”, Fifth Edition, McGraw HillBook Company.

Lubis, A. and Boyle, J.T. (2004) “The Pressure Reduction Effect in Smooth PipingElbows – Revisited”, International Journal of Pressure Vessels and Piping”,Vol 81, pp.119-125.

Moncarz, P.D., Shyne, J.C., and Derbalian, J.K. (1987). “Failures of 108-inch steelpipe water main”, J Performance of Constructed Facilities, ASCE, 1(3), pp.168-187.

Phillips, R. V., Triay, R. Jr., and Marynick, S. M. (1972). “Pipeline Problems—Brittle Fracture, Joint Stresses, and Welding.” J. AWWA, WaterTechnology/Distribution, 64(7), 421-429.

Popov, E.P. (1990) “Engineering Mechanics of Solids”, Prentice Hall InternationalSeries in Civil Engineering and Engineering Mechanics, New Jersey.

Rajah, S., Jeyapalan, J.K., Saleira, W.E., McCabe, W.M., and Grodt, R. (2004) “SoilStructure Interaction Effects in Thrust Restraint Systems of Buried Pipelines”,In Proceedings of the 2004 Pipeline Division Specialty Conference “PipelineEngineering and Construction: ‘What’s on the Horizon’”, ASCE, San DiegoCalifornia.

Tawfik, M.S. and O’Rourke, T.D. (1985). “Load Carrying Capacity of Welded SlipJoints”, J of Pressure Vessel Technology, 107(1), pp. 36-43.

Van der Pyl, L.M. (1953). “Bibliography on Bourdon Tubes and Bourdon TubeGages”, ASME Paper 53-IRD-1., presented at the Eighth National InstrumentConference, Chicago, Illinois, September 1953.

Watkins, R.K, Card, R.J., and Williams, N. (2006) “An Investigation into the Historyand Use of Welded Lap Joints for Steel Water Pipe”, Conference Proceedingsfor 2006 ASCE Pipelines Conference, Chicago, ASCE, August 2006.



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longitudinal stresses in pressure pipeline design – a critical review

Documents

steel pipelines

high pressure pipelines

pipeline failure

ground pipelines

primarydesign stresses

current design practice

buried pipelines

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