Proceedings of the 2012 9th International Pipeline Conference

IPC2012
September 24-28, 2012, Calgary, Alberta, Canada

IPC2012-90662

FITNESS-FOR-SERVICE ANALYSIS OF SKELP-END WELDS IN SPIRAL PIPES

Ming Liu and Yong-Yi Wang Steve Rapp

Center for Reliable Energy Systems Spectra Energy
Dublin, OH, USA Houston, TX, USA

KEYWORDS pipes, the helical-seam (or spiral) pipes can offer cost
advantages for as much as 10% to 15% [1].
Fitness for Service, Skelp-End Weld, Spiral Pipe
The majority of the recently installed spiral pipes were
ABSTRACT
manufactured from steel coils. Some of the spiral pipes in the
For large diameter spiral pipes, there can be one skelp-end past were manufactured from steel plates. The coils and plates
weld (SEW) in every 5-7 joints of pipes. The industry are collectively termed skelps.
acceptance of SEWs is uneven although API 5L permits SEWs
The weld joining the skelps is often called skelp-end weld
in finished pipes. A joint industry project (JIP) [1] was formed (or strip/plate end weld in API 5L terms). The position of the
to develop uniformly acceptable inspection and test plans skelp-end weld (SEW) in a pipe is schematically shown in
(ITPs) for SEWs. The development was conducted through Figure 1. The skelps are firstly joined by a partial penetration
two parallel processes: (1) fitness-for-service analysis of the weld on what will become the inside diameter (ID) side of the
SEWs under a variety of loading conditions expected in their pipe. The string with the partially completed skelp-end weld is
life time and (2) consensus building based on the best practice fed into the pipe forming process. The helical welding is
and quality control protocols. completed from ID and outside diameter (OD) sides as the pipe
This paper details the fitness-for-service analysis of SEWs. is being formed. The skelp-end weld is then completed from
A companion paper provides a summary of the recommended the OD side by applying OD welding. As a result, the partially
ITPs developed in the JIP [2]. In the fitness-for-service completed skelp-end weld is subjected to pipe forming strains.
analysis, the SEWs were subjected to a variety of loading Skelp-End Weld
conditions covering construction, commissioning, and normal
service with and without internal pressure. For in-service
loading, both static and cyclic loading was considered. The
extensive fitness-for-service analysis demonstrated that there is
no inherent integrity risk associated with the SEWs when these
welds are manufactured, tested, and inspected using generally
accepted quality control measures applied to helical seam
welds. Additional inspection and quality control for coil end
properties and T-joints are recommended in the companion
paper. Helical Weld

INTRODUCTION Figure 1 Skelp-end weld in a helical seam pipe [3]

In many of the recent onshore large diameter pipeline The acceptance of the SEW in finished line pipes varies
construction projects in North America, more helical-seam greatly [1]. Although API 5L provides provisions for the
welded pipes (SAWH or COWH in API 5L [3] terms) have testing and acceptance of SEWs, some companies do not accept
been installed than longitudinal-seam welded pipes (SAWL or SEWs at all. Other companies may accept SEWs on a case-by-
COWL in API 5L terms). This trend is expected to continue for case basis, often after additional tests specified by purchasers.
the foreseeable future. Compared to the longitudinal-seam

1 Copyright © 2012 by ASME

FITNESS-FOR-SERVICE ANALYSES determined from statistic analyses of material properties
Scope of FFS Analysis provided by project sponsors (X70) and those collected in
CRES material database (X80). The pipe and weld properties
During onshore pipeline construction, many processes such
were measured in pipe longitudinal and circumferential
as lowering-in, back fill, tie-in, and hydrotest can generate
direction, respectively. The weld mismatch was calculated by
tensile stresses in the pipe and cause tensile rupture. Among all
normalizing the weld tensile strength by the pipe tensile
the above processes, pipe lowering-in usually produces the
strength. The heat affected zone (HAZ) mismatch was
highest longitudinal stresses. Hydrotest usually produces the
determined by the ratio between the HAZ and pipe material
highest stress in pipe circumferential (hoop) direction. Due to
hardness. Mismatch greater than one represents overmatching.
the Poisson’s effect, the internal pressure from hydrotest can
It is seen that the weld is about 5% overmatched while the HAZ
also induce longitudinal stresses.
is about 5% undermatched (softened).
On the other hand, lowering-in and field cold-bending can
Table 1 Representative properties used in analysis
apply bending deformation to a pipe. The pipe could buckle on
the compressive side when the deformation is excessive. YS UTS uEL n WM HAZ
Grade Y/T
(ksi) (ksi) (%) (CSA) Mismatch Mismatch
In addition to the construction processes, the operating
loadings may also induce tensile and compressive stress to the X70 76 90 0.84 10.6 20.1 1.05 0.95
pipe. The operating pressure and temperature variations may X80 86 98 0.88 8.8 26.9 1.05 0.95
induce cyclic stresses to the pipeline and lead to fatigue flaw
growth. For finite element analyses (FEA), full tensile stress-strain
curves were created for the materials given in Table 1 with the
The FFS analyses performed in this paper covered the CSA Z662 equation [4]:
stresses from lowering-in, cold bending, hydrotest, and
n
operation. The tensile, compressive, and fatigue failure modes

    
were investigated.     0.005  y   , (1)
 
Geometry Parameters E  E   y 
Figure 2 shows a schematic drawing of a SEW where some where  and  represent engineering strain and stress,
basic parameters are given. The helical angle () is the angle respectively. The E is the material’s Young’s modulus, y is the
between the helical weld and pipe circumference. The length material’s yield strength, and n is the strain hardening exponent.
of the skelp-end weld equals to the coil width. The intersection The pipe stress-strain curves are shown in Figure 3. The
of the helical and skelp-end welds is commonly referred to as stress-strain curves of the weld metal and HAZ material were
the T-joint or T-weld. generated by scaling the corresponding pipe stress-strain curves
To avoid some practical issues in construction (e.g., girth according to the strength mismatch level in Table 1 .
weld inspection), a minimum distance of 300 mm (11.8 inch) is
800
required between the SEW and the finished pipe end [3]. In the
following analyses, the SEW is assumed to be sufficiently away
from and not affected by the pipe end. 600
Eng. Stress (MPa)

Skelp-End Weld

400
T-Joint

X70 Pipe Material

D 200 X80 Pipe material

(Helical Angle)
w
(Coil Width)  (Pipe OD)

0
T-Joint
0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14
Eng. Strain (mm/mm)

Helical/Spiral Weld Figure 3 Representative X70 and X80 stress-strain curves used
in analysis
Figure 2 Skelp-end weld geometry parameters Assessment of Tensile Integrity
Material Properties Assessment Procedures
Two pipe grades, X70 and X80, were selected for FFS For tensile failure assessment, existing flaws of certain
analyses. The pipe and weld properties are summarized in sizes were assumed to be present. These flaws are associated
Table 1. The material properties of the helical seam weld and with natural weld features which fall below weld repair criteria.
SEW were the same. The material properties in Table 1 were

2 Copyright © 2012 by ASME

The crack driving force (CDF) in fracture mechanics CTOD under pure Mode I loading. The Mode I (opening) and
terminology was used to measure the magnitude (or intensity) Mode III (slide) components of the CTD are referred to as
of loading to the flaw. The CDF is associated with applied CTDI and CTDIII, respectively.
stress or strain. A failure event is postulated to occur when the Table 2 FEA matrix
FEA Matrixof lowering-in
- Lower-in induced tension(tension)
and Hydrotest and hydrotest
CDF reaches the material’s toughness. The toughness is
Helical Wall Flaw
usually measured experimentally. A lower bound conservative WM HAZ OD (D )
Coil Width
(w)
Angle Flawed Thickness
High-
Low
Depth  No. of
Grade ( ) (t ) D /t Length Cases
OM UM Weld
toughness value can be estimated based on the past data of
(inch) (mm) (inch) (mm) (degree) (inch) (mm) (mm) (mm)
similar materials. 36 914 18.6 Skelp-
FEA Models X70
1.05 0.95 36 914
54 1372 28.5 end
weld 0.50 12.7 72
0.00
2  50 16
X80 72 1829 39.5 1.27
A typical finite element model used in the tensile integrity 72 1829 39.5
Helical
weld
assessment is shown in Figure 4. The pipe with an elliptic
surface breaking flaw on pipe ID surface was modeled with (a)
three-dimensional (3D) solid/brick elements. The flaw was
located on the fusion line (i.e., HAZ flaw) of the SEW at the T-
joint. Due to the slight HAZ softening and weld metal (WM)
overmatching, the HAZ flaw is considered to represent the
worst case. The total length of the model was set to be 6OD to (b) ID Surface h

avoid any end effect on the flaw behavior. Skelp-End Weld

HAZ HAZ
The weld high-low misalignment was modeled by relative
shift of the coils on the two sides of the skelp-end weld. The a
misalignment was uniform along the skelp-end weld and was (c)
ID Surface
gradually reduced to zero over the width of the helical weld at
the T-joints.
Figure 4 FEA model for flaws in skelp-end welds
The FEA matrix is shown in Table 2. The pipe OD and Mode I
wall thickness were 36 inch and 0.50 inch, respectively. The (Opening)

coil widths were varied and the helical angles were from 19 to CTD : Crack tip displacement
40. The maximum high-low misalignment was 1.27 mm (i.e., CTD I : Mode I crack tip displacement
CTD III: Mode III crack tip displacement
10% of the wall thickness). The flaw size was assumed to be 2-
mm deep and 50-mm long.
Lower-in

CTD I
The pipe lowering-in process was simulated by applying CTD III

axial tensile displacement at pipe ends. Given the small flaw

size relative to the pipe diameter, the uniform tension loading is Mode III
a reasonable representation of the tension side of the pipe under (Out-of-Plane Shear)

global bending. Figure 5 Schematic drawing of mode I and III loading and
Due to the orientation of the SEW, the flaw in the SEW crack tip displacement (CTD)
experiences primarily mixed Mode I and Mode III loading.
The Mode I loading opens the two flaw surfaces while the
Mode III loading creates relative slide between the two flaw
surfaces along the length of the flaw (see Figure 5). Figure 6
shows a typical plastic strain contour near the flaw tip. Due to
softened HAZ properties, the plastic strain in the HAZ is larger
and spreads wider than the strain in the weld metal. The
relative opening and sliding of the flaw surfaces are evident.
A number of crack driving force representations, such as
Figure 6 Typical plastic strain contour and deformation near the
the stress intensity factors (K), J-integral (J), and crack tip
flaw tip (Figure 4(c) location)
opening displacement (CTOD) have been developed to evaluate
the intensity of the crack-tip fields, such as stresses and strains. To help understand the effect of helical angles, a so-called
The crack tip displacement vector (CTD) has been developed load angle was introduced. The load angle is defined as the
and used for mixed mode fracture problems [5]. The CTD is angle between the flaw length and the primary stress (see
defined as the relative displacement of the two points on the Figure 7). For lowering-in analysis, the primary stress is along
two flaw surfaces at a given distance (0.5-mm in this work) the pipe longitudinal direction.
behind the flaw tip as shown in Figure 5 and Figure 6. The 0.5 The comparison of the CTDI and CTDIII induced by the
mm is selected so that the CTD is very close to the traditional lowering-in tensile stress is shown in Figure 8 for the X70 pipe

3 Copyright © 2012 by ASME

with a helical angle () of 40. The lowering-in tensile stress Hydrotest
induces higher Mode III loading than Mode I loading. Similar During hydrotest, tensile stress up to the yield strength of
relations were observed for other helical angles and pipe the pipe can be generated along the pipe circumferential
grades. direction (hoop) due to applied internal pressure. At the same
time, tensile stress can be generated in the pipe axial (or
longitudinal) direction (axial) due to axial restraint and the
Tension
Poison’s effect (see Figure 11). The hoop and axial stress can
Load Angle be calculated with the following equations,
Load Angle
PD , and PD ,
 hoop   axial  
2t 2t
Flaw in Skelp-end weld Flaw in helical weld
where P represents the internal pressure and D and t represent
Figure 7 Definition of load angle under longitudinal tension the pipe outside diameter and wall thickness, respectively. The
The total CTD driving forces of SEW flaws are shown in symbol  represents the Poison’s ratio of the pipe material. It
Figure 9 for different helical angles. For comparison, the flaws should be noted that the total resultant stress (total) is not
in helical welds were also analyzed. The helical weld high-low perpendicular to the flaw surface (i.e., the skelp-end weld) for
misalignment was 10% of wall thickness (maximum allowed) any given helical angles. If  = 0.3 is used, the total is
and uniform along the whole helical weld. The assumption is perpendicular to the flaw when the helical angle equals to
believed to be conservative. Figure 9 shows that the CTD 16.7. The helical angle is defined as the angle between the
driving force increases as the load angle increases. The SEW helical weld and pipe circumference direction (as shown in 2).
high-low misalignment has very little effect on the CTD driving 0.06
force for the SEW flaws. However, the helical weld high-low Lower-in Tension, X70 Pipe
misalignment increases the CTD driving force for the flaws in 0.05
D = 36 in, t = 0.5 in, w = 72 in,  = 40
a = 2 mm, 2c = 50 mm, h = 0.0 mm
CTD Driving Force (mm)

the helical welds. Very similar results were found for the X80
pipes (see Figure 10). The CTD driving forces of the X80 0.04 CTD_I
pipes are slightly higher than those of the X70 pipes. CTD_III
To assess the fitness-for-service of the welds, it is 0.03

necessary to compare the CTD driving force with an

appropriate measure of material’s toughness. The concept of 0.02

apparent toughness is introduced here. The apparent toughness

0.01
is the initiation toughness under low crack-tip constraint
conditions, e.g., a planar flaw in a structure subjected to
0.00
predominantly tensile loading. The apparent toughness is 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
higher than a conventional toughness measured by testing high- Applied Stress / Yield Strength
constraint specimens. The concept and the applications of the
apparent toughness may be found in a number of publications Figure 8 Mode I and Mode III crack driving force (CTDI and
[6,7,8,9,10]. CTDIII) induced by pipe lowering-in tensile stress
Past experience indicated that the lower bound apparent 0.10
Lower-in Tension, X70 Pipe
toughness (low-constraint) appropriate for the current D = 36 in, t = 12.7 mm
Helical Angle / Load Angle
assessment of the X70 and X80 pipes and welds is 0.2 mm. a = 2 mm, 2c = 50 mm
CTD Driving Force (mm)

0.08
The average apparent toughness can be 0.4-0.6 mm or higher. Solid lines: h = 0.00 mm 40/ 50 (HLW)
For all the cases analyzed, the maximum CTD driving force is Dashed lines: h = 1.27 mm
0.06 HLW: Flaw in helical weld
less than 0.1 mm when the applied stress reaches the pipe yield SEW: Flaw in skelp-end weld 40/40 (SEW)
strength. The CTD at the applied stress that equals to the yield
strength is much smaller than this lower bound toughness. 0.04
Therefore the integrity of the SEWs is sound under the tensile 29/29 (SEW)

stress of the pipe lowering-in process. 0.02

19/19 (SEW)
Figure 9 and Figure 10 also show that the CTD increases
as the load angle increases and this trend is not affected by the 0.00
flaw locations (i.e., in skelp-end or helical weld). Therefore the 0.6 0.7 0.8 0.9 1.0 1.1 1.2
flaws in the skelp-end weld can be treated as those in the Applied Stress / Yield Strength
helical weld as long as the load angle is properly adjusted. This Figure 9 CTD driving force under lowering-in tension (X70)
observation is further demonstrated in hydrotest analysis.

4 Copyright © 2012 by ASME

 0.10 0.04
Lower-in Tension, X80 Pipe Helical Angle / Load Angle Hydrotest, X70 Pipe
D = 36 in, t = 12.7 mm D = 36 in, t = 0.5 in, w = 72 in,  = 40
a = 2 mm, 2c = 50 mm a = 2 mm, 2c = 50 mm, h = 0.0 mm
CTD Driving Force (mm)

0.08

CTD Driving Force (mm)

Solid lines: h = 0.00 mm 40/ 50 (HLW) 0.03
Dashed lines: h = 1.27 mm CTD_I
0.06 HLW: Flaw in helical weld
SEW: Flaw in skelp-end weld 40/40 (SEW) CTD_III
0.02
0.04
29/29 (SEW)
0.01
0.02 19/19 (SEW)

0.00 0.00
0.6 0.7 0.8 0.9 1.0 1.1 1.2 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
Applied Stress / Yield Strength Applied Stress / Yield Strength
Figure 12 Model I and Mode III crack driving force (CTDI and
Figure 10 CTD driving force under lowering-in tension (X80) CTDIII) induced by hydrotest
Helical Weld 0.05
Hydrotest, X70 Pipe
D = 36 in, t = 12.7 mm Helical Angle / Load Angle
0.04 a = 2 mm, 2c = 50 mm

CTD Driving Force (mm)

hoop total Solid lines: h = 0.00 mm 19/71 (SEW)
Dashed lines: h = 1.27 mm
0.03 HLW: Flaw in helical weld
29/61 (SEW)
SEW: Flaw in skelp-end weld
axial
0.02 40/50 (SEW)

Flaw 40/ 40 (HLW)

0.01
Skelp-end Weld

Figure 11 Schematic drawing of bi-axial stress under service 0.00

To simulate the hydrotest process, internal pressure was 0.6 0.7 0.8 0.9 1.0 1.1 1.2
applied while the pipe ends were restrained from longitudinal Hoop Stress / Yield Strength
displacement. The hydrotest also applies a mixed mode loading Figure 13 CTD driving force under hydrotest (X70 Pipe)
of Mode I and Mode III to the flaw. As shown in Figure 12, 0.05
unlike in the lowering-in process, the hydrotest generates Hydrotest, X80 Pipe
D = 36 in, t = 12.7 mm Helical Angle / Load Angle
higher Mode I loading than Mode III loading. a = 2 mm, 2c = 50 mm
0.04
CTD Driving Force (mm)

The total CTD driving forces are shown in Figure 13 and Solid lines: h = 0.00 mm
Dashed lines: h = 1.27 mm
Figure 14 for X70 and X80 pipes, respectively. The primary HLW: Flaw in helical weld 19/71 (SEW)
0.03
stress from hydrotest is applied in pipe circumference direction. SEW: Flaw in skelp-end weld
Therefore, the load angle is the angle between the flaw length 29/61 (SEW)
and the pipe circumference. The CTD driving force increases 0.02
as the load angle increases no matter where the flaw is located 40/50 (SEW)

(skelp-end weld or helical weld). But, the weld high-low 0.01

40/ 40 (HLW)
misalignment shows much greater effect on the CTD driving
force than what was found in the lowering-in simulation. The
0.00
weld high-low misalignment increases the CTD driving force 0.6 0.7 0.8 0.9 1.0 1.1 1.2
by approximately 50% in comparison to the cases without high- Hoop Stress / Yield Strength
low misalignment. It is also found that, unlike the lowering-in
Figure 14 CTD driving force under hydrotest (X80 Pipe)
analysis results, the X70 pipes showed higher CTD driving
forces than the X80 pipes. This is probably due to the bi-axial Operation
stress effect and the change of the main loading mode. Similar to the loading conditions in hydrotest, a pipe
In general, the CTD driving force due to hydrotest is much experiences bi-axial tensile stress during normal operations.
smaller than that from the lowering-in process. The maximum One principal stress (hoop) is in the pipe circumferential
CTD driving force of all the cases that were analyzed is much direction which is directly generated by the internal pressure.
smaller than the expected toughness of the materials. Therefore The other principal stress (axial) is along the pipe axial (or
the assumed SEW flaw should be safe under hydrotest. longitudinal) direction and is usually generated by the Poison’s

5 Copyright © 2012 by ASME

effect. The flaw is therefore under mixed mode I (opening / The FEA matrix is shown in Table 3. The pipe OD was 36
normal stress) and III (out-of-plane shear) loading as that in inch and two wall thicknesses, 0.50 and 0.75 inch, were
hydrotest. Since the maximum allowable operating pressure analyzed. The corresponding D/t ratios were 72 and 48,
(MAOP) is less than the hydrotest pressure, the results obtained respectively. Three coil widths, corresponding to helical angles
in the hydrotest analysis can be applied for the operation of 18.6, 28.5, and 39.5, were studied. The effect of the SEW
condition conservatively. high-low misalignment was also analyzed and the maximum
In addition to the axial stress generated by the axial misalignment is assumed to be 1.5 mm.
constraint and Poison’s effect, coincidental forces may apply Table 3 FEA matrix for bending analysis
additional tension or bending load to the pipe in the FEA Matrix - Lower-in induced bending, field bend, and in-service bending
longitudinal direction. It should be noted that only the
Helical
accidental forces that are less than those applied in the pipe WM HAZ OD (D )
Coil Width
Angle
Wall High-
Grade (w) Weld Thickness (t ) D /t Low No. of
( )
lowering-in process are considered in this analysis. OM UM Cases
(inch) (mm) (inch) (mm) (degree) (inch) (mm) (mm)
It is known that the internal pressure is beneficial to the 36 914 18.6 Skelp-
tensile stress capacity (in stress-based design regime). In X70
54 1372 28.5 end
0.50 12.7 72 0.00
1.05 0.95 36 914 weld 32
another word, the resistance to tensile rupture (in stress-based X80 72 1829 39.5 0.75 19.1 48 1.50
Plain
design regime) is increased by the internal pressure. Therefore, N/A
Pipe
the results from the lowering-in analysis can be applied to those Pipe with skelp-end weld
accidental force conditions conservatively. Enforced Rotation Enforced Rotation

Assessment of Compressive Capacity

Assessment Procedures L = 6D

The buckling failure mode is associated with a pipe’s

D
tolerance to global bending. At initial loading stage, a pipe
gradually deforms with the increase of bending moment. The
deformation is relatively uniform along the pipe length. Upon
reaching a critical load level, the deformation starts to localize Skelp-end weld T-Joint at the
Helical weld
in certain locations. Further loading usually causes the compressive side

localized deformation in one of the locations to grow rapidly Figure 15 FE model for bending analysis
and the bending moment to decrease. The curvature or strain Lower-in/Cold Bend
associated with the peak bending moment is often used as the
critical bending strain. In general, buckled pipes with finite The lower-in and cold bend induced bending stress was
size wrinkles can still contain pressure. Therefore, the buckling simulated by applying rotation at both pipe ends. One pipe end
is not considered as an ultimate limit state. However, the was allowed to translate freely in the longitudinal direction.
buckled pipes offer reduced resistance to further deformation A deformed pipe from the lower-in simulation is shown in
and may fail by fatigue in the buckled area. Figure 16 where a buckle was formed on the bottom (i.e.,
The focus of the analysis is whether the material property compression) side of the pipe. Due to the highly localized
variation and weld geometry discontinuity near the SEW, deformation near the buckle, the compressive strain at the
especially at the T-joint, could reduce the tolerance of pipes to bottom of the pipe is not uniform along the pipe length. An
the compressive stress on the compression side of the pipe. average strain is often used to evaluate the compressive strain.
The average strain is calculated using the relative rotations
The critical compressive buckling strain of pipes between the two cross-sections of a pipe segment of a given
containing SEW was calculated in FEA. The calculated critical gauge length centered at the buckling location as shown in
strain was then compared with the maximum strain given by Figure 16. A gauge length of 2D is often used where D
the design codes to see if the critical strain is reduced by the represents the pipe outside diameter. The average strain was
SEW. The overall buckling analysis procedure has been calculated using the following equation,
D ( 2  1 )
extensively documented in [11,12].
FEA Models  ave  where l gauge  2 D (2)
Figure 15 shows a typical FE model used in compressive 2 l gauge
capacity analyses. The pipe was modeled with 3D solid/brick
Typical relationship between the bending moment and the
elements. The total length of the pipe was 6OD which is
average compressive strain from the finite element simulation
similar to the dimension used in most experiments. To simulate
is shown in Figure 17 for X70 pipes. Three pipe conditions are
the thick plates attached to the pipe ends in most experimental
shown: (1) a plain pipe without helical or skelp-end weld; (2) a
bending tests, the pipe ends were modeled as rigid plane. One
pipe with helical and skelp-end weld without weld high-low
of the T-joints was put on the bottom side of the pipe where the
misalignment; (3) a pipe with helical and skelp-end weld with
maximum compressive strain is expected.
1.5-mm SEW high-low misalignment.

6 Copyright © 2012 by ASME

 D = 914.4 mm;  = 28.52; t = 0.75 in; fp = 0; X70
lgauge lgauge 9
2 2
8

7
Enforced Enforced

Moment (MN  m)
Rotation Rotation 6
1 2
5
Bulking Location
4 Plain Pipe
Figure 16 Buckled pipe under bending deformation Skelp-End Weld, h = 0 mm
3 Skelp-End Weld, h = 1.5 mm
The average compressive strains corresponding to the
maximum bending moment are normally referred to as the 2
critical compressive strains. It is seen that without high-low 1 X70, No pressure, = 29, D/t = 48
misalignment, the SEW does not affect the critical strain. The
0
high-low misalignment, however, can greatly reduce the critical 0 0.5 1 1.5 2 2.5 3 3.5 4
strain. The critical strains of the pipes with different helical Compressive Strain (%)
angles are almost identical. In other words, the critical Figure 17 Moment vs. compressive
No Pressure strain (no pressure)
compressive strains were found to be independent of the helical 6
angles. Similar results were found for X80 pipes. Elastic Equation
CAN/CSA Z662-96
The FEA calculated critical strains of the X70 pipes (D/t = 5 DNV-OS-F101
48 and 72) are compared with the maximum design strains C-FER Zero Pressure Design
FEA: X70, No HiLow
from some pipeline design standards in Figure 18. Without

Critical Strain (%)

FEA: X70, HiLow = 1.5 mm
4
high-low misalignment, the calculated critical strains for pipes
X70 No Pressure
containing SEWs are all greater than the maximum design
3
strains allowed by the codes. However, the calculated critical
strain with 1.5-mm high-low misalignment and D/t = 72 is
2
slightly lower than that given in CSA Z662 but is still higher
than those given in DNV-OS-F101 and C-FER design curve. It
should be noted that the elastic equation given in Figure 18 is 1
an ideal elastic solution to the buckling problem. It is known to
over-predict the buckling strain and is not taken by any of the 0
20 40 60 80 100
design codes or standards. The elastic equation is presented D/t
only for reference and validation purpose.
Figure 18 Calculated critical compressive strain vs. design
Operation strains in codes
D = 914.4 mm;  = 28.52; t = 0.75 in; fp = 0.8; X70
Similar bending analyses as those described in previous 6
section were performed with the addition of internal pressure.
The finite element model and matrix were the same with those 5
shown in Figure 15 and Table 3. In the analysis, the MAOP
was applied first where the pipe end was free of any
Moment (MN  m)

4
displacement restraint and load. The bending load was then
applied by enforcing a specified rotation at the pipe end. 3
Typical bending moment vs. average compressive strain curves Plain Pipe
are shown in Figure 19. The results were found almost 2
Skelp-End Weld, h = 0 mm
Skelp-End Weld, h = 1.5 mm
identical for all helical angles and for X80 pipes. By
comparing these results with those in Figure 17, it is seen that
1
the internal pressure increases the critical compressive strains.
X70, Pressure = 80%SMYS,  = 29, D/t = 48
Therefore, the strain capacity is sufficient when compared with
0
the maximum design strains shown in Figure 18. 0 0.5 1 1.5 2 2.5 3 3.5 4
Compressive Strain (%)
From the above analysis, it can be concluded that the SEW
does not pose any threat to the pipe integrity under static Figure 19 Moment vs. compressive strain (with pressure)
service loads. Fatigue Assessment
Pressure and temperature oscillations during normal
operations can generate cyclic stress in a pipe. As a result, the
flaws in the skelp-end weld may undergo fatigue growth. The
cyclic stress consists of two principal components, i.e., hoop

7 Copyright © 2012 by ASME

and axial (as shown in Figure 11). The flaw is under mixed should not be a concern for the skelp-end welds with initial
mode I and III fatigue loading conditions. The magnitude of flaws less than 2 mm (depth)  50 mm (length).
each loading mode depends on the angle between the total Table 4 Operating conditions
stress (total) and the flaw surface (i.e., helical angle).
Pipe and Operating Conditions
In the following analysis, the mixed mode fatigue problem OD WT Grade Design MAOP hoop axial total
was simplified to a pure mode I problem where the applied (in) (mm) (ksi) Factor (ksi) (ksi) (ksi) (ksi)

stress was assumed to be the total and always perpendicular to 36.0 12.7 80 0.8 1.78 64.0 19.2 66.8

the flaw surface regardless of the helical angle. Therefore, this Table 5 Stress spectrum for fatigue assessment
simplified assessment is independent of the helical angles. The
assessment results are believed to be conservative since the Stress Spectrum for Fatigue Assessment

materials’ resistance to mode I fracture and fatigue are usually Description P/P
P hoop axial total Cycles/
(ksi) (ksi) Year
the lowest among all three loading modes. (ksi) (ksi)
Daily 2% 0.04 1.28 0.38 1.34 365
The pipe geometry and designed operating condition are Monthly 10% 0.18 6.40 1.92 6.68 12
given in Table 4. The MAOP yields a design factor of 0.8. The Seasonal 20% 0.36 12.80 3.84 13.36 2
maximum allowed flaw size under the MAOP was determined Upset Condition 50% 0.89 32.00 9.60 33.41 4
with the ECA procedure in API1104 Appendix A where the Hydrotesting 125% 2.22 80.00 24.00 83.52 N/A

Option 1 method was used with the CTOD toughness between P: Operation pressure; P: Pressure oscillation

0.10 mm and 0.25 mm. Table 6 Predicted flaw growth vs. service time
The assumed stress spectrum used in the fatigue
BS7910 Mean+2SD Growth Curve R>0.5 in Air
assessment is given in Table 5 where the oscillation of the
pressure was given as a percentage of the MAOP. It should be No High-Low High-Low = 10% Wall Thickness
noted that only one hydrotest cycle was applied. The Length Depth Length Depth
Year Year
combination of the high pipe grade (X80) and the high MAOP (mm) (mm) (mm) (mm)
(Design Factor = 0.80) yields conservative results since it 50.0 2.0 0 50.0 2.0 0
generates high cyclic stress. The assessment results can be 51.6 2.7 167 51.6 2.7 73
conservatively applied to lower grade pipes (such as X70) and 53.2 3.4 281 53.3 3.4 124
lower design factors (such as 0.72). In addition, the assessment 54.9 4.1 357 55.0 4.2 158
results can be conservatively applied to pipes with wall 56.6 4.9 410 56.6 4.9 181
thickness greater than 12.7 mm. CONCLUSIONS
The fatigue growth rate (da/dN) was assumed to follow the Fitness-for-service analysis was conducted on SEWs under
Paris’s Law and be in the form of a power law function of the a variety of loading conditions including construction,
stress intensity factor range (K) as da/dN = C (K)n. The two- commissioning, and full-pressure service. The FFS analyses
stage mean plus two standard deviation (2SD) growth covered the stresses from lower-in, cold bending, hydrotest, and
properties for R  0.5 in air recommended in BS 7910:1999 for operation. The tensile, compressive, and fatigue failure modes
weld fatigue assessment were used. The constants were given were investigated.
in the following,
In comparison to helical welds, no inherent risk was found
Stage A (K  196 N/mm3/2), C = 2.110-17 and n = 5.10, for SEWs. Using material property data from actual SEWs, the
Stage B (K > 196 N/mm3/2), C = 1.2910-12 and n = 2.88, fitness-for-service analysis has demonstrated that the pipes
containing SEWs are safe under both static and cyclic loading
where the unit of da/dN is mm/cycle.
conditions. The project team believes spiral pipes containing
The initial flaw depth and length were assumed to be 2 mm SEW can provide satisfactory service when manufactured,
and 50 mm, respectively; and the weld high-low misalignment inspected, and accepted by the recommended procedures.
was assumed to be 0.00 mm or 1.27 mm (10% of wall
The fitness-for-service of the SEWs is established on the
thickness). The stress intensity factor solutions of Anderson’s
basis of (1) safety margins under a variety of loading conditions
[13] were used in the assessment. A stress magnification factor
and (2) comparison of expected performance of spiral welds
of 1.33 was applied to the case with weld high-low
and SEWs under similar material property and loading
misalignment per BS7910:1999.
conditions. The actual pipe properties in a given pipe order or a
The predicted flaw growth vs. the service time is shown in grade can have large variations permitted by applicable codes
Table 6, where the last row indicates the predicted maximum and standards (e.g., API 5L and CSA Z245.1). Anti-corrosion
allowable flaw size at the MAOP from the ECA procedure in coating is known to affect the tensile property of the pipes. The
API1104 Appendix A. It is seen that the predicted service life Charpy impact energy of the welds (deposited weld metal and
is 410 and 181 years for the cases without and with high-low heat-affected zone) is likely to have variations too. These
misalignment, respectively. Therefore, the fatigue failure variations are known and they affect both spiral welds and
SEWs. The material properties used in the fitness-for-service

8 Copyright © 2012 by ASME

analysis should not be interpreted as representing a particular
grade, wall thickness, or other attributes of the data used. With Evaluation against Experimental Data,” Proceedings of the
the possible material property variations in mind, the safety 9th International Pipeline Conference, Paper No. IPC2012-
margins shown in the analysis are sufficiently high that 90660, September 24-28, 2012, Calgary, Alberta, Canada.
satisfactory service of SEWs can be expected when the 11 Liu, M. and Wang, Y.-Y., “Modeling of Anisotropy of
recommended QA/QC procedures [1,2] are applied. TMCP and UOE Linepipes,” Proceedings of the Sixteenth
ACKNOWLEDGMENTS International Offshore and Polar Engineering Conference,
The financial support from the following sponsors is (ISOPE 2006), San Francisco, USA, May 28 – June 2, 2006,
gratefully acknowledged: Alliance Pipeline, Berg Steel Pipe, pp. 221-227.
Corinth Pipe Works, El Paso, Enbridge, Evraz Inc. N.A., 12 Liu, M. and Wang, Y.-Y., “Modeling of Anisotropy of
Panhandle Energy, PSL N.A., Spectra Energy, Stupp, and TMCP and UOE Linepipes,” International Journal of
TransCanada Pipeline Ltd.
Offshore and Polar Engineering Conference, Vol. 17, No. 4,
December 2007, pp. 288-293.
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13 Anderson, T.L., Thorwald, G., Revelle, D.J., and Lanaud,
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4 CSA Z662, 2007.
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Tier Tensile Strain Models for Strain-Based Design – Model

9 Copyright © 2012 by ASME