Selection of a_completion_material_for_a

Paper No.
111
SELECTION OF A COMPLETION MATERIAL FOR A SOUR OIL FIELD USING FRACTURE
MECHANICS CONCEPTS AND STRESS CORROSION CRACKING TESTING
Albert0 Valdes and Raymundo Case
Intevep, S.A. - Research and Technological Support Center of Petr6leos de Venezuela, S.A.
PO Box 76343, Caracas 107OA, Venezuela
ABSTRACT
This paper describes the methodology used and highlights the experimental results that led to the choice of
the completion material for an oilfield expansion with a large variation in the concentration of aggressive
gases (both sweet and sour), chloride content, temperature, pressure and water content. The testing
program was formed by corrosion (general, localized and DCB), metallurgical and mechanical tests.
Finally, Fracture Mechanics concepts and calculations were used to determine the maximum defect size
that could be allowed for a safe operation of the new wells.
Keywords: SSCC, High strength steels, Fracture Mechanics, Sour Service, Chrome steels, CO2, H2.S
INTRODUCTION
In oil and gas production, corrosion related failures can make a significant impact over the costs involved
in the actual exploitation of a given reservoir, particularly when a sour corrosive environment is present
due to the amount of H2S and CO2 dissolved in the water produced or condensed. Theses conditions will
lead to pitting, crevice, high general corrosion, Sulfide Stress Corrosion Cracking (SSCC) or Stress
Corrosion Cracking. The relative severity of each of these forms of environmentally related damage will
depend on the actual or expected conditions present during operation of the well.
A cost effective material selection methodology for OCTG grade steels, is critical for an economically
efficient exploitation of the reservoir. The selection criteria for OCI’G grade steels are guided by the
mechanical properties, resistance to stress cracking and to general and localized corrosion at the expected
operating conditions. However the lack of a complete knowledge of the effect of the environmental
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parameters on these aspects may lead to the use of corrosion resistant alloys that are outperforming or
more expensive then necessary.
This situation underscores the importance of developing a suitable OCTG selection methodology for high
strength steels, that can take into account the effect of the entire set of environmental parameters defining a
sour corrosive media. This way a cost effective selection can be made for each particular condition found
in the field. The selection of a corrosion resistant alloy (CRA) is particularly delicate because of the high
investment costs associated. Additionally, the susceptibility of several CRA alloys towards SSCC at near
ambient temperatures, and common stress corrosion cracking at higher temperatures associated with
localized corrosion, further complicates the alloy selection. This behavior shown by CFU alloys is related
to the partial pressures of CO* / H$, chloride concentration and temperature as well as to the content of
specific alloying elements [l-3].
In order to characterize in a quantitative manner the effect of the major environmental parameters on the
susceptibility to cracking in several CFL4 alloys suitable testing protocols are needed, which takes into
account the particular conditions for which these materials were designed. The results obtained with this
methodology must be analyzed taking into consideration concepts and calculations from Fracture
Mechanics [4]. However, these procedures can not give a complete picture about the behavior of the
material if sensible aspects such as pitting resistance, fracture toughness and local stress requirements are
not properly assessed.
Environmentally Assisted Cracking Assessment using Fracture Mechanics
Fracture mechanics has been used to describe the mechanical parameters that govern crack initiation and
growth. These concepts have been applied to the study of environmental assisted cracking (EAC) and
particularly to SSCC in high strength steels, mainly by defining a threshold stress concentration factor
through which the critical conditions for steady state crack growth can be assessd [4].
Several test methods have been published elsewhere to determine the stress corrosion cracking behavior
using Fracture Mechanics. The specimens utilized can be classified in KI decreasing or KI increasing with
crack growth [4]. The most common techniques use specimens designed for Mode I cracking, where the
crack is tension loaded.
The main feature of the KI decreasing specimens is that the stress corrosion crack continues to grow as
long as there is enough free elastic deformation energy ahead of the crack tip. Therefore the threshold KI is
achieved when the crack growth stops. In the study of SSCC, the most common kind of specimen used is
the DCB type standardized by the NACE International specimen which is of the KI decreasing type [5].
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The DCB specimens as a testing method for assessing the susceptibility to SSCC of high strength OCTG
carbon steels have been described in several publications in the literature [6-81. It has been used either as a
complimentary methodology that can provide useful information for specific situations, or as a mean to
define in a quantitative way the susceptibility to SSCC using the threshold stress concentration or Klsscc
parameter.
The usefulness of the threshold stress concentration factor to assess the susceptibility to particular forms of
cracking, such as SSCC, is based on the relationship between the increase of environmental severity to
produce SSCC and the decrease of the Ktssc- parameter value [6,8,9]. The results reported in the literature
clearly show a much higher sensibility of the Krsscc parameter when compared to other common methods
used to assess SSCC susceptibility, such as slow strain rate or constant load [7,10].
This behavior shown by the KIss~ parameter with the relative severity of an environment, defined by
solution pH, temperature and content of acid gases, become useful to develop material testing protocols
that take into account the actual or expected environmental conditions at given well or reservoir [7,8].
Damage Tolerance Analysis Assessment
Implicit in the concept of the stress concentration factor is the relationship between the applied stress to a
given structure and the expected defect dimension, as expressed in the following equation :
Where K is the stress concentration factor, Q is the remote applied stress, a is the defect depth and f a
shape factor that depends of the geometrical characteristic of the structure where the crack is present.
Given a particular value of the stress concentration factor and mechanical stresses applied to a particular
structure, it is possible to assess the necessary defect dimensions that will produce such K value.
This concept, known as damage tolerance analysis, is widely applied in fracture mechanics to assess the
defect dimensions that could lead to damage in a structure. For assessing environmental cracking the most
important curves are associated with the threshold stress concentration, Krm and the fracture toughness or
Krc . The first curve defines the conditions required for the onset of cracking and the last one for fracture
itself.
Figure 1 depicts a typical damage tolerance set of curves. In this figure the lower bound curve is related to
the required defect depth and length that will initiate a crack. The upper bound curve is related to the
fracture toughness and shows the maximum allowable crack dimension that will produce an unstable
propagation or fracture. In the space between both curves the cracks propagate at a constant rate. Therefore
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the useful life of a structural element subject to stress corrosion cracking will be controlled by either the
actual crack growth rate and the difference between Krrn and Krc values of the material.
The area under the curve related to Krm is controlled by the resistance of the material to localized corrosion
that will produce or conduce to stress corrosion cracking. For CRA alloys the aspect of most concern is
pitting or crevice resistance, as well as the SSCC resistance itself [2,3].
This paper describes a test protocol developed for obtaining the necessary information to draw the damage
tolerance curves for a set of CRA alloys under actual environmental. This analysis was used as the main
criteria for selecting the most adequate material that will lead to a cost effective solution with maximum
reliability.
EXPERIMENTAL
The test protocol was applied to three CRA’s: two different modified 13Cr alloys and one 15Cr alloy,
being considered as candidate completion materials of new wells (3.5”, 12.7 #), as a replacement of the
22Cr used previously in the same oilfield. The 22Cr alloy was included in the study for comparison. The
purpose of the tests was fourfold: (1) to establish the resistance of the materials to several forms of
corrosion damage under the actual environmental conditions where they would be used, (2) to obtain
relevant data for a damage tolerance analysis and, (3) to draw guidelines for a purchase specification
document and, (4) to define corrosion performance related parameters for the manufacturing quality
control program.
Typical range of environmental conditions, corresponding to the South Lake Oilfield in western
Venezuela, are shown in Table 1. Depending on the selection criteria, the choice of a CRA for these
conditions could balance from a borderline performance of the alloy to an over design in material selection.
Both being inappropriate, the former because of safety considerations, the latter because of the higher
costs involved in several aspects of the completion. Damage tolerance curves were used to determine
which alloy could be safely used.
Design
The wide range of operating conditions of the South Lake field provides ground for several corrosion
problems. The main emphasis was made on the susceptibility to SSCC, localized and general corrosion.
In order to optimize the number of tests a set of four experimental conditions was chosen (see Table 2).
The test parameters for assessing the susceptibility to SSCC were selected for the shut-in wellhead because
this is expected to be the most severe condition for sulfide induced cracking. Similarly, the test parameters
for resistance to localized and general corrosion were those of the flowing wellhead when these forms of
111/4

corrosion should be important due to the higher local temperature. All tests were performed at two H2S
and chloride concentration levels.
Materials Description
The nominal compositions of the alloys tested is shown in Table 3. All steels tested had a quenched and
temper uniform martensite matrix with no carbides present. The mechanical properties of these materials
are shown in Table 4. All steels had a yield strength in the 95 Ksi range, except the 22Cr Duplex alloy
which classified as a 110 Ksi material. The Krc values shown in Table 4 were obtained from CTOD tests
published elsewhere [111.
Testing Methodology and set-up
The test protocol was based on several standard techniques used to assess the SSCC susceptibility, as well
the resistance to general and localized. To obtain the high temperature and pressure required in the
experimental design several autoclaves were adapted to perform the DCB, localized and crevice corrosion
specimens.
Figure 4 shows the autoclave setup used in the testing protocol, which consist of a stirred Hastelloy C
vessel with 2 liters of internal volume, temperature control and gas mixture supply.
All experiments were run in duplicate (DCB) or triplicate (general and localized) tests. For the DCB tests
only two specimens were placed in each autoclave in order to keep the exposed surface area to solution
volume ratio below the value of lo-15 ml/cm3 suggested in the TMO177-96 standard. For the general and
localized corrosion tests the same criterion was maintained. Results are reported as an average in each
case.
SSCC assessment
The susceptibility to SSCC was determined by means of the DCB technique, which is widely described in
the literature. A sub-size type of experiment was used, the most important dimensions of which are
shown in Table 5. These dimensions were chosen according to NACE TM0177-96 standard [12]. All
DCB specimens were machined from actual tubing samples.
In order to obtain a reliable KIsscc value using the DCB specimens it is necessary to set an appropriate
crack opening displacement and to establish an adequate test time. The crack opening displacement for
22Cr (alloy A) was taken from NACE TM0177-96 standard, whereas that of the other alloys (B-D) was
determined by an in-house developed procedure.
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Results obtained by Altem Current Potential Drop (ACPD) [7] indicate that for a high strength (P-l 10)
alloy an stable crack growth requires more than 720 hr. to show. Thus, separate tests were performed at
720 hr. and 1440 hr. in order to establish the actual crack detention required to assess a reliable KIsscc
value. All DCB specimen were fatigue precracked until obtaining an initial pre-crack length of 1 mm. The
calculation of the Ktsscc values was made using the procedure described in the NACE TM0177-96
standard [121.
Localized corrosion resistance
Two forms of localized corrosion were considered, pitting and crevice, due to its relation with the integrity
of the tubing wellhead. Figure 2 depicts the typical test coupon and its dimensions used to assess the
pitting and crevice resistance. Each specimen was ground up to 1 pm, to simulate actual mill surface. The
presence of pitting was verified by visual inspection at 100X. The diameter distribution was measured
counting the observable pitting population at this magnification with an image analyzer adapted to an
optical microscope.
The experimental setup consists of two test coupons joined face to face by means of a screw and bolt,
isolated from the sample with PFTE washers. Figure also shows the dimensions of the test coupons used.
The coupons for crevice resistance had the same surface treatment of those for pitting resistance.
General corrosion resistance
The general corrosion rate was measured by the weight loss method, according to ASTM G-l standard
[131. Figure 3 shows the dimensions of these coupons. Each specimen received the same surface treatment
as in the localized corrosion test.
Methodology of Analysis
As mentioned before, the aim of the present study was to determine the most convenient CRA alloy for the
operating conditions found in the South Lake field, using a fracture mechanics approach based on damage
tolerance. In order to estimate the parameters for the damage tolerance calculations a stress analysis to the
tubing using the actual completion design specified for the South Lake oilfield was carried out. The
damage tolerance curves were calculated with a software package developed for Fitness for Service
analysis for pressure vessels [ref.], based on the mechanical well completion design.
The damage tolerance curves were used to obtain the limiting flaw size allowed by the fracture toughness
(Krc) and the threshold stress concentration (Krss& of each material. From these curves and the results of
the localized corrosion tests the estimated lifetime of the well tubing was estimated.
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EXPERIMENTAL RESULTS
Susceptibility to SSCC
The results of the DCB tests are shown in Table 6. Steel C showed around 6 mm of crack propagation
under condition 1 corresponding to a Krssc~ value of 34.8 MPadm, and 5 mm at condition 2
corresponding to a Krsscc value of 35.8 MPadm. These values for steel C where obtained at 1440 hr. of
test time, which corresponds to an average crack growth rate of 6.94~10.~ nun/h and 8.94~10.~ for
conditions 1 and 2 respectively. Steels A, B and D showed no crack growth in any of the tests and
therefore these tests were classified as invalid according to NACE TM0177-96 standard. The applied
stress intensity factor for these steels is shown instead in Table 6.
Complementary tests were performed by the steel manufacturers in order to verify the susceptibility to
SSCC by means of tensile constant load tests under the same conditions used for the DCB tests. These
tests were performed at 90% of SMYS for an exposure time of 720 hr. The examination for cracking was
made visually at 100X. The results obtained are shown in Table 7. No cracks or failures were obtained for
steels A, B and C. However, steel D showed cracking at condition 2. This showed that the crack opening
displacement applied for steel D was not enough to reach the onset of crack propagation in the DCB tests.
Independent tests were performed on steel B with a modified condition 2, i.e. increasing the partial
pressure of H2S to 2.62 KPa (0.38 psi), using DCB and constant load test methodologies [14]. The
results showed crack propagation at 720 hr. and 1140 hr. of exposure time, giving a KIsscc of 26 MPadm,
and exhibiting failure in the constant load test at 90% of the SMYS. The average crack growth was 5.2
mm in the 720 hr tests, resulting a calculated propagation rate of 7.2~10~~mrn/hr.
Localized Corrosion Tests
The results obtained in the pitting and crevice tests are shown in Table 8. The results of the pitting tests
show that only steel A is resistant to both conditions. Steels B and C were susceptible to pitting corrosion
at the highest H$S partial pressure with steel B being the more resistant of the two. Steel D showed pits
under both conditions tested.
The pit diameter distribution for steel B is shown in Figure 5. Pit diameters spread from 20 pm to 180 pm,
with the average around 80 pm. From the average pit diameter the penetration rate was calculated as 1.45
mm/y (58 mpy). Figure 6 shows the pit diameter distribution for steel C, the average pit diameter is 1.5
mm, which gives an estimated penetration rate of 8.6 mmlyr. (344 mpy). Figure 7 shows the pit diameter
distribution for steel D at condition 3 which has the highest H$ partial pressure, the average size is 0.15
mm resulting in a penetration rate of 1.9 mm/yr. (76.8 mpy).
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The results of the crevice tests (see Table 8) indicate that steels C and D show damage under the testing
condition 3. Steels A and B did not show evidence of crevice corrosion susceptibility. These results were
used only as a qualitative indication.
General Corrosion Tests
The results of the general corrosion tests are shown in Table 9. Corrosion rates for steels A, B and C were
in all instances below 0.01 mm/yr. (less than 0.4 mpy) corresponding to a high corrosion resistance. Steel
D showed the highest corrosion resistance of the group for both conditions.
DISCUSSION
The damage tolerance curves obtained from applying the methodology, are shown in Figures 7 through
10. Each curve indicates the relationship between the flaw dimensions (width and depth) required to
produce crack growth under the test conditions. The upper bound indicates the crack size that will produce
fracture. In order to simplify the present discussion, the damage tolerance curves obtained for each CRA
steel evaluated will be discussed separated.
Since for several of the steels it has been not possible to assess a reliable value for Krsscc, an empirical
value of 30 Mpa m*0.5 will be used for reference purposes in the damage tolerance curve analysis.
Actually this particular value of fracture toughness seems to be associated with a threshold value of
susceptibility to SSCC in high strength carbon steel [6,7, lo].
The use of the 30 MPadm value in the analysis of damage tolerance curves of the CBA alloys evaluated
resides in the comparison between the DCB and tensile tests performed within the testing protocol. In the
different performed, it has been verified that when the Krsscc value was above 30 MPadm, the tensile test
showed no cracking or failure [7, lo]. Also it was observed that if Ktsscc value measured is below the
reference value failure was verified in the tensile tests.
Therefore, although the threshold value of 30 MPadm is actually experimentally evaluated for high
strength carbon steels, it’s employed in the present work as reference to indicate a lower boundary to
simplify the analysis of the damage tolerance curves.
Behavior of Steel A
The damage tolerance curves obtained for the steel A are shown in Figure 7. The curve at 30 MPadm
represents a lower damage tolerance limit above which crack growth may be expected, whereas the curve
at 315 MPadm sets an upper boundary beyond which catastrophic cracking may happen. For steel A,
neither the DCB nor the tensile tests yielded evidence of susceptibility to SSCC. This leads to think that
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this steel in immune to attack under the conditions tested. Taking into account that pitting may occur under
actual field conditions, a threshold value of 30 MPadm was used as a conservative estimate for the lower
defect tolerance curve for steel A.
The lower and upper tolerance defect limits are very separated, this indicates that in the event of apparition
of a crack initiating defect the onset of fracture will occur when the crack depth is almost equal to the wall
thickness of a typical 3.5” (12.7 #) tubing string or, tire crack length would be comparable to the internal
perimeter . This indicates the likeliness that the performance of steel A be “leak before break”. Also, the
results of the localized corrosion tests showed that there is no evidence of pitting or crevice corrosion
under the experimental conditions. The overall judgment is that steel A satisfy the requirements of
resistance to environmental degradation. This steel, however, was retained in the study for performance
comparison since it is not an economic alternative.
Steel B
The damage tolerance curve for steel B is shown in Figure 9. Since the tests results allowed to calculate an
actual value of Ktsscc and there was presence of pitting under the two test conditions, two lower boundary
curves were generated corresponding each one to a test condition.
The lower damage curve correspond to a Krsscc value of 26 MPadm obtained in the lower chloride
environment. The 35 MPadm shows the estimated lower boundary for environmental assisted cracking
that could be caused by the pitting found on the high chloride tests. The calculations leading to Figure 9
indicate that defects in the range 2 - 4 mm in depth can produce a crack.
The estimated Time-to-failure (TTF) of the tubulars for the 26 MPadm curve was assessed from the
calculated crack growth rate and pit penetration rate. For a closed well head condition the ‘ITF was
calculated to be 1.7 years. This ‘ITF requires an initial defect length of 0.8 mm which is unlikely to happen
in view of the results of the pitting tests which have shown an average pit depth one order of magnitude
smaller. The estimated ‘RF of the tubulars for the 3.5 MPadm curve is 2.8 years. As previously, this
condition is quite seldom to occur because it will require an initial pit depth of 1.5 mm, which is believed
unlikely to occur.
Steel C
The set of damage tolerance curves associated with steel C is shown in Figure 10. Again two curves are
plotted for the lower limit, both corresponding to the Krsscc values obtained in the tests at 1440 hr. Both
KIsscc values are similar and the curves tends to overlap.
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For the lower curve corresponding to 34.8 MPadm the estimated TTF is 0.2 yr., based on the crack
propagation and pit penetration rate measured in the tests. The required defect size to initiate the crack
growth is 0.07 mm in diameter which is below the average diameter measured in the tests. Therefore, it
can be inferred that the cracking produced under this condition could be related to SCC.
The lower curve corresponding to 35.8 MPadm, yields an estimated time to fracture of 0.2 years, similar
to the previous condition. However, since no pitting was observed under the corresponding tests this TTF
can not be associated to a pit penetration rate. This indicates that steel C could be susceptible to SCC if a
defect would form with the necessary dimensions.
Steel D
The damage tolerance curves obtained for the steel D are shown in Figure 11. The tensile tests performed
on this steel have shown that it is susceptible to SSCC, although no reliable value Kisscc was obtained.
Similarly to steel A the reference curve calculated at 30 MPadm was used for calculating the lower damage
tolerance limit.
The ITF was estimated from crack growth rate reported in the literature for 410 steel [151, which has
composition and microestructural features similar to steel D. A value of 7~10~~mm/s was calculated for the
crack growth of steel D.
Under these premises the estimated life time of the tubing was estimated in 1 year. Most of this time would
be consumed in pit growth to reach the required size to produce a crack. The calculated required pit
diameter is 1.7 mm, which is close to the average value recorded in the pitting test of steel D under
conditions 2 and 3.
General Comparison
Based on the damage tolerance curve analysis applied to the different steels a quantitative approach to an
adequate selection can be made. In this particular case the aspect of reliability can be expressed in terms of
the longest time to failure.
Steel A exhibits the most suitable behavior since it does not show any specific evidence of SSCC
susceptibility and shows no localized corrosion damage. The corresponding lower defect tolerance limit is
only referential of the worst estimated possible environmental condition that could conduce to EAC. In this
case the high fracture toughness values of steel A allow a leak before break behavior.
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Steel B is prone to SSCC, however, the required defect sizes that could lead to stable crack propagation is
well over the maximum pit size distribution found in the tests. Under these circumstances it is believed that
a failure is unlikely to occur.
Steels C and D, show evidence of susceptibility to EAC and SSCC respectively. Also, the defect
dimensions are within the measured pit size distribution. Under these conditions it is likely that a failure
will occur during service, reducing drastically the useful life of the well completion.
Finally, using the analysis of the damage tolerance curve the final decision made was to use either steel A
or B under the operational conditions of the South Lake Field. This recommendation was finally expressed
as a purchase specification for the tubing string and as quality control parameters for the manufacturing
process.
CONCLUSIONS
1. A test protocol was developed and used to choose the most suitable CBA steels for the operational
conditions of the South Lake Oilfield. The protocol consists of DCB and Tensile tests to define either
EAC or SSCC susceptibility, and also general and localized corrosion tests. The final decision on the
CBA alloy to be used was based on suitability, as well as economic criteria (not covered in this paper).
The analysis of the data, using fracture mechanics criteria, led to the construction of damage tolerance
curves.
2. Under the conditions tested, the most suitable steels were found to be steel A and B. This decision was
based on the performance of these steels under the different tests performed, and the resulting damage
tolerance curves, which show a very high resistance of these steels to environmental degradation.
3. The use of the damage tolerance curve analysis allowed to assess the conditions under which a given
steel could be used safely. This analysis allowed to discard steels C and D as not suitable since the
resulting pit size distribution is within the minimum defect size that will produce either SSCC or EAC
under the expected operating conditions.
4. The damage tolerance curve analysis proved to be a suitable method to combine the results of specific
tests (SSCC, general and localized corrosion) in order to assess the conditions that could lead to
fracture or stable crack propagation in the CBA steels.
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1.
2.
3.
4.
5.
6.
7.
8.
9.
10.
11. Intevep, S.A. Internal report
A. Ikeda, T. Kudo, Y. Okada, S. Mukai, F. Terasaki. “Corrosion Behaviors of High Alloy Oil
Country Tubular Goods for Deep Sour Gas Well”. CORROSION / 84, Paper No 206, (Houston,
TX: NACE International 1984).
Y. Miyata, Y. Yamante, 0. Fumkirni, H. Niwa, K. Tar&i. “ Corrosion of new 13 Cr Martensitic
Steel OCTG in Severe CO2 Environment”. CORROSION 95, Paper No 83. , (Houston TX: NACE
International 1995).
M. Ueda, T. Kushida, T. Mori. “ Evaluation of SSC Resistance on Super 13 Cr Stainless Steel in
Sour Applications”. CORROSION 95, Paper No 80, (Houston TX: NACE International 1995).
H.L. Ewalds, R.J.H. Wanhill.. Fracture Mechanics , Edward Arnold (Ed.). Londres, Delftse
Uitgevers Maatschappij, 1985, p.p. 193-205.
B.F. BROWN Stress Corrosion Cracking & Corrosion Fatigue of High Strength Steels, DMICS
Report 210,26-28. 1964.
M. Kermani, R. Macguish, J. Smith, R. Case, J. Vera. ‘The effect of Environmental Variables on
Crack Propagation of Carbon Steels in Sour Media”. Oil and Gas Production and Refining,
International Corrosion Conference (ICC), Houston 1993.
R. Case, J. Vera, C. Sequera, “ The Use of ACPD Technique in Assessing Cracking Propagation
and Kisscc in High Strength Steels OCTG Materials in Sour Media”. CORROSION 97, Paper No
97046. (Houston TX: NACE International 1997).
R. Case, J. Vera, A. Castro, “ The Relationship Between Hydrogen Permeation and Sulfide Stress
Cracking Susceptibilty of OCTG Materials at Different Temperatures and pH Values”. CORROSION
97, Paper 97047, (Houston TX: NACE International 1997).
M. Kermani, D. Harrop, R. MacGuish, J. Vera. “Sulfide Stress Cracking of Downhole Tubular
Steels”. CORROSION 9 1, 199 1, Paper No 272. (Houston TX: NACE International 199 1)
R. Case, J. Vera, A. Viloria, M. Staia, “ Effect of the Environment on the Sulfide Stress Cracking in
High Strength Steels by Fracture Mechanics” (in Spanish). 1” NACE Latin American Corrosion
Conference, 2 (1994), paper 94100.
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12.
13.
14.
15.
NATIONAL ASSOCIATION OF CORROSION ENGINEERS, Laboratory Testing of Metals for
Resistance to Sulfide Stress Cracking in H2S Environments. TM0177-90. 1990.
AMERICAN SOCIETY FOR TESTING MATERIALS, Practice for preparing, Cleaning and
Evaluating Corrosion Tests Specimens, Standard Gl.
Private Communication of the steel manufacturers
J. Yu, R. Brook, R. Hutchings and A. Tumbull. “Stress Cosrrosion Testing of AISI 410 Steel in
Sour Environments using Side Grooved DCB Specimens” in Proceedings of Life Prediction of
Corrodible Structures. Paper # 46, Cambridge, 1991.
TABLE 1
RANGE OF OPERATING CONDITIONS OF THE SOUTH LAKE OILFIELD
Agent Level
Co2 max. 800 psi
H2S 0.16 - 0.56 psi
Chloride 100 -
ppm 30000 ppm
Temperature 77 F (25 C) - 330 F (166 C)
TABLE 2
EXPERIMENTAL DESIGN
Test condition Method Test time Parameter assessed
1
Temperature 25 C
DCB test 336,720 and KISSCC
0.56 psi H2S, 800 psi CO2
100 ppm chloride WA’=@
1440 hrs SSCC susceptibilty
2
Temperature 25 C DCB test
336,720 and Ksscc
0.16 psi H2S, 800 psi CO2
30000 ppm chloride
(NACE D)
1440 hrs
SSCC susceptibilty
3
Temperature 166 C Pitting, crevice 336 hrs
0.56 psi H2S, 800 psi CO2 and weight loss
Resistance to general
and localized corrosion
30000 ppm chloride
4
Temperature 166 C Pitting, crevice 336 hrs
0.16 psi H2S, 800 psi CO2 and weight loss
Resistance to general
and localized corrosion
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TABLE 3
NOMINAL COMPOSITION OF THE STEELS
Steel % C % Si %Mn %P %S % Cr % Ni %Mo %N
A 0.015 0.44 1.58 0.024 0.001 21.96 5.57 2.93 0.13
B 0.010 0.21 0.43 0.017 0.001 11.91 5.73 1.95 -
C 0.12 0.14 0.86 0.013 0.001 14.6 1.54 0.52 0.08
D 0.025 0.25 0.46 0.013 0.001 13.1 4.0 1.0 0.049
TABLE 4
NOMINAL MECHANICAL PROPERTIES OF THE STEELS
Steel
A
B
C
D
Yield strength Tensile strength Elongation (%) Average Fracture
WV WV hardness (HRC) toughness
(MPadm)
814.3 896.3 26.8 28.3 315.3
712.2 814.3 27.2 24.8 271.0
71105 861.8 23.2 23.0 183.0
760.0 838.1 21.0 24.5 304.0
TABLE 5
DIMENSIONS OF THE DCB SPECIMENS
steel
Width Height Length Wedge Thickness Crack Opening
b-d (mm) b-4 Bn (=) displacement
(mm)
A 5 25 100 2.90 0.51
B 5 25 100 2.88 0.71
C 5 25 100 2.87 0.52
D 5 25 100 2.90 0.33
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TABLE 6
RESULTS OF THE DCB TESTS
Material Test condition 360 hrs 720 hrs 1440 hrs
A 1
2
B 1
2
C 1
2
D 1
2
NCG > 20
NCG > 19
NCG > 40
NCG > 41
NCG > 42
NCG > 40
NCG > 20
NCG > 22
NCG > 20
NCG > 21
NCG > 41
NCG > 39
NCG > 40
NCG > 43
NCG > 17
NCG > 20
NCG > 22
NCG > 20
NCG > 42
NCG>41
34.8
35.8
NCG > 18
NCG > 20
NCG = no crack growth
TABLE 7
RESULTS OF THE TENSILE TESTS
Material Test condition Tensile test result
1 No failure
A
2 No failure
1 No failure
B
2 No failure
1 No failure
C
2 No failure
1 No failure
D
2 Failed
111/15

TABLE 8
RESULTS OF THE LOCALIZED CORROSION TESTS
Material Test condition Pitting test Crevice test
A 3 No pitting No crevice
4 No pitting No crevice
B 3 Small dispersed pitting (Al, B 1) No crevice
C 3 Large pitting (A2, B 1) Large crevice pits
D 3 Lar&na&d&pkd pitting (Al, B 1)
4 Small dispersed pitting (Al, B 1) No crevice
TABLE 9
RESULTS OF THE GENERAL CORROSION TESTS
Material
A
B
C
D
Corrosion rate
Test condition (mm/ye@
3 0.005
4 0.01
3 0.03
4 0.01
3 0.01
4 0.009
3 0.09
4 0.22
111116

0.4
0.35
0.3
m 0.25
B 0.2
8
0.15
0.1
0.05
0
7 1
T
0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 5.5
LENGTH 2c
Figure 1 .- Typical set of damage tolerance curves, showing the
respective stable crack growth zone and fracture zone.
.11.11
Figure 2 .- Schematic diagram of the tests
samples used in the localized corrosion tests,
showing their typical dimensions.
Figure 3.- Schematic diagram of the
weight loss coupon used in the general
corrosion tests, indicating the typical
dimensions.
Figure 4 .- Schematic diagram of the testing setup.
111117

Figure 5 .- Histogram showing the pit
diameter distribution of steel B at test
condition 3.
30.
25.
20.
c
d 15.
10.
5.
OT I ! ! I, T
.Ol .02 II3 .04 .05 .06 .07 .08 09 .I .,I
Dhmoter (nm)
Figure 7 .- Histogram showing the pit
diameter distribution of steel D at test
condition 3.
KIW71MPa'nN!
L
0.1 1 10 100 iooo
Ddecl
LmgL(am)
Figure 9 .- Damage tolerance curves for a
tubing of 3.5” external diameter made of
steel B.
14-
12-
%10-
8 a-
6-
0 12 3 4 5 6 7
Diameter (mm)
Figure 6 .- Histogram showing the pit diameter
distribution of steel C at test condition 3
KIC=367
MPaWQ.5
k
1 10 100 1000
Oafad
Lmgth
(mm)
Figure 8 .- Damage tolerance curves for a tubing
of 3.5” external diameter made of steel A.
LKIC.lB3MPa'nN!
1 10 100 1000
C&i Langth
2c(mm:
Figure 10 .- Damage tolerance curves for a tubing
of 3.5” external diameter made of steel C.
111118

KICd04.7 MPa*m*0.5
=L
10 100 1000
Defect Length 2c (mm)
Figure 11 .- Damage tolerance curves for a tubing of 3.5” internal diameter made of
steel D.
111119

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