Proceedings of the Eleventh (2001) International Offshore and Polar Engineering Conference Stavanger, Norway, June 17-22, 2001 Copyright © 2001 by The International Society of Offshore and Polar Engineers ISBN 1-880653-51-6 (Set); 1SBN 1-880653-55-9 (VoL IV); ISSN 1098-6189 (Set)

Sour Service X65 Seamless Linepipe for Offshore Special Applications

E. Anelli andD. Colleluori, CENTRO SVILUPPO MATERIALI S.p.A., Roma, Italy J.C. Gonzalez, SIDERCA S.A.I.C. (DST Group), Campana, Argentina

G. Cumino, DALMINE S.p.A. (DST Group), Dalmine, Italy H. Quintanilla andM. Tivelli, TAMSA S.A. (DST Group), Veracruz, Mexico

ABSTRACT

The recent trends in the design criteria, laying technologies and service conditions for offshore pipelines call for high performance seamless pipes tailored to the specific applications required by the market. Experiences that have been successfully performed by DST Group in the large-scale production of seamless pipes for special offshore applications are presented in this paper.

A 0.1C-IMn steel, with optimized Nb and V contents, was developed for quenched and tempered (Q&T) seamless pipes (O.D. 219.1 mm, W.T 15.4 mm), which showed excellent mechanical and corrosion performances after full-scale-simulated reeling.

For heavy wall pipes (O.D. 304.5 mm, W.T. 34 mm) a Mo-V-Ti steel was selected in order to guarantee a suitable combination of strength, yield to tensile ratio (YS/UTS) and work hardening ratio (WHR) even at high temperatures (160°C). Adequate strain aging and HIC resistance was also obtained in such Q&T seamless pipes, which have been successfully delivered for a high temperature, high pressure offshore production line.

KEY WORDS: High Temperature High Pressure Pipeline, seamless pipe, sour service, strain-aging

INTRODUCTION

The operating conditions of submarine pipelines become increasingly demanding and require materials able to operate at great depths for the transportation of very high temperature and pressure fluids. In addition, the pipelay operations, which are involved when linepipes are installed by means of the reel method, cause repeated bend cycles. These cyclic loads affect material stress-strain properties. Moreover, reeling is currently applied to an increasing range of pipe geometries, being the present limit around 16" O.D., 30 mm thick pipes.

Accumulated plastic strain reduces the ductility and toughness of the pipe material. Strain aging and toughness tests must be carried out following specified procedures when the pipeline is to be exposed to more than 2% plastic strain during installation (DNV, 2000). Plastic deformation of the pipe may also decrease the resistance to sulfide stress corrosion cracking.

The ratio between the yield stress and the ultimate tensile strength (YS/UTS) should not exceed a certain critical value in order to guarantee an extra safety margin. Related to YS/UTS ratio are the strain hardening properties of the steel. Strain hardening is important for local buckling stability in plastic bending (Gresnigt, 1994). In order to apply a limit state or strain based design (Bai and Damsleth, 1997) we need a much more detailed knowledge of the material properties such as the stress-strain relationship at various temperatures between ambient and the operating condition.

Modern seamless pipes are suitable for offshore applications, provided a proper metallurgical design and processing route are set-up. In this framework, the results of recent experiences are presented in the production of high performance quenched and tempered (Q&T) seamless pipes for special offshore applications. Sour service grade API 5L X65 produced by DST for flowlines and for production pipes are considered in this paper.

The metallurgical design criteria and process conditions to manufacture these linepipes are presented together with the most significant results of a comprehensive program of characterization, including microstructure, yield and tensile strengths, work hardening ratio (WHR), notch toughness, stress-strain curves at various temperatures (tensile and compression), strain aging susceptibility, HIC and SSCC resistance. To assess these properties both small and full- scale testing methods have been extensively used.

METALLURGICAL BACKGROUND FOR SELECTING THE OPTIMUM PROCESS CONDITIONS

Seamless pipes of medium size (O.D. up to 16") are presently produced by a hot rolling process carried out in the following main stages: piercing, rolling at retained mandrel mill and sizing. Quenching and tempering treatments are performed on the pipes in order to refine the microstructure and obtain the required properties (GonzAlez, 1998).

A rational approach to the design and production of these materials requires the quantitative knowledge of the effects of steel chemistry, thermomechanical and heat treatment variables on the microstructure and final mechanical properties. The influence of microalloying additions, thermomechanical and Q&T practices on austenite

242

refinement, phase transformation and response to heat treatments of low-C steels for seamless linepipes was investigated by dilatometry, hot torsion tests, mathematical modeling, pilot and industrial trials (Anelli, 1992, 1994a, 1994b, 1996, 1997). This study allowed to establish the guidelines to select chemical composition and process conditions.

Hot Piercing and Rolling

Hot deformation conditions - e.g. temperature (T), strain (e), strain rate (g')-typical of piercing (T = 1200-1300 °C, e> 1, ~' = 0 . 3 - 3 s "~) and of the first rolling passes in the retained mandrel mill (T = 1100 - 1190 °C, e > 0.5, e' = 4 - 8 s "~) are able to activate the dynamic recrystallization (DRX) of austenite in C-Mn and microalloyed steels (Anelli, 1994b). A DRX austenite grain size (AGS) of 15 to 25 gm is typical of microalloyed steels deformed at 1100 - 1250 °C. These grains can grow rapidly unless Ti microadditions are used.

Static recrystallization is present at the final rolling passes in the retained mandrel mill. Because the applied deformation schedule incorporates low strains at the end of the sequence, recrystallization- induced grain growth can result. The AGS of as-rolled pipes ranges from 35 to 55 lain for V steels and 30 to 40 gm for V-Nb microalloyed steels, depending on the hot rolling schedule and final wall thickness. The controlled microaddition of Ti (0.015%) helps in maintaining the AGS of as-rolled pipes into values between 20 to 30 gm, even in the case of high finish rolling temperatures (heavy wall pipes).

Quenching

The final AGS depends on the austenitizing temperature and holding time, nature and size distribution of precipitates present after hot rolling and natural cooling. The more uniform the as-rolled microstructure, the easier it would be to homogenize the austenite.

Laboratory tests and industrial trials have shown that to avoid the formation of coarse austenite grains (AGS > 25 gm) in low carbon steels (0.08-0. I 1%C) the heating temperature has to be lower than 900 °C for C-Mn steels; however, V-Nb steels and V-Ti steels can be safely austenitized up to 920 to 950 °C to dissolve V-rich precipitates, without problems due to the pinning effect of Nb (C, N) and Ti (C, N) fine particles which hinder grain boundary movement.

Of concern in a quenching process of seamless pipes are the effects of through-thickness cooling rate gradients, induced by surface water cooling, and the sequence of transformation and the resultant microstructure and hardness profile.

The production of these pipes requires the quantitative knowledge of the effects of steel chemistry and cooling rate on the microstructure. A specific test program was performed with the main objective of measuring the phase transformation characteristics of austenite under continuous cooling conditions by dilatometry and metallography (construction of CCT diagrams). Mathematical modeling, which links the basic principles of heat transfer and microstructural phenomena have been effectively applied in this field.

The volume fraction of microstructural constituents and hardness of as-quenched linepipes were predicted as a function of the local cooling rate, calculated by a thermal model based on the general Fourier heat equation, by means of an Artificial Neural Network (ANN) trained on a selected database of CCT diagrams of linepipe steels.

Simulations carried out by the model and industrial trials have shown that linepipes with wall thickness less than 16 mm can be effectively quenched passing continuously through water jets produced

by nozzles arranged in a series of rings forming a tunnel. In the case oI heavy wall pipes, external and internal quenching is needed to reduce hardness gradients and make more homogeneous the as-quenched structure. This type of quenching is carried out by dipping the pipe in a tank containing stirred water. During quenching the pipe is under rotation and an internal water jet is used, too.

A few examples of through-thickness hardness profiles of as- quenched pipes of Nb-V steel (W.T. = 15.4 mm, external quenching) and Mo-V-Ti steel (W.T. = 34 mm, external and internal quenching) are compared in Fig. 1.

"1" (n ¢}

6-

-1-

250

240

230 220

210

200 190

180 170

N b - V Steel (W.T. = 15.4 mm) Externa l Q u e n c h i n g

i ' i i ' .... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . i ........... ¢ . - l ~ o ' c ~ o ~ l . . . . . . . .

: 1- ..... 3 o'crock [ . . . . . / ............ I . . . . . . ! ....... I -*-- - Bo'e~ock / .......

i I --~-. 9o'c,ock /

0 2 4 6 8 10 12 14 16 Depth (mm)

-r- ff) ® ¢..

32

Mo-V-T i S tee l (W,T. = 34 m m ) Externa l and Internal Q u e n c h i n g

240250230 ' ' ' ' ' ' ' 1 ' ' ' ' . ' ' ' : i I',t :; ~ " ' ! '

220 i '

2 , o ! ........ i .... 200 , : . i ,00'i 180. " ' , 170 . . . . . . . . . . . . . . . . . . . . . . . . . .

2 4 6 8 10 12 14 16 Depth (mm)

Figure 1. Through-thickness hardness profiles (external surface = 0 mm depth).

The hardness gradient for pipes with W.T. < 16 mm is typically less than 1.2 HV~dmm, although external quenching is used. In addition to the more effective cooling, the improved hardness uniformity of the heavy wall pipe is owed to the higher hardenability induced by the presence of 0.26% Mo.

In Fig.2 the phase transformation characteristics of both steels are compared. In order to have a volume fraction of bainite and martensite higher than 50% and a fine ferrite microstructure, the cooling rate should be greater than 40 °C/s for the Nb-V steel, whilst the Mo addition allows to reach a similar as-quenched structure and hardness values at half of the cooling rate.

A predominant structure of mixed lower bainite and martensite appears beneficial to reach adequate toughness levels after tempering. Also fine and homogeneous austenite grains are important, especially to avoid the formation of coarse bainitic structures that contribute to reduce toughness.

243

1 0 0 0

800

,~ 60o

4 0 0

2 0 0

1 0 0 0

o• 8 0 0

6 0 0

400 ¢,

200

Figure 2.

. . . . . . . . i . . . . . . . . i . . . . . . . . i . . . . . . . . l

Nb- V Steel

I ~ 1!~0 80 40 20 10 5 I [

I 307 246225200 188 175 161 ~ I . . . . . ' - I , . . . . . i . . . . . . . . t . . . . . . . . i

i0 i 1 0 2 1 0 3 l 0 ~

Time ( s )

. . . . . . . . i . . . . . .

mass (%'~ C 0.11 Mn 1,02 Si 0.28 S 0.0015 P 0.009 Ni 0.01

ICr 0,13 [Mo 0.09

C u 0 . 1 8

Ti <0.005 V 0.071 Nb 0.018 AI 0.026

~o ~ ~o 6

Mo-V-Ti Steel mass ( % :

"---.. c o.1o

~ 1 Mn 0.98 . Si 0.25 S 0.0018

~ P 0.006 ' - - Ni 0.06

Cr 0.30 Mo 0.26

I Cu 0.16 Ti 0.020 V 0.068 N b <0 .005

I ' l l ' ~ ° s ° 4° 2° ~° ' ~ I AI 0.025 3"/8 273 246 228 19"J l~/O 146 "

I01 10 2 10 3 10 4 10 5 10 6

Time ( s ) Continuous cool ing transformation (CCT) diagrams. AGS = 8 p.m.

Tempering

The tempering conditions, mainly temperature and the presence of elements - such as V and Mo - able to give precipitation hardening and to slow the recovery/recrystallization process, are the controlling factors for the combination of strength and toughness, for a given as- quenched microstructure. Usually, increasing the tempering temperature for a same strength level, through appropriate microalloying additions, leads to better low-temperature impact properties. For linepipes of X65 grade, the tempering temperatures should be higher than 600 °C to have high toughness levels, suitable to satisfy the specifications requirements, especially after strain-aging.

MICROSTRUCTURE AND STANDARD MECHANICAL PROPERTIES

On the basis of the metallurgical design criteria previously discussed and the material requirements established by the oil companies, two chemistries (Table I) were selected for the production of sour service grade API 5L X65 flowlines (O.D. = 219.1 mm, W.T. = 15.3 mm) and production pipelines (O.D. = 304.5 mm, W.T = 34 mm).

The Q&T treatments were calibrated for the specific chemical composition and pipe size in order to develop a suitable microstructure able to fulfil the target mechanical properties.

Some examples of the microstructures formed through the pipe thickness are shown in Fig. 3 for both the products. The microstructures appear fine and uniform, and reflect the features of the

original as-quenched microstructures in terms of repartition oI constituents and AGS.

Nb-V Steel 0V.T. = 15.4 mm) Mo-V-Ti Steel (W.T. = 34 nun External Quenching External~d Internal Quenching

• ~ ~ , - k . . ~ TM ~ ~ "~ ,...~7.~ ." v'..,~ , % - ~ ' . .~ , :. " : . . -

' ~ - - . , . . ~ , .~.:~ .. . . . ~ g - ~ ' ~ " % " ~ : ~ ' Z ~ "~ ~" "~'~":~i ' ~ ~ ' " ~

• ~.~ ". ~d,, ~"~ ,z~ ~ " " £

. . . . . . . . . . . . . . . . . . . ' ~ ~ ' - ~ ! , . ~

b)

c)

Figure 3. Typical microstructures of Q&T pipes at different through- thickness positions: a) external side; b) mid-thickness; c) internal side.

Steels Nb-V Mo-V-Ti

Values Min Max Mean Mitt Max Mean C(%) 0.09 0.12 0.10 0.09 0.12 0.11

Mn (%) 0.94 1.00 0.96 0.93 1.00 0.96 Si(%) 0.19 0.27 0.23 0.21 0.26 0.23

P (ppm) 70 130 90 50 80 70 S(ppm) 10 20 t3 10 30 20 Nb (%) 0.019 0.026 0.023 - <0.005 <0.005 I14o (%) 0.08 0.10 0.09 0.25 0.27 0.26 Cr(%) 0.10 0.17 0.12 0.01 0.06 0.03 V(%) 0.040 0.053 0.045 0.059 0.068 0.063

Ni (%) 0.I0 0.11 0.1 ! 0.04 0.07 0.05 Cu (%) 0.12 0.18 0.14 0.08 0.16 0.13

Sn (ppm) 70 100 80 60 110 70 Al(%) 0.022 0.036 0.027 0.020 0.030 0.025 T/(%) 0.002 0.004 0.003 0.013 0.020 0.017 Ca (%) 10 28 16 10 20 17

. N(ppm) 70 103 83 44 76 59 Ceq 0.310 0.350 0.330 0.329 0.376 0.348 Pcm 0.171 0.185 0.202 0 .171 0.185 0.202

Table I. Product analysis from 50 heats of the Nb-V steel and 100 heats of the Mo-V-Ti steel.

244

The reproducibility of both chemical compositions (Table I) and heat treatments for various heats allowed to maintain the variability of mechanical properties in a limited range (Table II), satisfying the required values.

Tensile test at room

temperature (*)

Hardness HVIo

Transverse Charpy- V

Notch Energy (J)

(*) (o)

Values

YS (MPa)

UTS (MPa)

YS/UTS

El (%) Ext. (9

MieL

Int. (9

-40°C

-lO°C

Steels

Nb-V W.T. = 15.4 mm

Min Max Mean

460 560 500

540 630 600

0.83 0.89 0.85

36 44 41

180 238 210

178 213 192

165 205 185

170 290 250

302 450 411

Transverse full size specimen Below the surface (2 mm depth)

Mo-V-Ti W.T. = 34 .0 mm

Min Max Mean

463 569 528

561 633 599

0.82! 0.90 0.88

22 48 36

169 233 197

176 216 196

163 231 190

50 120 80

144 290 201

Table II. Mechanical properties.

Another key factor was the set-up of suitable steelmaking and continuous casting practices able to control the content of non metallic inclusions and the segregation levels. The small length and the low volume fraction of oxides and sulfides, combined with the absence of severe banding, allowed to reach a high resistance to hydrogen embrittlement. HIC testing, carried out according to NACE TM 02 84- 96, using solution A, gave values of the parameters CLR, CTR and CSR (Table III) well below the maximum acceptable limits required by the most demanding Specifications (10%, 2.5% and 1%, respectively).

Steel (W.T.)

Nb-V

(15.4 ram)

Mo-V-Ti

(34 mm)

Specimen CLR Specimen CTR Average Average

I 2 3 1 2 3

0 0 0 0 0 0 0 0

6.2 8.7 3.1 6.0 0.26 0.32 0.25 0.28

SPECIAL TESTING FOR THE ASSESSMENT OF IN-SERVICE PERFORMANCE

High Temperature Stress-Strain Behavior

The design of lines for transport of fluids at high temperatures requires the detailed knowledge of material properties at the operating temperature. Of course, the strength variation versus temperature is a key factor for pipes for such applications.

For both pipe types, a reduction in strength (6 to l0 %) was actually observed when temperature was increased up to 80-100 °C. At higher temperatures (140 to 160 °C) the values of YS and UTS remained similar to those measured at 80-100 °C (Fig.4).

700

650 ..............

"-" 600 O.

~ ' 5 5 0

5OO

i .................

U

i I • Nb-VSteel(VgT.=lS.4mm) ~ 450 I i - -=- -Mo.V-Ti Steel (W.. T. = 34 ram)

1. . , . . ! . 1 J . I 1, 4000 20 40 60 80 100 120 140 160 180 200

Temperature (°C)

700

650 [ Nb-VSteel(W.T.=lS.4mm) ; .... Mo-V-Ti Steel (W.T. = 34 ram)

600

D ....~; 550 [ ~

>" 500 "

450 ................

400,'----J -: . . . . .--J -- . . . . - . . . . . . . . . . . . . . . . . . . . . .--t--.--v-.--- 0 20 40 60 80 100 120 140 160 180 200

Temperature (°C)

Steel Specimen (W.T.) 1 2 3

Nb-V 0 0 0

(15.4 ram)

Mo-V-Ti 0.002 0.004 0.005

(34 mm)

Table III.

CSR Cracks Average Position

0

0.004 mid-thickness

HIC results according to NACE TM 0284-96, solution A.

Figure 4. Yield stress (YS) and ultimate tensile strength (UTS) ot Q&T linepipes at various temperatures.

It is important to note, however, that the yield strength remained well above the minimum acceptable value for all the temperatures considered.

In the ease the linepipe is to be used in a strain based design application at high temperatures, a specific testing program could be required to confirm material characteristics. Cyclic tests with load between zero strain and maximum compressive strain of 1% to be performed at room temperature and at high temperatures (up to 160 °C)

245

could be demanded. Another type of material assessment is hot tensile testing after compression at room temperature.

Table IV shows the results obtained on a Q&T Nb-V steel pipe (OD = 355.2 mm, WT = 22.2 mm), grade API 5L X65, for a testing program including:

= tensile test at room temperature (RT); * tensile test at 115°C; * tensile test at RT after 0.5% compression;

tensile test at 115°C after 0.5% compression.

Values

Min

Max

Mean

Tensile test at room

temperature YS YS/UTS WHR

(MPa) (-) (-)

463 0.82 0.92

569 0.90 0.94

528 0.88 0.93

Compression test at 160°C

YS WHR (MPa) (-)

413 0.82

545 0.88

490 0.85

Compression at room

temperature

(%)

Compression

Stress at 0.5%

(MPa)

Tensile Testing

Temperature

(°C) Rto.s Stress at

0.5% T.E.

(MPa)

Stress at

3% T.E.

(MPa)

0 0.5 0 0.5

528 533

22 22 115 115

532 501 505 485

575 575 565 561

UTS 636 638 608 608

(MPa)

Rto.5/UTS 0.85 0.78 0.83 0.80 (-)

Table IV. Effect of 0.5% compression at room temperature on tensile properties at 22 and ! ! 5 °C.

The tests were carried out on longitudinal specimens using a 100 kN servo-hydraulic tensile/compression machine with special grips allowing to work at high temperatures with controlled axial loading condition according to ASTM E21 and EN 10002 - part 5a Standards. Specimen heating was performed with a lamp furnace controlled by • microprocessor. Specimen temperature was measured by a K type termocouple. Deformation was monitored by suitable clip. Data arising from clip, load cell and thermocouple were acquired by a dedicated PC.

As expected, after compression and tensile testing, due to the Bauschinger effect, the stress recorded at 0.5% total elongation (R r05) is decreased. The amount of the Bauschinger effect depends mainly from the compression strain and is practically independent from the examined temperature range (22 to 115 °C). The Rto.5 reduction is 20 to 30 MPa after 0.5 % compression.

Also the work hardening of the material is a key factor in the strain- based design for pipelines. The typical values of YS/UTS and WHR (stress at 0.5% T.E./stress at 3% T.E.), at various temperatures are shown in Table V and in Fig.5.

Table V. Work hardening characteristics at room temperature and at 160°C.

CO 1"-

CO >-

0 . 9 2

0.90 . . . . . . . . .

0 . 8 8 .

0 8 0 . . . . . . . . . . . . _ - i . . . . . . . i . . . . -

0 . 7 6 [- ........ , ......... , ................... : . . . . . . . . . . . . i ....... i [ . . . . . . . . . . . . [

o7, i i I I " I " ~ . . . . . . . . . . . . . . i ......... I . . . . . - - . - - M o - V - T i S t e e l O t E T . = 3 4 r a m )

0.721- ........ ] ...... ; . . . . . . I . . . . . [ . . . . . . . . . . . . . . t ~. I

o7o[, ; , , , , i , I , , , I , [ , , , i 0 20 40 60 80 100 120 140 160 180 200

Temperature (°C)

I Z -r"

0 . 9 8

0 . 9 6

0 . 9 4

0 . 9 2

0 . 9 0

0 . 8 8

! I I I • • . ] . . . . . . ~ . . i 1

I i

....... , ~ ••! -I . . . . . . i i •

-4 ': I z i 0 " 8 6 1 I I . I

r ! i I I / I I I i :

0 . 8 4 [ - . . . . . . . F - ~ • N b . V S t e e l ( W . T . = 1 5 4 m m ) [

0 " 8 2 I t I . . . . M ° ' V ' T i S t e e l O N ' r ' = 3 4 m m ) t ' I

0 8 0 - i I I I i F I , i 0 2 0 4 0 6 0 8 0 1 0 0 1 2 0 1 4 0 1 6 0 1 8 0 2 0 0

Temperature (°C)

Figure 5. Work hardening characteristics as a function of temperature (longitudinal round specimens machined at pipe mid- thickness).

At ambient temperature, the results were quite satisfactory, while tow YS/UTS ratio and high work hardening rate were observed at temperatures higher than 100 °C. In performing the tensile test on cylindrical specimens, the strain-stress data have been automatically

246

recorded to be used in deriving equations for applying the non-linear Finite Element Models. A few examples of actual stress-strain characteristics, measured for the same material at various testing temperatures are shown in Fig.6.

t~ o.. v ~9 09

C~ n

v

800

700

600

500

400

300

200

100

0 0

800

700

600

500

i i

n =0.11

T = 1 0 0 ° C

o o.' o o'. s o. o strain

• =

n = 0 ~ 2 ~

400

300

200

100

o;

T = 1 6 0 " C

I .~-v ~,.., (w.r, = ~.~ ,,,m) ] [ ................. M ~:Y:E.~!~.*L!w:r:.~_~ m~.!J

o,bs 020 strain

Figure 6. Stress - strain curves at various temperatures (longitudinal round specimens machined at pipe mid-thickness).

The work hardening coefficient "n", evaluated from the experimental data using as fitting curve the equation ~ = K e " (e > 0.01), describes in a better way the residual deformation capability after yielding.

The higher is the work hardening rate, the larger is the curvature which the pipe can withstand without buckling (Bruschi, 1994).

Strain-Aging Susceptibility

For linepipes to be used in installation systems introducing an accumulated plastic strain > 2%, additional requirements do apply, concerning their strain-aging susceptibility (DNV, 2000).

In order to assess this issue, Charpy V - notch testing was performed on samples removed from finished pipes, successively deformed by uniaxial tension at strains of 3% and 5%, and artificially aged at 250 °C for one hour before testing.

The impact energy and the brittle fracture appearance are plotted in Fig.7 for the base and strain-aged materials. The ductile to brittle transition temperature is increased of 10 to 15 °C and 15 to 25 °C after 3% and 5% deformation, respectively.

N b - V S t e e l ( W . T . = 1 5 . 4 m m ) M o - V - T i S t e e l ( W . T . = 3 4 m m )

~ ~ , ~ " ~ ! ' B a s e M a t e r i a ~ ~ ~ ~ , t i

I + '

+ , i + [ + i + i , , i

t 1 2 4 0 , . +- : '

1 8 0 . J ......... Y l + ~ . - i . ~ - . . . ~ ......... ~ ......... e Q

/ ~ ~ ~ , ~ * I I - * ' B t l t t l e a r e a l I ! i y. + ! + i / ,20I ........... ~ : : : i + +--1 + - : , ....................... / r , ~ + L - 4 1,0 00t+4+ 4 ++-I +M+ + + bo

f ' t ~ + ' ~ ! S t r a i n a g e d 3 % + 2 5 0 ° C - l h ~ ~ i ] 3 O O ~ 1 " ~ l F F '" " r ' l ~ i " i " " } . . . . . . . . . . . . . . . . . . . . . r " " -- F F 11" ; . J , . " 111 i . } F.F F H . ' -- ~H ' 111F'. I ] ~ I - - - - ~ - - ' k ~ i ~ - - - - ~ a - - - - ~ } . . . . . . . . . . . ~ I O O ~

: ~ ~ : o + = + ! , ~ * t , ' • n • +

i ! . . . . . . . . . . . . . . i . . . . . . . . . . . . V ' " ' " ' " ; " " ' " '~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S O

Z , 2 t J "["'""t+ ' . . . . . . . . . . . . . . . . i i i ~i . . . . . . . . . . . . . . . . . . ! i .................... i .... | ............... i ~ " ~

...................................... ................................................ t+i++ ............................ ............................ t +o m + i 4 + + : i i ~".* I

m i . . . . ' + ' S t r a i n a g e d 5 % + 2 5 0 ° C - l h . . . . P" 3 O O l F " ~ : ................... i f ..................... ~ - --I I -~ ' - - + - + - - , ~ - - i - " ~ - - 4 ' i - - , | 1 O0

4.+--++-, . . . . . . . . . . . . . . . .

*, + ~ + ~ ~ ~. P i

t ' ° + " ' '°° -~-- ~ ~ + - " > . . . . . > - r - - r - - h - - ~

+ 1 t ~o ~ - ~ - - - ~ - _ ~ ~._....L._.__~...,...._.. *.*. $_.._.i_ ...~ ................ ~o

Ol i * ~ i . i I ~.~,,.d. ! / f + ~ J ; , ' F ' ~ & - - / 0 -100 -80 -.80 ,40 -20 0 20 40 -'IQ~ -80 -60 ,40 -20 0 20 40

Temperature (°C) T e m p e r a t u r e (°C)

Figure 7. Evolution of impact properties after artificial strain-aging.

Impact tests carried out on the deformed materials, without treating at 250 °C, gave an increase of the fracture appearance transition temperature (50% FATT) of 10 to 15 °C. Only about half of the toughness deterioration can be attributed to aging. A very much reduced strain-aging sensitivity was therefore found for all the test materials. The slight F A T T increase exhibited even from Ti- microalloyed steel, when nitrogen should be fully bound to Ti, indicates that carbon in solution should be responsible of aging at 250 °C. Of course the strain-aged materials also exhibited some strengthening, but the yield strength increment remained below the maximum accepted value of specified minimum yield strength (SMYS) + 100 MPa. Again, only 20 to 30 MPa of strengthening can be attributed to strain-aging.

Full Scale Tests

A programme of full scale tests was conducted on flowlines in order to assess specific in-service performances required by the customer.

Reeling and straightening simulation

The trials were intended to subject 12.5 m long pipe to the Ioadings likely to be developed when the flowlines are installed and recovered using the reel method. The simulations were carried out following a specific procedure set-up for the simulation of the deformation cycles that can be experienced a reel ship, using a test arrangement in U.K.. The test system basically consists of two spine beams, hinged together at one end. Permanently attached to the inner face of one spine beam is the reeling former and on the other a straightening former which is interchangeable, The radius of the curvature of the reeling former was approximately equal to the radius of curvature of the hub of the drum (i.e. 8.2 m). The extreme fiber strains developed in the linepipe material are given by the following relationship:

Outside radius of pipe cross section /Radius of curvature.

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A straightening former with 38 m radius was used to bend the length of the pipe in the reverse direction after it had already been bent around the reeling former. An axial load of 200 kN could be applied during bending. The operations of reeling and straightening were repeated a second time. The four cycles intended to simulate also the stages required to recover the pipeline at a later date. The total accumulated strain was estimated 12.8 %. The pipe was positioned in the test arrangement such that the 9 o' clock position is that which would experience the maximum extreme fiber longitudinal tensile strains when bending is applied. Tensile and impact toughness tests were performed on samples removed from the 3, 6, 9 and 12 o' clock positions. Results are summarized in Table VI. A slight yield strength reduction was observed at 3 and 9 o' clock positions for the longitudinal specimens. Toughness is practically unchanged.

Also the SSC resistance of the flowlines submitted to reeling and straightening simulations was assessed. Four point bent beam specimens were stressed at 90% of the actual yield strength. Specimens, whose dimensions were 115 mm x 15 mm x 5 mm, were removed from the 3, 6, 9 and 12 o' clock positions and ground up to 300 mesh grit paper. The specimens (four for each position) were washed, degreased in tetrachloroethylen, dried and tested in the autoclave (20 1). Testing solution was prepared according to NACE TM 01 77-96 solution A. The acidity value was obtained through small additions of HCI up to reach the desired pH values, 3.5 and 2.7, respectively.

Pos.

3 I o'clock !

6 o'clock l

9 o'clock

12 o'clock

Longitudinal specimens

YS UTS YS/ El (MPa) (MPa) UTS (%)

465 593 0.78 43.0 460 592 0.78 42.7

494 590 0.84 44.7 474 581 0.82 44.7

458 597 0.77 44.7 465 583 0.80 44.7

486 585 0.83 45.7 465 570 0.82 39.8

Transverse specimens

YS UTS YS/ El (MPa) (MPa) OTS (%)

493 584 0.84 35.8 509 608 0.84 38.8

501 603 0.83 37.8 491 601 0.82 41.7

480 581 0.83 41.7 526 610 0.86 41.7

480 597 0.80 39.8 488 602 0 .81 41.7

Pos.

3 o'clock

6 o'clock

9 o'clock

12 o'clock

Transverse Charpy - V notch

Energy (J)

-40°C -10oc 260 269 262 273 265 270

n.d. md.

256 285 279 280 260 278

252 292 262 2 8 6 265 295

Table VI. Mechanical properties after full-scale simulated reeling and straightening of 12.5 m long line pipe of Nb-V steel (O.D. = 219 ram, W.T. = 15.4 ram). Full size specimens.

The solution was deaerated by high purity nitrogen, flowing for more than 24 hours before starting the test. Electrical insulation of specimens was made by internal coating of teflon foils to prevent galvanic coupling. H2S was introduced at ambient temperature and bubbled for one hour.

Then the temperature was controlled at 23 +/- 2 °C. The pH of the solution was measured before starting the test and just after completion. The testing conditions are reported in Table VII.

Test

No.

A

B

Solution pH pH NACE TM 0177-96

Initial Final 50 g/l NaCI

4 g/l CH3CONa 2.71 4.18

50 g/I NaCI 4 g/l CH3CONa 3.48 4.55

Test Temperature

No. ¢C)

A 23+2

B

Total Pressure

bar) 1

Partial H2S Pressure

t b~) i

23 + 2 1 1

Table VII. Autoclave testing conditions.

Because no cracks were detected on the specimen surfaces by stereoscopic microscope examination (magnification 50 X), after washing in deionized water and removing of corrosion products, the SSC resistance of Q&T linepipes remains high although an accumulated strain was introduced by simulated reeling.

Full ring test

Full ring testing of a girth welded pipe was carried out in accordance with the procedure published in the HSE Document OTI 95 635, at an applied toad of 72.8 % of the SMYS in NACE TM 01 77-96 solution A. The loading was applied by means of an internal turnbuckle. The girth weld was at the center of the sample and the length was such that residual stresses would be retained. The test duration was 30 days. The composition of the test solution introduced into the pipe sample was as follows:

5% NaC1, 0.5% CH3COOH, initial pH = 2.6, final pH = 3.8

The H2S was bubbled through the solution throughout the test. The H2S concentration was checked regularly and maintained at 2300 ppm or greater. The full ring sample was controlled at a temperature of 25 +/-4 °C. Hydrogen permeation measurements were carried out by Devanathan-type cells fitted to the external surface of the pipe section. On-line monitoring was performed. The ultrasonic inspections revealed that no indications developed during the course of the test. After testing the section was drained and unloaded. Wet magnetic particle inspection of the weld inside surface detected no defects.

The pipe submitted to full ring test met the acceptance criteria under the imposed test conditions and tolerated a hydrogen flux equivalent to a permeation current of 100 mAm -2.

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5. CONCLUSIONS

The following conclusions were drawn from the extensive testing program on seamless Q&T pipes for special offshore applications:

• The typical operation conditions of a flowline were effectively simulated by small and full scale testing methodologies and detailed properties of materials were determined.

• The results clearly showed that a reasonable balance between customer specifications, processing capabilities and product properties can be found for Q&T seamless pipes.

• More specifically, a 0.1C-1Mn steel, with optimized Nb and V contents, was developed for quenched and tempered (Q&T) seamless pipes (O.D. 219.1 mm x W.T 15.3 ram), which showed excellent mechanical and corrosion performances after full-scale- simulated reeling.

• For heavy wall pipes (O.D. 304.5 mmx W.T 34 ram) a Mo-V-Ti steel was selected in order to guarantee a suitable combination of strength, yield to tensile ratio (YS/UTS) and work hardening ratio (WHR) even at high temperatures (160°C).

• Adequate strain aging and HIC resistance was also obtained in such Q&T seamless pipes, which have been successfully delivered for a high temperature, high pressure offshore production line.

Gonz~dez, JC, Tivelli, M, Quintanilla, H, and Cumino, G (1998). "DST Experience in High Performance Linepipe and Flowlines for Sour Service", Proceedings of the 40 th Mechanical Working and Steel Processing Conference, ISS, 1998, pp 621-626.

Gresnigt, AM, Dijkstra, OD, and Rongen, HJM van (1994). "Design ot Pipelines in High Strength Steel", Proceedings of the Fourth International Offshore and Polar Engineering Conference (ISOPE), Vol. II, Osaka, Japan, April 10-15, 1994, pp 186-194.

REFERENCES

Anelli, E, Poli, A, Oulhadj, A, and Cumino, G (1992). "Development o f Microalloyed Steel Seamless Tubes", Proceedings Conf. On Pipeline Reliability, Calgary, Canada, June 1992, III-6/ppl-11.

Anelli, E, Manigrasso, M, and Cumino G (1994a). "Microalloyed Steels for On-Line Treated Seamless tubes", Proceedings of the 3~ ~ Mechanical Working and Steel Processing Conference, ISS, Vol. XXXII, Baltimore; USA, October 1994, pp 161-169.

Anelti, E, Farsetti, P, and Cumino, G (1994b). "Prediction of the Response to Quenching and Tempering of Low Alloy Steels for Seamless Tubes", Proceedings of the 36 th Mechanical Working and Steel Processing Conference, ISS, Vol. XXXII, Baltimore; USA, October 1994, pp 149-159.

Anelli, E, Farsetti P, and Cumino G (1996). "New Process for On-Line Normalizing of Seamless Pipes", Proceedings of the 38 ~h Mechanical Working and Steel Processing Conference, ISS, Vol. XXXIV, Cleveland; USA, October 1996, pp 285-295.

Anelli, E, Cumino G., Gonz~lez, JC (1997). "Metallurgical Design of Accelerated-Cooling Process for Seamless Pipe production", Proceedings from Materials Solutions '97 on Accelerated Cooling/Direct Quenching of Steels; Indianapolis, USA, 15-18 September 1997, pp 67-75.

Bai, Y, and Damsleth, PA (1997). "Limit-State Base Design of Pipelines", Proceedings of the International Conference on Offshore Mechanics and Arctic Engineering (OMAE), Vol. V, pp 119-127.

Bruschi, R, Curti, G, Dumitrescu, A, and Vitali, L (1994). "Strength Criteria for Hot Pipelines Susceptible to Euler-Bar Buckling", Proceedings of the Fourth International Offshore and Polar Engineering Conference (1SOPE), Vol. II, Osaka, Japan, April 10-15, 1994, pp 135-148.

DNV OS-F101 (2000), "Submarine Pipeline Systems 2000", Det Norske Veritas Offshore Standard OS-F I 01, Norway, pp !-204.

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