key issues related to modelling of internal corrosion of oil and gas pipelines–a review

8/13/2019 Key issues related to modelling of internal corrosion of oil and gas pipelinesA review

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Experimental investigation on gasliquid two-phase slug ow enhancedcarbon dioxide corrosion in vertical upward pipeline

Donghong Zheng, Defu Che * , Yinhe LiuState Key Laboratory of Multiphase Flow in Power Engineering, Xian Jiaotong University, Xian 710049, China

a r t i c l e i n f o

Article history:Received 2 January 2007Accepted 3 August 2008Available online 22 August 2008

Keywords:C. Slug owC. Mass transferC. Wall shear stressC. Carbon dioxide corrosionC. Corrosion rate

a b s t r a c t

The carbon dioxide corrosion behavior of API N80 grade steel enhanced by gasliquid two-phase verticalupward slug ow has been both mechanistically and experimentally investigated. It is found that thehydrodynamic characteristics of slug ow, such as the direction alternated wall shear stress, the uctu-ation of wall normal stress, and the mass transfer near the wall, have signicant effects on the carbondioxide corrosion process. It is difcult to form dense, compact, and protective corrosion product lmin the corrosion process, which is dominated by general corrosion, and can develop into pitting and mesaattack due to localized corrosion. An empirical correlation is suggested to predict slug ow enhanced car-bon dioxide corrosion. It is found that the mass transfer in corrosion product lm can be neglected andthe slug ow enhanced carbon dioxide corrosion is dominantly controlled by mass transfer or by both of mass transfer and electrochemical corrosion reaction.

2008 Elsevier Ltd. All rights reserved.

1. Introduction

An important consideration in oil and gas industry is multi-phase transport from remote wells for much more economicaltransport of oil and gas combined. The multiphase transport pipe-lines are mostly made of carbon steel and low-alloy steel, whichare able to meet many of the mechanical, structural, fabricationrequirements and may offer considerable capital savings over themore expensive alloys [1,2] . However, the frequently encounteredmultiphase uids may contain signicant levels of CO 2, which incombination with free water can make the pipeline environmentseriously corrosive, resulting in the ow enhanced CO 2 corrosion,thus the damage to the interior of carbon steel pipeline walls,and the decrease in pipeline lifetime and even possible shut downof the pipeline. Over the last few decades, with the development of the crude oil exploitation technology and long distance multiphasetransport technology, the problem of ow enhanced CO 2 corrosionhas made the economic losses increasingly serious, and it is espe-cially true for the oil and gas multiphase transport in the offshoreand deep-sea oil elds.

Corrosion is a surface damage process. Therefore, what is goingon at the metal surface has a profound effect on the corrosion [3] .Many aspects of uid dynamics related to or determined by theinteractions between uid and metal surface are of importanceto corrosion. The changes in uid hydrodynamics, turbulence, wallshear stress, mass transfer, electrochemical corrosion, the forma-

tion and the destruction of the corrosion product lm, are all inti-mately related to the hydrodynamic boundary layer and thediffusion boundary layer in the vicinity of the metal substrate. Ithas been found that metallurgical characteristics of pipelines andhydrodynamic characteristics of ow are important factors inu-encing corrosion rate.

There are remarkably different hydrodynamic characteristicsfor different ow regimes, resulting in different mechanisms of ow enhanced corrosion. Among the gasliquid two-phase ow re-gimes, the vertical upward slug ow is highlighted potentially asthe most aggressive and commonly encountered regime, which al-ways exists in oil and gas pipelines with higher oil and gas produc-tion rates, and causes a high corrosion rate of carbon steel [4,5] .The remarkable hydrodynamic characteristic of the slug ow isthe ow intermittence, that is, the ow is pseudo-periodicallyalternated by Taylor elongated bubble with an annular falling li-quid lm around it and a portion of succeeding liquid slug, simplydescribed as a sequence of slug unit. The Taylor bubble occupiesnearly the whole cross-section of the tube with a bullet-shapedfront and a at tail prole, while the annular falling liquid lm isassimilated by the succeeding liquid slug entraining many smallbubbles. Although slug ow appears a well-ordered ow, it ishighly complicated with an unsteady nature, which is inuencedby many factors, such as the velocities of gas and liquid, void frac-tion, pressure shock, density wave, liquid slug frequency and thephysical properties of gas and liquid.

In this paper, using the limiting diffusion current technology,conductivity probe technology and digital high-speed video sys-tem, mass-loss method, electrochemical impedance spectroscopy

0010-938X/$ - see front matter 2008 Elsevier Ltd. All rights reserved.doi:10.1016/j.corsci.2008.08.006

* Corresponding author. Tel.: +86 29 82665185; fax: +86 29 82668703.E-mail address: [email protected] (D. Che).

Corrosion Science 50 (2008) 30053020

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(EIS), linear polarization resistance (LPR), scanning electronmicroscopy (SEM), and X-ray diffraction (XRD), the vertical upwardslug ow enhanced CO 2 corrosion is investigated both experimen-tally and mechanistically. The major characteristics investigatedinclude the shear stress and mass transfer coefcient in the nearwall zone, the effect of hydrodynamic characteristics on thecorrosion process, CO 2 corrosion rate, the morphological and com-

position analysis of corrosion product lm, the coupling character-istics of hydrodynamics and electrochemical corrosion, etc.

2. Experimental

2.1. Experimental set-up

The experiments are conducted in a two-phase ow set-upmanufactured from 316L (UNS S31603) stainless steel, which isshown in Fig. 1. The detailed description of the experimental set-up has been reported in the previous paper of the authors [6] .

2.2. Hydrodynamic test section

The hydrodynamic test section is a vertical 35 mm inside diam-eter, 5 m long Plexiglas pipe, schematically shown in Fig. 2. Theexperimental liquid phase is the synthetic electrolyte of 0.5 mol/L NaOH0.01 mol/L K 4Fe(CN)60.01 mol/L K 3Fe(CN)6 by dissolvinganalytical-reagents in distilled water, and the experimental gasphase is N 2.

The detailed description of the hydrodynamic measurementprinciples has been reported [6,7] . The pressure differential witha distance of 1570 mm is measured by the 1151DP type capaci-tance differential pressure transmitter. A pressure transmitter isinstalled downstream to measure the local absolute pressure. Wallmass transfer probe and wall shear stress probe are used to mea-sure the mass transfer coefcient and the shear stress near the wallzone using the limiting diffusion current technology. A couple of single-sensor conductivity needle probes with the same geometri-cal structure are used to measure the average void fraction, localbubble size, slug length distribution, slug frequency, Taylor bubblevelocity, etc. A digital high-speed video system is installed down-stream of test section to directly measure two-phase ow charac-teristics by shooting the photographs of the ow at a high-speed.An armored copperconstantan thermocouple is installed to mea-sure the uid temperature. The temperature and pressure are

maintained constant at 40 C and 0.2 MPa throughout theexperiments.

2.3. Electrochemical corrosion test section

The electrochemical corrosion test section is a vertical 35 mminside diameter, 5 m long 316L stainless steel tube, schematically

shown in Fig. 3. The experimental liquid phase is the syntheticsolution of 0.5 mol/L NaCl by dissolving analytical-reagent in dis-tilled water, and the experimental gas phase is CO 2 . The tempera-ture and pressure are maintained constant at 68 C and 0.21 MPathroughout the experiments. The pH value of solution is around4.5 and the conductivity is about 5.1 X 1 m 1. Corrosion ratesare recorded using electrochemical impedance spectroscopy (EIS)and linear polarization resistance (LPR), double-checked withmass-loss method. The pH is monitored throughout the experi-ments using standard litmus paper.

To measure the electrochemical corrosion characteristics in situusing EIS and LPR, four sets of electrochemical corrosion measure-ment devices with the same three-electrode conguration areush-mounted to the wall of the pipe to minimize ow distur-bances, as shown in Fig. 4. All coupons are machined from the par-ent tubular section into cylindrical specimens with 10 mm indiameter, 6 mm in thickness. All of the surfaces are coated withepoxy, except leaving a bottom surface exposed to the corrosivesolution with an area of 0.785 cm 2 . In each three-electrode mea-surement device, the working electrode (WE) is made of API N80mild steel, the counter electrode (CE) and reference electrode(RE) are all made of 316L stainless steel. The martensitic micro-structure of sampled N80 mild steel indicates that this materialhas been quenched and tempered. The chemical compositions of N80 mild steel and 316L stainless steel are presented in Tables 1and 2 .

Before each test run, anaerobic grade CO 2 gas is used to deoxy-genate the solution, lowering the dissolved oxygen level to lessthan 30 l g/L from the formation of oxides. Moreover, the dissolvediron level is controlled below 15 mg/L, avoiding the saturated ironion effect on the corrosion process. During the experiments, theoxygen and iron ion levels are regularly monitored. The couponsare rstly polished progressively with 400-, 800- and 1200- grit sil-icon carbide abrasive paper, then degreased with acetone andrinsed with alcohol many times, dried, weighed, and nallyush-mounted. At the end of the electrochemical corrosion mea-

Fig. 1. Schematic diagram of experimental set-up: (1) gas storage vessel; (2) gas supplement device; (3) freezing dryer; (4) liquid supplement tank; (5) secondary separator;

(6) secondary pump; (7) liquid storage tank; (8) oil bath heater; (9) primary pump; (10) liquid phase orice owmeter; (11) twin-screw compressor; (12) gas surge tank; (13)gas phase orice owmeter; (14) two-phase mixing device; (15) test section; (16) non-return valve; (17) cut off valve.

3006 D. Zheng et al. / Corrosion Science 50 (2008) 30053020


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surements in situ, these coupons are carefully disassembled fromthe test section, rinsed, dried, weighed, stored in a glove bag andsealed in helium gas environment, preventing the corrosion prod-uct from oxidization and destruction, to be further morphologicallyanalyzed.

The electrochemical corrosion measurements are conductedusing an EG&G Princeton Applied Research (PAR) Potentiostat/Gal-vanostat Model 273A and a 5210 lock-in-amplier. The EIS mea-surements are carried out at the open-circuit potential withamplitude of 10 mV sinusoidal alternating current (AC) potentialin the sweep frequency range of 10 mHz to 10 kHz. The acquiredspectra are analyzed based on the principle of an equivalent circuitusing nonlinear least-squares curve tting algorithms provided by

the accompanying software. The LPR measurements are performedby polarizing the working electrode in the range of 20 mV fromopen-circuit potential with a scan rate of 0.1 mV/s, and the auto-matic IR drop compensation is used. In calculating corrosion ratesfrom linear polarization measurements, anodic and cathodic Tafelslopes are assumed to be 120 mV/dec. At the end of experiments,the potentiodynamic sweeps for Tafel polarization curves are doneat direct current (DC) potential range 500 mV below and 300 mVover the open-circuit potential with a scan rate of 0.1 mV/s.

3. Results and discussion

3.1. Hydrodynamic characteristics

There exist four processes in the ow enhanced corrosion as fol-lows [3,8,9] :

1. The formation of reactants in the bulk solution.2. The reactants transport from the bulk solution to the metal

surface.3. Electrochemical corrosion reaction at the metal/solution

interface.4. The corrosion product transport from the metal/solution inter-

face to the bulk solution.

Since electrochemical corrosion occurs on the metal surface, thehydrodynamic boundary layer and diffusion boundary layer exerta pivotal effect on the ow enhanced corrosion [3,10] . Being deci-sive for the ow enhanced corrosion, the extremely complicatedboundary must be simplied here. Starting from the wall, the

hydrodynamic boundary layer can be further subdivided into fourregions: a diffusion sublayer, a laminar viscous sublayer, an inter-mediate buffer layer, and an outer logarithmic turbulent layer. Inthe innermost diffusion sublayer, the mass is dominantly trans-ported by molecular diffusion as a result of concentration gradient.In the laminar viscous sublayer, the velocity and turbulent stressesof the ow are damped to zero, leaves only the viscous stresses of laminar ow to act on the wall; the random movement is themechanism of individual molecules and the mass is transportedby convection diffusion. In this sublayer, the laminar molecularviscous forces dominate over turbulent forces and the ow isessentially laminar. In the outer logarithmic turbulent layer, turbu-lence plays a signicant role, and mass is transported by turbulenteddy diffusion where the random movement of macro-volumes of

solution controls the diffusion. In the intermediate buffer layer, theeffects of both molecular viscosity and turbulence are of the equal

Fig. 2. Schematic diagram of hydrodynamic test section.

D. Zheng et al./ Corrosion Science 50 (2008) 30053020 3007


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importance, and mass is transported by both convection diffusionand turbulent eddy diffusion.

The mass transport in the boundary layer can be measured bymass transfer coefcient k. The viscous energy loss within the tur-bulent boundary layer, namely the intensity of turbulence in theuid acting on the wall, can be directly measured by wall shear

stress sw . The wall pressure of uid perpendicular to the pipe wallcan be measured by wall normal stress r n .

Consequently, the hydrodynamic characteristics, especially thetransport characteristics in the near wall zone, are of signicant ef-fect on corrosion. A higher mass transfer coefcient can cause anincrease in corrosion rate due to the increased transport of bothreactants and corrosion products. A higher wall shear stress canlead to higher corrosion rate by preventing the protective corrosionproduct lm from formation, or by retarding the lm growth, andby removing or destroying existing lm. A higher wall normal

stress uctuations can also exacerbate the corrosion rate by facili-tating the turbulence formation, making the corrosion product por-ous. The failure of the pipelines by ow enhanced corrosion usuallyoriginates from pitting and mesa attack, and is aggravated by mi-cro-turbulence further created by uid ows.

To better understand the slug ow enhanced corrosion, it isnecessary to have a deep insight into the transport characteristicsin the near wall zone. As shown in Fig. 5, the Taylor bubble is alarge axisymmetric bullet-shaped bubble with spherical cap and

Fig. 3. Schematic diagram of electrochemical corrosion test section.

Fig. 4. Schematic diagram of electrochemical three-electrode measurement sys-tem: (1) supporting seat; (2) coupon; (3) seal packing ring; (4) built-in tting; (5)gland nut.

Table 1

Chemical composition of N80 carbon steel (mass%)

C Si Mn P S Cr Mo Ni Nb V Ti Cu Fe

0.24 0.22 1.19 0.013 0.004 0.036 0.021 0.028 0.006 0.017 0.011 0.01 Balance

Table 2

Chemical composition of 316L stainless steel (mass%)

C Si Mn S Cr Mo Ni Fe

0.02 0.80 1.10 0.02 17.00 2.00 14.00 Balance



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nearly at bottom occupying almost the whole cross-section of thetube. Due to the buoyancy, the Taylor bubble pushes aside the li-quid phase toward the tube wall in the Taylor bubble nose zoneat point A, and a falling liquid lm starts to form at point B. The

annular developing falling liquid lm runs downward with an in-creased velocity all the way around the Taylor bubble, and thinsas it falls, until the wall friction force is able to balance the gravi-tational force of the lm and a terminal constant thickness andvelocity of a fully developed falling liquid lm are attained. Whenthe fully developed annular falling liquid lm ows over the Taylorbubble into the wake, the point C, where the annular falling liquidlm plunges into the succeeding liquid slug, encountering the up-ward bulk liquid ow, and its velocity is diminishing until the fall-ing liquid lm is fully assimilated by liquid slug at D. The uidows upward in the zone from the point D to the point A 0 at whichthe succeeding Taylor bubble comes.

Fig. 6 shows the instantaneous wall shear stress, wall normalstress, and mass transfer coefcient proles in slug units at super-

cial velocities U SG = 0.45 m/s, U SL = 0.344 m/s. The acquisitionvoltage U is obtained using the conductivity probe, the wall shear

stress s w is gained using the limiting diffusion current wall shearstress probe, the wall normal stress r n is achieved using pressuretransmitter, and the mass transfer coefcient k is acquired usingthe limiting diffusion current mass transfer probe. The limiting dif-fusion current technology is based on the redox principle of potas-sium ferri- and ferro-cyanide couple. The high level of acquisitionvoltage U represents the Taylor bubble, the corresponding durationtime is attributed to the length of Taylor bubble; the low level of acquisition voltage U represents the liquid slug, the duration timecorresponds to the length of liquid slug. Whereas the high levels of wall shear stress, wall normal stress, and mass transfer are associ-ated with falling liquid lm, the low levels are associated with theliquid slug, and the skew lines are associated with the transitionfrom one zone to another. By comparing the proles of acquisitionvoltage, wall shear stress, wall normal stress, and mass transfercoefcient, the effects of the Taylor bubble nose zone, the Taylorbubble wake zone and the penetration length of the falling liquidlm in the liquid slug on the wall shear stress and the mass trans-fer coefcient can also be obtained. As shown in Fig. 6, althoughthe transitions of the wall shear stress and the mass transfer coef-

cient from the low level to the high level are almost simulta-neous, the transitions lag behind the acquisition voltage,resulting from the existence of LA in the Taylor bubble nose zone;and the transitions of the wall shear stress and the mass transfercoefcient from the high level to the low level are also almostsimultaneous, the transitions also lag behind the acquisition volt-age, resulting from the existence of LC in the Taylor bubble wakezone, as shown in Fig. 5.

From the above analysis of hydrodynamic structure for the slugunit, and from the validation of experimental results, the masstransfer for the slug unit can be classied as four parts as follows:

1. Mass transfer in the Taylor bubble nose zone from A to B withcoefcient kTBn.

2. Mass transfer in the falling liquid lm zone from B to C withcoefcient kTBf .

Fig. 5. Physical description of a slug unit.

0.0 0.2 0.4 0.6 0.8 1.0 1.2

210100

210200

210300

210400

210500

Time (s)

n ( P a )

U SG=0.45 m/sU

SL=0.344 m/s

6

8

10

12

14

16

k f e r r o - c y a n i

d e / 1 0 -

4 (

m / s )

U SG=0.45 m/s U SL=0.344 m/s

-40

-20

0

2040

w / 1 0 ( P a )

U SG=0.45 m/s U SL=0.344 m/s

2

3

4

5

U SG

=0.45 m/s U SL

=0.344 m/s

A c q u i s i

t i o n v o

l t a g e

U ( V )

Fig. 6. Experimental results of wall shear stress and mass transfer coefcient.



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3. Mass transfer in the Taylor bubble wake zone from C to E withcoefcient kTBw .

4. Mass transfer in the remaining liquid slug zone from E to A 0

with coefcient kLS.

Designating the downward wall shear stress positive, the wallshear stress in the slug unit can be divided into two parts as

follows:1. Positive wall shear stress swf from B to D due to the falling

liquid lm, and2. negative wall shear stress swLS from D to B 0 due to the liquid

slug.

Experimental results have been found that both the mass trans-fer coefcient and wall shear stress in each zone increase with theincrease of supercial gas velocity, and decrease with the increaseof supercial liquid velocity. For the falling liquid lm zone, as thesupercial gas velocity is increased, the falling liquid lm terminalvelocity is increased and the falling liquid lm thickness is re-duced, which results in the increases of the velocity gradient andthe concentration gradient, thus increased mass transfer coef-cient and wall shear stress. Whereas as the supercial liquid veloc-ity is increased, the falling liquid lm terminal velocity isdecreased and the falling liquid lm thickness is increased, whichresults in the decreases of the velocity gradient and the concentra-tion gradient. For the Taylor bubble wake zone, increasing thesupercial gas velocity leads to an increased falling liquid lm ter-minal velocity, resulting in more turbulent, disorderly and chaoticwake, and the mass transfer coefcient is increased; the increasedsupercial liquid velocity leads to a decreased mass transfer coef-cient. For the liquid slug zone and the Taylor bubble nose zone,the increase of the supercial gas velocity gives rise to the increaseof void fraction, resulting in increased probability of bubble impacton the wall and a larger turbulence of liquid slug, thus increasedmass transfer coefcient and wall shear stress; whereas a contrarychange is enhanced by increasing the supercial liquid velocity. Itcan also been found that the mass transfer coefcient kTBf is max-imum, the mass transfer coefcient kTBn is minimum; the magni-tude of the wall shear stress associated s wf with the falling liquidlm is always larger than that of swLS associated with the liquidslug.

Due to the complicated hydrodynamic characteristics of slugow, there is still no widely applied normalized formula valid toaccurate determination of the mass transfer coefcient and wallshear stress.

Mass transfer prediction is of great importance in computingcorrosion rates, particularly for the transport based models, andis also the key for the coupling of CO 2 corrosion and hydrodynam-ics of slug ow. The global formula for mass transfer prediction in

fully developed single phase liquid full-pipe turbulent ow can beexpressed as

Sh a Reb Sc c 1

where a = 0.023, b = 0.8 and c = 0.33 for well-known fully developedsingle phase liquid full-pipe turbulent ow correlation developedby Chilton and Colburn [11] . Sh is the Sherwood number kL/Dd ,Re the Reynolds number qUL/l , and Sc the Schmidt number l /(qDd).

Based on the single phase mass transfer coefcient formula, acorrected normalized mass transfer coefcient formula for gasli-quid vertical upward slug ow is suggested as follows:

Sh C 0 :023 Re0 :8 Sc 0:33 2

namely

kTP 0 :023 CRe0:8 Sc 0 :33

DdL 0 :023 C qULl

0 :8 lqDd

0 :33 DdL

3

C 0 :27 1 Fr s

b 0 :5

4

where C is the corrected factor considering the difference of gasli-

quid slug ow from the single phase ow; kTP the mass transfercoefcient (m/s) in gasliquid two-phase slug ow; Dd the diffusioncoefcient (m 2 /s), for the hydrogen ion (H +) Dd = 9.31 10 9 m 2/s,for the ferrous ion (Fe 2+ ) Dd = 7.12 10 10 m 2 /s; L the characteristiclength (m); U the characteristic velocity (m/s); l the characteristicdynamic viscosity (Pa s); q the characteristic density (kg/m 3); Fr s,(U SL + U SG)/( gD)0.5 the Froude number for the mixture velocity of supercial gas phase velocity U SG and supercial liquid velocityU SL ; and b the Taylor bubble length fraction, ratio of Taylor bubblelength LTB to slug unit length LSU = LTB + LLS.

The experimental results show the falling liquid formation dis-tance from the Taylor bubble nose LA is short, and the correspond-ing mass transfer coefcient kTBn is minimum. Therefore, kTBn canbe neglected. The proposed expressions for calculating other masstransfer coefcient are shown in Table 3 . Good agreement can befound with maximum derivation within 10%.

In Table 3 , q L and l L are the physical properties for liquid phase,q G and l G the physical properties for gas phase, D the diameter of the pipe, haTBi the average Taylor bubble void fraction, ha LSi theaverage liquid slug void fraction, hU axial i the average axial velocityin the Taylor bubble wake zone, the calculation can be referred to[6,7] , g the gravitational acceleration.

The average mass transfer coefcient in a slug unit can be cal-culated as follows:

K Xki f i 5 where k i is the mass transfer coefcient kTBf , kTBw , and kLS in eachpart of a slug unit, f i the weight factor based on the length fractionof each part, LTB/LSU, LB/LSU, (LLS LB)/LSU, respectively.

Fig. 7 shows the average mass transfer coefcient of H + for ver-tical upward gasliquid slug ow at different supercial gas and li-quid velocities.

The wall shear stress can be calculated as follows:

sw f qU 2

2 6

where f = 0.046 Re 0.2 is wall friction factor, Re = q LU /l , U the char-acteristic velocity (m/s), L the characteristic length (m), q the char-acteristic density (kg/m 3), l the dynamic viscosity (Pa s).

The proposed expressions for calculating the wall shear stresss wf associated with falling liquid lm and the wall shear stress s wLSassociated with liquid slug are listed in Table 4 . Good agreementcan be found with maximum derivation within 10%.

In Table 4 , U GLS and U LLS are the gas bubble velocity and liquidvelocity in liquid slug, respectively.

Leaving out the directions of wall shear stress, the average wallshear stress in a slug unit can be calculated as follows:

sw jswf j b j swLS j 1 b 7

Table 3

The proposed expressions

kTBf kTBw kLS

U (m/s) 9.916[ gD(1 ha TBi 0.5 )]0.5 hU axial i U SG + U SL L (m) 4 0.5 D(1 ha TBi 0.5 ) D Dq (kg/m 3) qL qL (1 ha LSi ) + q Gha LSi qL (1 ha LSi )

+ q GhaLSil (kg/(ms)) l L l L (1 ha LSi ) + l Gha LSi l L (1 ha LSi )

+ l GhaLSi



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Fig. 8 shows the average wall shear stress for vertical upwardgasliquid slug ow at different supercial gas and liquidvelocities.

The wall normal stress r n , which represents the wall pressure of uid perpendicular to the pipe wall, is constant with small uctu-ations when the only-liquid ows in a straight smooth pipeline. Insuch case, wall normal stress has a weak effect on the corrosion.However, in the gasliquid two-phase slug ow, the larger wallpressure uctuations are caused by the pseudo-periodical inter-mittence of Taylor bubble and liquid slug, exerting a large addi-tional mechanical force onto the metal surface. Thereby, thelarger wall pressure uctuation is a very important parameter con-sidered in the slug ow enhanced corrosion. Fig. 9 shows the max-imum wall normal stress uctuation values for vertical upward

gasliquid slug ow. It can be seen that the maximum uctuationsare insensitive to the different supercial gas and liquid velocities.

From the above analysis, it can be seen that the intermittence of Taylor bubble and liquid slug is an essential characteristic. Asshown in Fig. 6, when a Taylor bubble and the succeeding liquidslug sequentially pass through a given cross-section of the tube,the direction of wall shear stress changes twice, which also causeslarger uctuations of both mass transfer coefcient and wall nor-mal stress. The intermittence characteristic is one of the importantfactors for the ow enhanced corrosion. The intermittence of slugow can be characterized by slug frequency F SU,

F SU N SU

D t 8

where N SU is the number of slug units at time interval D t .

Fig. 10 shows the slug frequency distribution at different super-cial gas and liquid velocities. It is indicated that the slug fre-quency is generally above 2 Hz, and the slug frequency increaseswith increased supercial liquid velocity, decreases with increasedsupercial gas velocity.

3.2. CO 2 corrosion characteristics

Fig. 11 shows the corrosion rates measured by LPR for CO 2 cor-rosion of N80 steel at different supercial gas and liquid velocities.

Table 4

The proposed expressions for wall shear stress

swf swLS

U (m/s) 9.916[ gD(1 (haTBi )0.5 )]0.5 U GLS + U LLSL (m) 4 0.5 D[1 (ha TBi )0.5 ] Dq (kg/m 3) qL qL (1 ha LSi ) + q Gha LSil (kg/(m s)) l L l L (1 ha LSi ) + l Gha LSi

0.30 0.35 0.40 0.45 0.50 0.55 0.60 0.65 0.7040

60

80

100

120

140

160

180

200

220U

SG=0.15 m/s

U SG

=0.25 m/s

U SG=0.35 m/sU

SG=0.45 m/s

w ( P a )

U SL

(m/s)

Fig. 8. Average wall shear stress for vertical upward gasliquid slug ow.

0.30 0.35 0.40 0.45 0.50 0.55 0.60 0.65 0.70100

200

300

400

500

600

700

U SG

=0.15 m/sU

SG=0.25 m/s

U SG

=0.35 m/sU

SG=0.45 m/s

M a x

i m u m

n ( P a )

U SL

(m/s)

Fig. 9. Maximum wall normal stress uctuation values for vertical upward gasliquid slug ow.

0.30 0.35 0.40 0.45 0.50 0.55 0.60 0.65 0.7020

25

30

35

40

45

50

55

60U

SG=0.15 m/s

U SG

=0.25 m/sU

SG=0.35 m/s

U SG

=0.45 m/s

k H +

/ 1 0 - 4 ( m / s )

U SL

(m/s)

Fig. 7. Average mass transfer coefcient of H + for vertical upward gasliquid slugow.

0.30 0.35 0.40 0.45 0.50 0.55 0.60 0.65 0.70

2

3

4

5

6U

SG=0.15m/s

U SG

=0.25m/s

U SG=0.35m/s U

SG=0.45m/s

F S U

( H z )

U SL (m/s)

Fig. 10. Slug frequency distribution at different supercial gas and liquid velocities.



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The corrosion rates keep constant within the 30-h exposure to cor-rosion environment in all experiments due to no corrosion productlm. Nevertheless, the corrosion rates remarkably decrease due tothe formation of the corrosion product lm for 5085-h exposures.For instance, under the condition of U SG = 0.45 m/s, U SL = 0.344 m/s,the corrosion rate changes from 19.7 mm/yr at the beginning of thecorrosion experiment to the smallest value of 16.11 mm/yr after75-h exposure. Further exposed to corrosion, under the larger masstransfer and wall shear stress situations, e.g., U SG = 0.45 m/s,U SL = 0.344 m/s, and U SG = 0.45 m/s, U SL = 0.518 m/s, the corrosionrates increase to 22 mm/yr and 16.4 mm/yr after 140-h exposure,respectively, higher than that within the initial 30-h exposure,namely 19.7 mm/yr and 15.3 mm/yr, respectively. Whereas under

smaller mass transfer and wall shear stress situations, e.g., whenU SG = 0.15 m/s, U SL = 0.446 m/s, and U SG = 0.15 m/s, U SL = 0.651 m/s, the corrosion rates decrease continuously from 10.21 mm/yrand 7.9 mm/yr at the initial exposure to 8.51 mm/yr and 5.8 mm/yr at the end of the experiments after 140-h exposure, respectively.In such situations, corrosion product lm has begun to build on thecoupons after 50-h exposure, and the lm steadily grows, makingthe corrosion rate dramatically decrease.

As shown in Fig. 11 , it can be seen that the average corrosionrate reaches its maximum when U SG = 0.45 m/s, U SL = 0.344 m/s,minimum when U SG = 0.15 m/s, U SL = 0.651 m/s in all experiments.According to the hydrodynamic characteristics analysis, as shownin Figs. 7 and 8 , the average mass transfer coefcient and the aver-age wall shear stress are maximum when U SG = 0.45 m/s,

U SL = 0.344 m/s, minimum when U SG = 0.15 m/s, U SL = 0.651 m/s.It is further validated that the hydrodynamic characteristic plays

a key role in the corrosion process, namely larger mass transfercoefcient and larger wall shear stress lead to a higher corrosionrate. Whereas at smaller mass transfer and smaller wall shearstress, the corrosion rate keeps lower and is prone to form the cor-rosion product lm on the coupon.

Figs. 12 and 13 show the potentiodynamic polarization curvesfor CO2 corrosion of N80 steel when U SL = 0.344 m/s andU SL = 0.651 m/s with different supercial gas velocities. It is alsocan be seen that the polarization currents of N80 steel are depen-dent on the near wall hydrodynamic characteristics.

When U SL = 0.344 m/s with U SG = 0.15 m/s and U SG = 0.45 m/s,except 140-h exposure, the cathodic branches of the polarizationcurves are similar for 1-h, 50-h and 70-h exposures. The cathodic

branches of the curves exhibit limiting current, which is a directconsequence of depletion of H + ions near the surface and the slowhydration rate of CO 2. Right below the corrosion potential, thecathodic reaction process is controlled by activation, and the exter-nal power provides extra electrons, making the cathodic reactionaccelerated; whereas farther negative to the corrosion potential,due to the adjacent active species depletion, the diffusion transportof electroactive species from the bulk solution to the electrode sur-face is dominated, and the cathodic reaction is diffusion controlledwith the existence of explicit limiting current. However, for theanodic branch of polarization curve, the current density increaseswith the increase in anodic polarization potential, and unlimitedsupply of iron atoms is available at the interface. The anodic reac-tion rates do not depend on the mass transfer process. Except 75-h

exposure, the anodic branches of polarization curves are similar for1-h, 50-h and 140-h exposures. It is found that the polarization

0 20 40 60 80 100 120 1405

10

15

20

25

30U

SG=0.15 m/s U SG=0.25 m/sU

SG=0.35 m/s U SG=0.45 m/sU

SL=0.344 m/s

C R ( m

m / y )

Time (h)

0 20 40 60 80 100 120 1405

10

15

20

25U

SG=0.15 m/s U SG=0.25 m/sU

SG=0.35 m/s U SG=0.45 m/sU

SL=0.446 m/s

C R ( m

m / y )

Time (h)

0 20 40 60 80 100 120 1405

10

15

20U

SG=0.15 m/s U SG=0.25 m/sU

SG=0.35 m/s U SG=0.45 m/s U

SL=0.518 m/s

C R ( m m

/ y )

Time (h)

0 20 40 60 80 100 120 1405

10

15U

SG=0.15 m/s U SG=0.25 m/sU

SG=0.35 m/s U SG=0.45 m/sU

SL=0.651 m/s

C R ( m m

/ y )

Time (h)

Fig. 11. Corrosion rates measured by LPR for CO 2 corrosion of N80 steel at different supercial gas and liquid velocities.



9/16

curves gradually change from general corrosion to pitting corro-sion. For 1-h and 50-h exposures, the polarization curves of N80show the typical shape of general corrosion; whereas for 75-hexposure, the curves show the pitting type, and the small passivecurrent densities should be attributed to the corrosion productlm, although the lm is not-fully covered and not compact; for140-h exposure, the corrosion product lm is no longer resistantto hydrodynamic characteristics and detaches from the surface.

Table 5 shows the shifting trend of corrosion potential and thecorresponding corrosion current density when U SL = 0.344 m/swith U SG = 0.15 m/s and U SG = 0.45 m/s. Since the corrosion poten-tial increases in the anodic direction with the formation of corro-sion product lm, the corrosion product lm may be considered

to have anodic anticorrosive capability.The polarization characteristics when U SL = 0.651 m/s with

U SG = 0.15 m/s and U SG = 0.45 m/s are similar to those whenU SL = 0.344 m/s. However, due to smaller mass transfer and smallerwall shear stress, a corrosion product lm is formed on the couponwhen U SG = 0.15 m/s, U SL = 0.651 m/s for 50-h, 75-h and 140-hexposures, and the tendency of pitting corrosion is greatest.

Figs. 14 and 15 show the typical Nyquist and Bode plots of EISspectra for CO 2 corrosion of N80 steel at U SL = 0.344 m/s andU SL = 0.651 m/s with different supercial gas velocities. WhenU SG = 0.15 m/s, U SL = 0.344 m/s, two semicircles appear after 75-hexposure, which means a corrosion product lm has been formedon the coupon. The rst semicircle is so small that it merges intothe second semicircle, cannot clearly distinguished from each

other, namely Rf is much smaller than the charge transfer resis-tance Rct . The rst semicircle at higher frequency would be due

to the corrosion product lm because a surface dielectric lm nor-mally has a small time constant and so has a phase angle shift inthe high frequency range. The measured results validate the con-clusion with Rf = 6.3 X cm 2 , C f = 1018 l F/cm 2 , the value of Rf beingone seventh of the Rct . This indicates that the corrosion productlm has non-protective capacity, and further validates the resultsof corrosion process measured by LPR and potentiodynamicpolarization.

As shown in Table 6 , the charged transfer resistance increasesfrom the 1-h exposure to 50-h exposure, and corrosion productlm starts to form on the coupon. After 75-h exposure, the chargedtransfer resistance increases to the maximum with corrosion prod-uct lm covered; however, the corrosion product lm is not-fully

covered and non-compact. The difference for the physical propertyof corrosion product lm from that of the metal substrate, leads tohigh tendency of pitting corrosion. In such a case, the corrosionproduct lm is detached from the metal substrate due to the effectof the hydrodynamics of slug ow. Thereby, the charge transferresistance Rct decreases to the minimum after 140-h exposure.

From the analysis of EIS spectra, it is found that the type of cor-rosion and the kinetics of the corrosion reaction whenU SG = 0.45 m/s, U SL = 0.344 m/s, and U SG = 0.45 m/s, U SL = 0.651 m/s are similar to those when U SG = 0.15 m/s, U SL = 0.344 m/s.

The impedances when U SG = 0.15 m/s, U SL = 0.651 m/s, areshown in Table 6 . Since corrosion product lm starts to build onthe coupons at the beginning of the exposure, the angle shifts inhigher frequency are very marked. The corrosion product lm

resistance Rf is very small, no more than one fth of the chargetransfer resistance Rct , which indicates that the corrosion product

-5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5-1000

-900

-800

-700

-600

-500

-400

-300 1 h 50 h

U SG=0.15 m/s

U SL=0.344 m/s

P o t e n

t i a l

( m V

) v s

3 1 6 L S S

Log Current density (A/cm 2)

-6.0 -5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5-1000

-900

-800

-700

-600

-500

-400

-300 75 h 140 h

U SG=0.15 m/s

U SL=0.344 m/s

P o t e n

t i a l

( m V

) v s

3 1 6 L S S


-5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5-1100

-1000

-900

-800

-700

-600

-500

-400

-300 1 h 50 h

U SG=0.45 m/s

U SL=0.344 m/s

P o t e n t

i a l ( m

V ) v s

3 1 6 L S S


-6.0 -5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5 -2.0-1100

-1000

-900

-800

-700

-600

-500

-400

-300 75 h 140 h

U SG=0.45 m/s

U SL=0.344 m/s

P o t e n t

i a l ( m

V ) v s

3 1 6 L S S


Fig. 12. Potentiodynamic polarization curves for CO 2 corrosion of N80 steel at U SL = 0.344 m/s.



10/16

lm is very porous, its anticorrosion capacity is very poor, despitethe corrosion product lm formation on the coupons at the initialexposure. Results of the polarization measurement and impedancemeasurement show that the corrosion lm has limited protectivecapacity in the vertical upward gasliquid two-phase slug ow en-hanced CO 2 corrosion due to the remarkable effect of the peculiarhydrodynamics.

It is still difcult for a fully covered, compact, and anticorrosivecorrosion product lm to form due to the effect of hydrodynamicsof slug ow. The hydrodynamic characteristics of vertical upwardgasliquid two-phase slug ow govern the corrosion mechanism,and are intimately related to the overall corrosion process. Masstransfer characteristic has an adverse effect on the corrosion rate,whereas the wall shear stress and wall normal stress are responsi-

ble for the nal morphological consequence. Under the peculiarhydrodynamics of pseudo-periodical intermittence of Taylor bub-

ble and liquid slug, the interactions of both the upward and down-ward wall shear stresses, the uctuation of wall normal stressmake the nal morphological consequence of the corrosionextraordinarily serious.

Fig. 16 shows the resultant stress and included angle for wallshear stress and the uctuation of wall normal stress. The high le-vel of acquisition voltage U represents the Taylor bubble; the lowlevel of acquisition voltage U represents the liquid slug. When aTaylor bubble passing by a given cross-section of the pipe, the po-sitive wall shear stress swf changes from 0 Pa to the maximum(swf )max , and the uctuation of the wall normal stress changes fromthe 0 Pa to the maximum ( D r n)max as well. Consequently, theresultant stress of the wall shear stress and the uctuation of wall

normal stress reaches its maximum, and the included angle be-tween the resultant stress and the wall shear stress increases from0 to the maximum (+90 ) as well. When a succeeding liquid slugpasses, the direction of the wall shear stress is changed upward,namely the negative wall shear stress swLS changes from 0 Pa tothe maximum ( swLS)max , and the uctuation of the wall normalstress is changed to 0 Pa again from the maximum ( D r n)max . Assuch, the resultant stress of the wall shear stress and the uctua-tion of wall normal stress changes to the minimum, and the in-cluded angle between the resultant stress and the wall shearstress decreases from +90 to the 0 as well. Due to the existenceof transition from the leading liquid slug to the succeeding Taylorbubble, the direction of the wall shear stress changes with stressstarting to increase, and the uctuation of wall normal stress in-

creases. as such, the included angle rstly changes to 90 , andthen dramatically changes to 0 , further to +90 .

-6.0 -5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5 -2.0-1600

-1400

-1200

-1000

-800

-600

-400

1 h 50 h

U SG=0.15 m/s

U SL=0.651 m/s

P o t e n

t i a l

( m V

) v s

3 1 6 L S S


-6.5 -6.0 -5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5 -2.0-1100

-1000

-900

-800

-700

-600

-500

-400

-300

-200

75 h 140 h

U SG=0.15 m/s

U SL=0.651 m/s

P o

t e n

t i a

l ( m V

) v s

3 1 6 L S S

Log Curent density (A/cm 2)

-6.0 -5.5 -5.0 -4.5 -4.0 -3.5 -3.0 -2.5 -2.0

-1000

-900

-800

-700

-600

-500

-400

-300

1 h 50 h

U SG=0.45 m/s

U SL=0.651 m/s

P o t e n

t i a l

( m V ) v s

3 1 6 L S S


-6 .5 -6 .0 -5 .5 -5 .0 -4 .5 -4.0 -3.5 -3.0 -2.5 -2 .0 -1 .5 -1.0-1000

-900

-800

-700

-600

-500

-400

-300

-200

95 h 140 h

U SG=0.45 m/s

U SL=0.651 m/s

P o t e n

t i a l

( m V ) v s

3 1 6 L S S


Fig. 13. Potentiodynamic polarization curves for CO 2 corrosion of N80 steel at U SL = 0.651 m/s.

Table 5

Corrosion potential vs corrosion current density

U SG = 0.15 m/s, U SL = 0.344 m/s U SG = 0.45 m/s, U SL = 0.344 m/s

Corrosionpotential (mV)(vs 316L SS)

Log corrosioncurrentdensity log i(A/cm 2)

Corrosionpotential (mV)(vs 316L SS)

Log corrosioncurrent densitylog i (A/cm 2)

1 h 600 3.018 630 2.7750 h 580 3.14 615 2.7875 h 540 3.27 560 2.86140 h 560 2.97 581 2.72



11/16

Both experimental results and previous studies by [12,13] showthat the main components of the corrosion product lm are ironcarbonate FeCO 3 and iron carbide Fe 3C. The N80 steel is corrodedaway as Fe 2+ , which can be precipitated out of the solution onthe coupon substrate as FeCO 3 when the concentration of whichexceeds the super-saturation level in the solution. Fe 3C, as the skel-eton of the metal in a classical sense, is the remains on the parentmetal substrate after the corrosion. Corrosion product lm is brit-tle, compared to the metal substrate, which has plastic physicalproperty. The brittle material is sensitive for the higher uidimpingement angle, such as 7590 , and the plastic material is sen-sitive to the smaller uid impingement angle, for example, 2045 .

Due to the higher resultant stress with included angle changingfrom 90 to the +90 accompanied with the frequency more than2 Hz, the hydrodynamics make the bonding between the metalmatrix and corrosion lm diminish, until the corrosion lm is de-stroyed and peeled off, leading to the exposure of partial or entiremetal to the corrosive reactant and occurring localized corrosion orgeneral corrosion.

Fig. 17 shows the corrosion product lm thickness on N80 steelcoupons vs different exposure time. When U SG = 0.45 m/s, U SL =0.344 m/s, and U SG = 0.15 m/s, U SL = 0.344 m/s, and U SG = 0.45 m/s,U SL = 0.651 m/s, the thickness of corrosion product lm rstly in-creases with the increased exposure time, reaching maximum after75-h exposure; and then decreases with increased exposure time.This can be elucidated as follows: the corrosion product lm starts

to form on the coupons, although it is not-fully covered andnon-compact; however, the localized corrosion is initiated due to

higher wall shear stress and wall normal stress; with the propaga-tion of the localized corrosion, the micro-turbulent intensity in-creases dramatically, which sometimes may be ten orders of magnitude higher than that in the smooth pipe; then the localizedcorrosion further aggravates, resulting in higher corrosion rate,more damage to the corrosion product lm, and more pitting.When U SG = 0.15 m/s, U SL = 0.651 m/s, due to the smaller wall shearstress and mass transfer of slug ow, there exists a favorable con-dition for FeCO 3 precipitation. A corrosion product lm starts toform on the coupons, builds-up, and further strengthens. Thismakes a thicker corrosion product lm, and decreases the corro-sion rate.

Fig. 18 shows the average area fraction covered by corrosionproduct lm on N80 coupons. It can be seen that the maximumarea fraction covered by corrosion product lm is no more than60% due to the effect of hydrodynamics of slug ow. Fig. 19 showsthe area fraction vs corrosion depth for N80 steel coupons after140-h exposure. It can be seen that the corrosion is dominatedby general corrosion with small pitting due to higher corrosiondepth with smaller area fraction in each experimental conditions.Under higher wall shear stress and mass transfer of slug ow, thereis higher corrosion rate, as well as more numerous and deeperpittings.

3.3. Coupling characteristics of hydrodynamics and corrosion

It is essential to develop a mechanistic model based on thesimultaneous and interaction of corrosion characteristics and

0

-10

-20

-30

-40

-50

-60

0

-10

-20

-30

-40

-50

0 10 20 30 40 50 600

-10

-20

-30

-40 1 h 50 h 75 h 140 h

U SG=0.15 m/s

U SL

=0.344 m/s

I m ( Z ) ( O h m - c m

2 )

Re(Z ) (Ohm-cm 2)

5.2Hz1.8Hz

1.1Hz

5.3Hz

-2 -1 0 1 2 3 40

20

40

60

80

100

| Z | ( O h m

- c m

2 )

Log Freq (Hz)

0 5 10 15 20 25 300

-5

-10

-15

-20 1 h 50 h 75 h 140 h

U SG=0.45 m/s

U SL

=0.344 m/s

I m ( Z ) ( O h m - c m

2 )

Re(Z ) (Ohm-cm 2)

5.2Hz

8.8Hz

6.1Hz6.2Hz

-2 -1 0 1 2 3 40

10

20

30

40

50

| Z | ( O h m - c m

2 )

Log Freq (Hz)

1 h 50 h 75 h 140 h

U SG=0.15 m/s

U SL=0.344 m/s

P h a s e ( d e g r e e )

1 h 50 h 75 h 140 h

U SG=0.45 m/sU

SL=0.344 m/s

P h a s e

( d e g r e e )

Fig. 14. Nyquist and Bode plots of EIS spectra for CO 2 corrosion of N80 steel at U SL = 0.344 m/s.



12/16

hydrodynamic characteristics, such as the rate equations of both

metal anodic dissolution and oxidants cathodic reductions, trans-port equations of reactants and products. Nevertheless, an effectivemechanistic model incorporating seamlessly all of these factors isnot available currently.

When the corrosion rate is electrochemical reaction controlled,mass transfer characteristic has no effect on the corrosion. How-ever, majority of elds and laboratory investigations have shownthat the corrosion are both controlled in gasliquid two-phase slugow enhanced CO 2 corrosion, namely the mass transfer mecha-nism and the electrochemical reaction mechanism are of equalimportance in determining the corrosion rate. In such a situation,the ow enhanced corrosion is not merely the simple summationof the electrochemical corrosion caused by the corrosive reactantsand the mechanical losses caused by the gasliquid two-phase slug

ow, the two causes are synergistical, namely there is a 1 + 1 > 2effect. There is a strong nonlinear coupling of three factors so that

the contribution of individual cause to corrosion cannot be distin-

guished easily. These three factors are the wall shear stress causedby the turbulence in the hydrodynamic boundary layer near thewall of the pipelines, the mass transfer caused by the turbulencediffusion in the diffusion boundary layer near the wall of the pipe-lines, and the electrochemical corrosion due to the corrosive reac-tants in the gasliquid two-phase ow.

Based on the analysis for the hydrodynamic boundary layer anddiffusion boundary layer in Section 3.1, the charge transport uxthrough the boundary layer will be considered according to theFicks rst law as follows:

i Dd nF d C d x

9

where Dd is the diffusion coefcient of the ionic species (m 2/s), n its

valence, F the Faraday constant (96,500 C), and d C /d x the concentra-tion gradient of the ion of interest.

0

-10

-20

-30

-40

-50

-60

0

-10

-20

-30

-40

-50

-60

0 10 20 30 40 50 600

-10

-20

-30

-40

-50 1 h 50 h 75 h 140 h

U SG=0.15 m/s

U SL=0.651 m/s

I m ( Z ) ( O h m - c m

2 )

Re(Z ) (Ohm-cm 2)

2.4Hz

8.8Hz

3.1Hz

-2 -1 0 1 2 3 40

20

40

60

80

100 1 h 50 h 75 h 140 h

U SG=0.15 m/s

U SL=0.651 m/s

| Z | ( O h m

- c m

2 )

Log Freq (Hz)

0 10 20 30 400

-10

-20

-30

-40 1 h 50 h 95 h 140 h

U SG=0.45 m/s

U SL=0.651 m/s

I m ( Z ) ( O h m - c m

2 )

Re(Z ) (Ohm-cm 2)

2.4Hz

3.1Hz

-2 -1 0 1 2 3 40

20

40

60

80 1 h 50 h 95 h 140 h

U SG=0.45 m/sU

SL=0.651 m/s

| Z | ( O h m - c m

2 )

Log Freq (Hz)

P h a s e ( d e g r e e )

P h a s e

( d e g r e e )

Fig. 15. Nyquist and Bode plots of EIS spectra for CO 2 corrosion of N80 steel at U SL = 0.651 m/s.

Table 6

Equivalent circuit model analysis

U SG = 0.15 m/s, U SL = 0.344 m/s U SG = 0.15 m/s, U SL = 0.651 m/s

Rct (X cm 2) H max ( ) C dl (l F/cm 2) Rct (X cm 2) H max ( ) C dl (l F/cm 2) Rf (X cm 2) C f (l F/cm 2)

1 h 28.61 49 1070 35.11 46.04 1463 2.12 63050 h 37.77 51.8 2342 36.27 56.18 1418 5.39 41775 h 45.12 54.2 3208 40.12 57.72 451 6.13 476140 h 26.08 46.6 1150 49.38 49.16 1345 6.63 572

Note: Rs = 4.1 (X cm 2).



13/16

As shown in Fig. 20 , there are two scenarios for the coupling of hydrodynamic characteristics and electrochemical characteristicsas follows:

1. Coupling without corrosion product lm, and2. coupling with corrosion product lm.

The reduction of H + is the main culprit for the CO 2 corrosion.Hence, only the H + reduction is of interest in the coupling modelfor convenience. On the naked metal substrate surface, for Scenario1 shown in Fig. 20, according to the Nernsts linear concentration

diffusion gradient model, the charge transport ux of H+

can be ex-pressed as

i DH nF C b C m

dd kTP nF C b C m 10

where kTP is the H + mass transfer coefcient (m/s); C b = 10 pH , thebulk concentration, (mol/L); C m the concentration at the metal/solu-tion interface (mol/L); n = 1 for H+, valence; dd the diffusion bound-ary layer thickness (m), which differs from the thickness of

hydrodynamic boundary layer.The Fe 2+ mass transfer has a great effect on the formation of the

corrosion product lm on the metal substrate surface. The corro-sion product lm will be not only the diffusion barrier betweenthe metal substrate and the corrosive medium, but also give riseto a concentration gradient of the electrochemical reactants. Whensuper-saturation concentrations of Fe 2+ and CO 23 next to the metalsubstrate surface are achieved, a corrosion product lm forms, asshown in Scenario 2 of Fig. 20. When the corrosion product lmstarts to build-up on the metal surface, the corrosion kinetics ischanged, the additional mass transport through the corrosion lmmust be also considered. The charge transport ux of H + can be ex-pressed as

i kTP nF C b C f 11

followed by

0.0 0.2 0.4 0.6 0.8 1.0 1.2-90

-60

-30

0

30

60

90

Time (s)

I n c l u d e d

a n g l e

( d e g r e e )

0

100

200

300

400

500

R e s u l

t a n t s t r e s s

( P a )

2

3

4

5

U SG=0.45 m/s

U SL=0.344 m/s

U SG

=0.45 m/sU

SL=0.344 m/s

A c q u

i s i t i o n v o

l t a g e

U ( V )

U SG=0.45 m/s U SL=0.344 m/s

Fig. 16. Resultant stress and included angle for wall shear stress and the uctuationof wall normal stress.

0 20 40 60 80 100 120 140 1600

2

4

6

8

10

12U

SG=0.45 m/s U

SL=0.344 m/s

U SG=0.15m/s U SL=0.344 m/s

U SG

=0.45 m/s U SL

=0.651 m/sU

SG=0.15 m/s U

SL=0.651 m/s

F i l m t h i c k n e s s

( m )

Exposure time (h)

Fig. 17. Corrosion product lm thickness on N80 steel coupons against differentexposure time.

0 20 40 60 80 100 120 140 1600

10

20

30

40

50

60U

SG=0.45 m/s U

SL=0.344 m/s

U SG

=0.15 m/s U SL

=0.344 m/sU

SG=0.45 m/s U

SL=0.651 m/s

U SG

=0.15 m/s U SL

=0.651 m/s

A r e a

f r a c

t i o n c o v e r e

d b y f i l m ( % )

Exposure time (h)

Fig. 18. Area fraction covered by corrosion product lm on N80 steel coupons.

0.10 0.12 0.14 0.16 0.18 0.20 0.22 0.24 0.260

5

10

15

20

25

A r e a

f r a c

t i o n

( % )

Corrosion depth (mm)

U SG

=0.15 m/s U SL

=0.344 m/sU

SG=0.25 m/s U

SL=0.344 m/s

Fig. 19. Area fraction of corrosion depth for N80 steel coupons after 140-hexposure.



14/16

i kf nF C f C m 12

where C f is the H + concentration on the corrosion product lm/solu-tion interface (mol/L), C m the H + concentration on the metal/solu-tion interface (mol/L), kf the mass transfer coefcient in corrosionproduct lm (m/s).

In Eqs. (10) and (12) , when the cathodic current is limited bythe transport of H + to the surface, then the concentration at thesurface is zero, C m = 0, namely the electrochemical reactions arecontrolled by the mass transfer of H +. This is valid solely for masstransfer-controlled corrosion.

There has been no advanced method to accurately determinethe H + concentration C m at the metal/solution interface so that itis hard to take into account the coupling of hydrodynamics inthe practical corrosion process, especially when the corrosion iscontrolled by electrochemical reaction or by both of mass transferand electrochemical reaction.

Double-checked by mass-loss method, based on the above anal-ysis on Scenarios 1 and 2, as shown in Fig. 20, and based on the for-mula for the charge transport ux of H + controlled by the masstransfer, an empirical correlation for the practical charge transfer

transport ux of H + is suggested to model slug ow enhancedCO2 corrosion as follows:

i P kTPt nFC b 13 1

kTPt

1

kTP

1

kf 14

where kTPt is the total mass transfer coefcient of H +; G the empir-ical coefcient, 0 < G < 1 for the corrosion controlled by electro-chemical reaction, G = 1 for the corrosion controlled by masstransfer, and G > 1 for the corrosion controlled by both; Eq. (14) onlyhas a physical rather than mathematical meaning in Scenario 1without corrosion product lm, namely kTPt = kTP and kf = 0.

Fig. 21 shows the total mass transfer coefcient kTPt and themass transfer coefcient kf of H+. It can be seen that the value of k

TPt rstly decreases within the 95-h exposure and then increases

after the 95-h exposure when U SG = 0.45 m/s, U SL = 0.344 m/s, andU SG = 0.15 m/s, U SL = 0.344 m/s, and U SG = 0.45 m/s U SG = 0.651 m/s; the value of kTPt reduces continuously when U SG = 0.15 m/s,U SG = 0.651 m/s. Nevertheless, the average value of kTPt at differentsupercial gas and liquid velocities is approximately equal to the

0 20 40 60 80 100 120 140 16010

20

30

40

50

60

70

U SG

=0.45 m/s U SL

=0.344 m/s U SG

=0.15 m/s U SL

=0.344 m/s U SG

=0.45 m/s U SL

=0.651 m/s U SG

=0.15 m/s U SL

=0.651 m/s

k T P t

/ 1 0 -

4 (

m / s )

Exposure time (h)

0 20 40 60 80 100 120 140 1600

100

200

300

1000

1500

2000

2500

k f / 1 0 -

4 (

m / s )

Exposure time (h)

Fig. 21. Total mass transfer coefcient kTPt and mass transfer coefcient in corrosion product lm kf of H+.

2 3H CO

-

3HCO

2-

3CO

+H

-OH

bC

f C

mC

Fig. 20. Schematics of mass transfer between bulk solution and metal surface.

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value kTP calculated based on the hydrodynamics of slug ow, asshown in Fig. 7. This means that the total mass transfer coefcientkTPt of H+ is dominated by kTP rather than kf . Hausler [14,15] foundthat the mass transfer coefcient kf in the corrosion product lm isdirectly proportional to the permeability of the lm and is inver-sely proportional to the thickness of the lm. From the analysison the corrosion product lm in Section 3.2 , it can been seen that

the thickness of the corrosion product lm in the micro-dimension,much thinner than the thickness of hydrodynamic boundary layer;moreover, the maximum area fraction covered by corrosion prod-uct lm is no more than 60%. Due to the effect of hydrodynamics of slug ow, the porous, non-compact, and not-fully covered corro-sion product lm has poor protection, which can be further vali-dated by the value kf . As shown in Fig. 21 , the value of kf is muchhigher than kTPt and kTP, as high as 15 times kTP when U SG =0.15 m/s, U SL = 0.344 m/s, and U SG = 0.45 m/s, U SL = 0.651 m/s, andU SG = 0.15 m/s, U SL = 0.651 m/s, even as high as 43 times kTP whenU SG = 0.45 m/s, U SL = 0.344 m/s. This indicates that the mass trans-fer coefcient kf in the corrosion product lm can be neglected incalculating the corrosion rates.

Ikeda et al. [16] found that there is a qualitative change in thecorrosion kinetics at temperatures around 60 C. Below 60 C, thecorrosion rate increases with increased temperature, accompaniedwith the formation of the non-protective corrosion product lmiron carbide Fe 3C; from 60 C to 90 C, the corrosion rate reachesits maximum with increased temperature, accompanied with theformation of the protective product lm iron carbonate FeCO 3 ;above 90 C, the corrosion rate decreases with increased tempera-ture due to the protective iron oxide Fe 2O3, which becomes themain component of the product layer. Therefore, at the experimen-tal temperature of 68 C, the corrosion rate is high, and the corro-sion process is controlled mostly by mass transfer or by both,which can also be validated by empirical coefcient G.

Fig. 22 shows the empirical coefcient G, it can be seen that itincreases with the increased exposure time throughout the exper-iments, which means that the corrosion process is initially con-trolled by electrochemical reaction, then by mass transfer, andfurther by both. At the initial exposure, there is no corrosionproduct lm on the coupons, and the empirical coefcient is be-tween 0.9 and 1, which further validates the reliability of themodel. With increased exposure time, due to the precipitationof corrosion product lm composed of iron carbonate FeCO 3and iron carbide Fe 3C, and due to the electronic conductibilityof Fe3C, the galvanic corrosion occurs with high potential, leading

to high localized corrosion, further propagating to pitting corro-sion and mesa attack; In such conditions, the localized corrosioncan bring forth the larger roughness of the wall, making the localturbulent intensity dramatically rise, compared to that in thesmooth pipe. Due to the difculty in taking into account the tur-bulent intensity in the localized corrosion with macro-parametersin a smooth pipe, a large deviation arises, which makes the

empirical coefcient G

larger than 1, even reach 1.5 whenU SG = 0.45 m/s, U SL = 0.344 m/s. Therefore, it is indicated that theslug ow enhanced CO 2 corrosion is dominantly controlled bymass transfer or by both.

4. Conclusions

The gasliquid slug ow enhanced carbon dioxide corrosionof API N80 grade steel in vertical upward pipeline has beenexperimentally investigated. The following conclusions can bedrawn:

1. Gasliquid two-phase vertical upward slug ow has peculiarhydrodynamics. When both the Taylor bubble and the succeed-ing liquid slug of a slug unit sequentially pass through a givencross-section of the tube, the direction of wall shear stress ischanged twice, which causes larger uctuations of both masstransfer coefcient and wall normal stress as well. The fre-quency of the pseudo-periodical alternation by Taylor bubbleand succeeding liquid slug is generally above 2 Hz.

2. Experimental results validate that the hydrodynamics of slugow have signicant effects on the CO 2 corrosion. A higher masstransfer coefcient can cause the increase of corrosion ratesince the transports of both reactant and corrosion productare increased. The higher wall shear stress can lead to highercorrosion rate by preventing the protective corrosion productlm from formation, or by retarding the lm growth, and byremoving or destroying existing lm. The higher wall normal

stress uctuations can also exacerbate the corrosion due tothe turbulence, making the corrosion product porous.

3. Within the initial 30-h exposure, the corrosion rate keeps con-stant; within the 5085-h exposure, the corrosion ratedecreases due to the formation of corrosion product lm; if the exposure time is further increased, due to the damage of corrosion product lm, accompanied with the galvanic corro-sion and pitting corrosion, the corrosion rate is increased, evenlarger than that at the initial exposure. The hydrodynamic char-acteristic plays a key role in the corrosion process, larger masstransfer coefcient and/or larger wall shear stress lead to ahigher corrosion rate, smaller mass transfer and smaller wallshear stress lead to a lower corrosion rate and is prone to formthe corrosion product lm on the coupons.

4. A not-fully covered, non-compact, non-protective corrosionproduct lm forms on the coupon due to the effect of hydrody-namics of vertical upward slug ow. It is found that the thick-ness of the corrosion product lm is no more than 12 l m, thearea fraction covered by corrosion product lm is no more than60%. The slug ow enhanced CO 2 corrosion is dominated bygeneral corrosion, and the corrosion can further develop intopitting and mesa attack due to localized corrosion.

5. Based on the formula for the corrosion process controlled bymass transfer, an empirical correlation is suggested to modelslug ow enhanced CO 2 corrosion. Because that there onlyexists a non-protective corrosion product lm due to the effectof hydrodynamics of slug ow, the mass transfer in corrosionproduct lm can be neglected. The slug ow enhanced CO 2 cor-

rosion is dominantly controlled by mass transfer or by both of mass transfer and electrochemical corrosion reaction.

0 20 40 60 80 100 120 140 1600.8

0.9

1.0

1.1

1.2

1.3

1.4

1.5

1.6U

SG=0.45 m/s U

SL=0.344 m/s

U SG

=0.15 m/s U SL

=0.344 m/s

U SG=0.45 m/s U SL=0.651 m/sU

SG=0.15 m/s U

SL=0.651 m/s

Expsoure time (h)

Fig. 22. Values for the empirical coefcient G .



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Acknowledgments

The nancial supports from the Natural Science Fund of China(50231020, 10372077) are gratefully acknowledged. The valuablediscussions with Prof. Lin HE of Material Science and EngineeringSchool of Xian Jiaotong University, Prof. Minxu LU of Material Sci-ence and Engineering School of Beijing Science and Technology

University, Dr. Zhenquan BAI of Tubular Goods Research Centerof China National Petroleum Corporation, and the kind help fromthem are also gratefully acknowledged.

References

[1] M.B. Kermani, D. Harrop, The impact of corrosion on the oil and gas industry, J.SPE Prod. Facilities 8 (1996) 186190.

[2] M.B. Kermani, A. Morshed, Carbon dioxide corrosion in oil and gas production a compendium, Corrosion 59 (2003) 659683.

[3] K.D. Efrd, Flow-induced corrosion, in: R. Winston Revie (Ed.), UhligsCorrosion Handbook, second ed., John Wiley & Sons, Inc., 2000, pp. 233248.

[4] Z.S. Mao, A.E. Dukler, Rise velocity of a Taylor bubble in a train of such bubblesin a owing liquid, Chem. Eng. Sci. 40 (1985) 21582160.

[5] W.P. Jepson, Modeling the transition to slug ow in horizontal conduit, Can. J.Chem. Eng. 67 (1989) 234.

[6] D.H. Zheng, D.F. Che, Experimental study on hydrodynamic characteristics of upward gasliquid slug ow, Int. J. Multiphase Flow 32 (2006) 11911218.

[7] D.H. Zheng, D.F. Che, An investigation on near wall transport characteristics inan adiabatic upward gasliquid two-phase slug ow, Heat Mass Transfer(2006), doi: 10.1007/s00231-006-0193-8.

[8] E. Dayalan, G. Vani, et al., Modeling CO 2 corrosion of carbon steels, in:Corrosion/1995, NACE International, 1995, Paper No. 118.

[9] S. Nesic, L. Lunde, Carbon dioxide corrosion of carbon steel in two-phase ow,Corrosion 50 (1994) 717727.

[10] D.T. Wasan, C.L. Tien, C.R. Wilke, Theoretical correlation of velocity and eddyviscosity for ow close to a pipe wall, AIChE J. 9 (1963) 567568.

[11] C.H. Chilton, A.P. Colburn, Mass transfer (absorption) coefcient predictionfrom data on heat transfer and uid friction, Ind. Eng. Chem. 26 (1934) 11831187.

[12] C. de Waard, Prediction of CO 2 corrosion of carbon steel, in: Corrosion/1993,NACE International, 1993, Paper No. 69.

[13] R. Zvauya, J.L. Dawson, Electrochemical reduction of carbon dioxide and theeffects of the enzyme carbonic anhydrase 11 on iron corrosion, J. Chem.Technol. Biotechnol. 61 (1994) 319326.

[14] R.H. Hausler, The mechanism of CO 2 corrosion of steels in hot, deep gas wells,Advances in CO 2 Corrosion, vol. 1, NACE, 1984.

[15] R.H. Hausler, D.W. Stegman, CO 2 corrosion and its prevention by chemicalinhibition in oil and gas production, in: Corrosion/1984, NACE, 1984, Paper No.363.

[16] A. Ikeda, M. Ueda, S. Mukai, CO 2 Behavior of Carbon and Cr Steels, Advances inCO2 Corrosion, NACE, 1984.

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key issues related to modelling of internal corrosion of oil and gas pipelines–a review

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