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Inhibitor performance under liquid droplet impingement conditions in CO2-containing environment
Corrosion Science, Vol. 34, No. 8, pp. 1299-1310, 1993 Printed in Great Britain.
0010-938X/93 $6.0tl + 0.00 © 1993 Pergamon Press Ltd
INHIBITOR PERFORMANCE U N D E R LIQUID DROPLET IMPINGEMENT CONDITIONS IN CO2-CONTAINING ENVIRONMENT J. H. GERRETSEN and A . VISSER Shell Research Arnhem, Billiton Research BV, P.O. Box 40, 6800 AA, Arnhem, The Netherlands Abstract--Oil and gas production and transport facilities are often constructed from relatively cheap carbon steel and, usually, inhibitors are then applied to mitigate corrosion, CO~ being a particular threat. A proper inhibitor formulation should reduce the corrosion rate by a factor of at least 10, under actual operating conditions. Severe corrosion may occur when film forming inhibitors are removed from the steel by high liquid shear forces. In gas production, impingement of liquid drolets carried in wet, CO2-containing gas may cause such erosion-corrosion. Current knowledge about the mechanism of flow induced removal of such protective inhibitors films is limited. The performances of commercial inhibitors under impact conditions are generally unknown and appropriate test methods for establishing critical conditions are lacking. The API RP-14E formula is often used for estimating critical erosion-corrosion velocities in uninhibited corrosive systems, but this formula is not applicable to inhibited systems. This paper focuses at correlating droplet impact velocity with performances of inhibitor films. Impact is simulated by collision of rotating steel coupons, mounted on a disc, against a liquid jet. To avoid drying of the samples between two collisions, the samples are also continuously wetted during rotation. A marked difference in performance between two commercial inhibitors was found. One inhibitor gave protection outside, but severe corrosion within the impact zone. The other inhibitor gave exactly the opposite behaviour. The equipment was shown to be capable of demonstrating differences in inhibitor performance as function of type and concentration. Critical impact velocities have been obtained for different inhibitors. The performance of the inhibitors tested were shown to be temperature and concentration dependent. The potential of an inhibitor to regain its protective properties, after failure above a critical velocity, could also be studied. This appeared to be a function of temperature for the two inhibitors tested. In principal, the equipment may be used for establishing critical velocities for the onset of erosioncorrosion under liquid droplet impact conditions for corrosion product films.
INTRODUCTION THE LIMITING p r o d u c t i o n v e l o c i t y b e y o n d w h i c h e r o s i o n - c o r r o s i o n in h i g h g a s p r o d u c t i o n o c c u r s , is u s u a l l y c a l c u l a t e d w i t h t h e A P I R P - 1 4 E f o r m u l a : 1 C Vlimiting erosio n -- ~ / ~ ,
(1)
where C may vary from 100 to 400, depending on the policy of the oil companies. The API formula is known to be inaccurate in many cases, usually underestimating t h e l i m i t i n g v e l o c i t y . T h i s is n o t s u r p r i s i n g as t h e f o r m u l a w a s o r i g i n a l l y d e r i v e d for predicting limiting velocities in the absence of inhibitor films or corrosion product scales. Because of the lack of another predictive tool, the formula has been used for Manuscript received 25 September 1992. 1299
1300
J.H. GERRETSENand A. VISSER
other operating conditions. The validity of the formula has recently been discussed by Smart, 2 who suggested that the physical relevance of the formula is given by a change in flow pattern at the critical velocity, namely from stratified multiphase flow to annular mist flow. In the latter case, damage may occur when the annular mist flow is disturbed and liquid droplets impinge on the surface. Flow disturbances can be caused by many geometries, such as bends, elbows, fittings, but also rough surface films may introduce significant turbulence. Current knowledge about inhibitor films and corrosion product performance under liquid droplet impact conditions is limited. When flow is completely smooth, the liquid shear stress is not likely to exceed the critical strain for failure of surface films and therefore is not high enough to cause damage. However, when flow is severely disturbed, damage to films may occur either by increased shear stress or by increased mass transport. Less severe flow conditions may readily remove films containing defects, prevent corrosion product films from forming, or inhibitor films from adhering. The latter is clear in the case of the application of oil soluble inhibitors and the occurrence of water drop-out. In the case of liquid droplet impact, the mechanism of which is briefly discussed in the next section, local shear stress can be extremely high. In addition to damage done by high local shear, damage may also occur by a fatigue mechanism in the case of repetitive impact. To combat erosion-corrosion in high production facilities, often constructed from carbon steel, inhibitors are applied. Today, inhibitor suppliers design the chemical composition of their inhibitors in such a way that they give a suitable protection against general corrosion under anticipated operating conditions. Their corrosion protective properties are usually judged in simple tests, like wheel/bottle tests. Rotating cylinder tests and even sometimes flow loop (smooth pipe flow) tests are increasingly being used to better establish their performance. Although an inhibitor is most likely to fail at places where flow disturbances occur, namely at bends, elbows and fittings, their persistence under more severe flow conditions is usually not thoroughly examined. Sometimes candidate inhibitors are also tested in the field. These field tests focus on general corrosion, while localized corrosion is the real threat, however. To judge the performance of candidate inhibitors for gas production operations where high flow velocities may occur they should specifically be screened on their performance at places where extreme conditions exist. For the specific conditions of annular mist flow a test facility has been built to examine inhibitor performance. In the test facility liquid droplet impact is being simulated by repetitive collisions of horizontally rotating coupons, mounted on the circumference of a disc, against a continuous vertical water jet. Conditions like temperature, ratio oil/water, type and concentration of inhibitor and water chemistry can be varied.
Droplet impact The theory of liquid droplet impact has been studied extensively in the past and excellent overviews can be found in literature. 3 Basically, the impact event resulting in damage can be divided in two parts. Initially the droplet is compressed only and there is no outward liquid flow. The contact area between droplet and the wall increases with a velocity greater than the compression waves. 4'5 Consequently the liquid in contact with the solid is com-
Inhibitor performance under liquid droplet impingement
1301
impact velocity
T ../ ['
\
,
', i
!
/
I q~ high pressure
v edge
steel
FIG. 1. Liquid droplet impact.
pressed. W h e n the speed of sideways m o v e m e n t b e c o m e s less than the compression wave velocity, pressure is relieved and o u t w a r d flow over the surface begins (Fig. 1).6 It has been pointed out that the liquid actually does not simply spread out over the surface but micro-jetting occurs, when the compression is released and the compressed liquid can expand. 7 T h e highest pressure occurs at the contact ring of the droplet before pressure is released. T h e pressure at the surface in the ring m a y easily be as high as 3 times the water h a m m e r pressure, whereas the velocity of the jets may be as high as 10 times the actual impact velocity. Especially at the edges of the contact ring, fatigue m a y occur. Outside the ring, d a m a g e is mainly due to high shear stresses. T h e water h a m m e r pressure is given by: Pw = plqv
(2)
where t~ (kg m -3) is the density of liquid, q (m s -1) is the velocity of sound in the liquid, v (m s -1) is the impact velocity. U n d e r operating conditions, the m a x i m u m pressure will not exceed the yield strength of C-steel (350 MPa). H e n c e , if metal erosion occurs it can only be caused by a fatigue mechanism. W h e n the angle, 0, between the edge of the compressed droplet and the metal is only a few degrees, the high pressure at the edge is reached, and jetting occurs (Fig. 1). E X P E R I M E N T A L METHOD Impact of liquid particles, as may occur in annular mist flow, is simulated in an experimental set-up, where horizontally rotating coupons hit a vertical water jet (Fig. 2). The samples are positioned in the vertical plane and mounted at the circumference of a disc. The samples collide against a continuous vertical water jet which has a diameter of 2 mm. In this way a cylindrical impact event is studied rather than single droplet impact. The mechanism leading to damage is considered to be similar, however. Preliminary experiments have shown that the samples dried immediately after an impact event, due to the centrifugal forces. Hence, after erosion by impact, corrosion could only take place for a fraction of the rotation time. Therefore a construction has been made to allow for continuous wetting during rotation. The construction consists of a reservoir mounted on the disk. A continuous water supply is arranged to the reservoir. Through nozzles liquid is spread over the samples during rotation. Drying of the samples by
1302
J . H . GERRETSENand A. VISSER
centrifugal forces is thus avoided. The water film over the surface of the samples will cushion the impact. In practice, a continuous liquid film will also always be present on the metal surface. The wheel rotates in a steel vessel which is continuously purged with either N2 or CO 2 containing less than 2 ppm O 2. The O 2 content is measured continuously, using an Orbisphere Oxygen sensor, mounted in the outlet gas stream. The steel vessel is placed on top of another vessel. The upper one is funnel shaped to avoid liquid hold-up. The lower vessel is used for preconditioning the liquid before starting an experiment and it supplies a liquid buffer (25 l) to avoid fast contamination with Fe ions and simultaneously a rapid pH increase. The tests can be operated with a two phase liquid (water/oil) flow. Kerosene, type Shellsol K, has been used as an oil phase. NaCl (1%) was added to the total amount of liquid. The liquid is recirculated continuously. Inhibitor can be injected directly to the liquid feed to the samples. The water/oil phases are separated in the lower vessel and, by means of two pumps, re-mixed just before entering the upper vessel. The actual water/oil ratio can be adjusted via the power of the two pumps. The water can be measured in a by-pass. The impact velocity can be varied between 9 and 50 m s -1. Corrosion is monitored in situ by means of pH and sensitive He measurements. During the experiments liquid samples are taken to measure the Fe concentration by ICP measurements. The pH measurements are carried out in the water phase, before re-mixing. H 2 monitoring takes place at the outlet side of the gas. The actual gas throughput is controlled by means of a Brooks mass flow controller. Samples were machined from steel 52-3, of which the chemical composition is given in Table 1. The samples were ground to 320 grit and degreased with alcohol and weighed before mounting. To avoid any galvanic action, the samples are isolated from the disc by PVDF washers. In view of a construction for in situ electrochemical measurements, not described in this paper, the contact angle between sample and
motor ] I
gas crtttIet to H2 ,fl 02 monitarin~
liquid feed gas inlet co2 / ~2
I sample
I" to conditioning vessel FIG. 2.
Schematic presentation of the jet and wheel test rig.
Inhibitor performance under liquid droplet impingement
1303
liquid jet is not exactly 90 ° but 88 ° . Consequently, the tests discussed here do not represent the worst case, the trends, however, are indicative. At the end of an experiment both the jet and the continuous water feed are shut off. The samples are then 'spin' dried for a few minutes. After releasing the pressure to ambient the samples are taken out and rinsed with kerosene and alcohol. After visual examination, corrosion products are removed using Clark's solution and weight loss is measured. The samples are then microscopically examined. The tests discussed in this paper are summarized in Table 2. Two different inhibitors were subjected to the tests. Inhibitor A is oil soluble and water dispersible. Inhibitor B is oil soluble and is not carried over to the water phase. The inhibitors were tested at 45 and 70°C. Also, base case experiments without inhibitors have been carried out. These tests were carried out at 27 and 45°C, to avoid any disturbances of formation of the solid corrosion products.
TABLE 1. Elements %
CHEMICALCOMPOSITIONOF STEEL 52-3
C
Si
Mn
P
S
Ni
Cr
Cu
0.19
0.44
1.41
0.013
0.027
0.08
0.17
0.15
Supplier: MCB+ Valkenswaard, The Netherlands.
pH
J
Tlme (h)
I
I
I
I
I
I
I
+
10
20
30
40
50
60
70
80
FI6.3.
Increase of pH during experiment 2.
1304
J.H. GERRETSENand A. VtssEa mg/i or mm/y 70.O0 60.00 50.00 40.O0 30.00
Corrosion/ pH
'
" .........
10.O0
.
0.00 3.50
3.70
3.90
4.10
4.30
4.50
4.70
pH
4.90
Fl~. 4. Calculatedrelation between pH and Fe2+ concentration (Ill) (mr V I), 45oc, 2 bar CO2, 1% NaCI. Corrosion rate predicted by the nomograms (o) and measured corrosion rate (A) (ram y-l).
accumufi~te~ Fe (rag)
[~[~
800
X
o~~
t
700 600 500 400
l
.~
200
100 0 4 0
~.20
~
40
60
80 i
"---0--- I.C~
i
meas ]
100
120 Time
1-12
FIG. 5. Fez~- accumulation, calculated from H2 measurements and Fez+ concentration obtained from ICP measurements (el, experiment 3. P = measured corrosion rate. Th = corrosion rate accordingto nomogram.
Inhibitor performance under liquid droplet impingement
mm/y or mg/I 45.00
1305
|7O C / 4 bar CO2/1% NaClal
35,00 30,00 25,00 20,00 i 15,00
.
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10,00 5,00
.
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pH rclation~
0.00
I
3.5
at"
3.7
d
I
I
i
3.9
4.1
4.3
4.5
[
oH 4.7
FZG. 6. CalcuLated relation between pH and Fe 2+ concentration (Ill) (mg 1-1), 70°C, 4 bar CO2, l% NaC1. Corrosion rate predicted by the nomogram s (o) and measured corrosion rate (A) (mm y-l).
accumulated F e (rag) 600
[~~
500
Ichangeto 70 C
I
4OO
200100
~
0 ~
I t0
~
I ol 20
o
P
I 30
4
. Th126 . ~
, 40 o
I.C.P. - -
50
I 60 82 meas
'l 70
! 80 Time (h)
FIG. 7. Fe 2÷ accumulation, calculated from H2 measurements and Fe2÷ concentration obtained from ICP measurements (o), experiment 4. P = measured corrosion rate, Th = corrosion rate according to nomogram, without scaling factor.
90
1306
J.H. GERRETSENand A. VISSER TABLE 2.
Exp.
TEST CONDITIONSAND RESULTS
Impact Inhibitor Water velocity Temp. Inhibitor c o n c . PCO2 cut (m s-l) (°C) type (ppm) (bar) (%)
1 2 3
15 30 10-30
27 45 70
A
100
1 2 4
4 5 6
10-30 10-35 10-30
45-70 45 70
A B B
100 100 100
4 4 4
Remarks
100 Little effect on corrosion rate 100 Little effect on corrosion rate 90 Failurebetween 15 and 20 m s-1, regaining protection at lower velocity 90 Failure,no regaining of protection 90 No failure 90 Bad performance
EXPERIMENTAL RESULTS AND DISCUSSION The test conditions and results are summarized in Table 2. First, tests have been carried out without inhibitors and without kerosene. The samples did not show any preferential attack after these tests. Figure 3 shows the increase of p H during an experiment at 45°C. A change in p H can be correlated with a corrosion rate by calculating the relation between Fe 2+ concentration and p H thermodynamically. This relation is given in Fig. 4. The corrosion rate obtained from the measured p H shift and the corrosion rates as predicted by the 'de W a a r d and Milliams' n o m o g r a m , 8 are also shown in Fig. 4. The corrosion rates showed to be somewhat higher than the corrosion rates expected from the n o m o g r a m , thus possibly indicating a small effect of impact velocity on corrosion rate. A good agreement was obtained between the average corrosion rate calculated from weight loss and from the increase in p H , 12 and 13 m m y-X, respectively. Optically the impact region appeared to be somewhat darker. There was no evidence of metal erosion, as was expected from the impact forces and yield strength of the metal. The amount of dissolved iron remained below the calculated solubility limit of FeCO3. Hence erosion of a product film and subsequent enhanced corrosion could not occur under these conditions. The corrosion rate in experiment 1 at 27°C was also a little higher than expected from the nomogram. Although corrosion rates predicted by the n o m o g r a m have usually been regarded as representing worst case, it seems that corrosion rates can b e c o m e higher than predicted under extreme flow conditions. In the experiments in which inhibitors were studied, the samples were allowed to corrode freely for about 1 or 2 h before injection of inhibitor. In test 3, Fig. 5, carried out at 70°C, the initial corrosion rate was approx. 20 m m y - l , which is below the corrosion rate predicted by the n o m o g r a m , using it without a scale factor (Fig. 6). On injection of the inhibitor, the corrosion rate decreased immediately from about 20 to 0.2 m m y-1. The impact at 10 m s -1 collision speed did not apparently prevent the inhibitor from protecting the steel. W h e n the impact velocity was increased to 3 0 m s -1, the inhibitor failed. The average corrosion rate, determined from H2 measurements, increased to 22 m m y-1. Decreasing the impact velocity resulted in a decrease of the corrosion rate. Apparently this inhibitor is capable of regaining its protection, once damaged. In the next stage of the experiment the velocity was increased step-wise. Figure 5 shows failure of the inhibitor above 20 m s -1. The corrosion rate at 15 m s -1 remained low, therefore it appears that indeed a critical erosion velocity can be obtained for this inhibitor; the corrosion rate being constant
MorphoLogy of samples after Liquid droplet impact, experiments
Continuous water feed
I00 ppm inhib. A 45/70 ° C V impact = variable r
]
profi Le
I00 ppm inhib. B 70oC V impact -- variable i
i
profile FIG. 8.
Photomicrographs of samples from experiment 4(a) and 7(b).
1307
Inhibitor performanceunder liquid droplet impingement
1309
under a critical velocity, above it the efficiency dropping to virtually zero. Other experiments had shown that the critical velocity is dependent on inhibitor concentration. At marginal concentration the inhibitor failed at a much lower velocity. A similar experiment was carried out at 45°C (Fig. 7). It was interesting to observe that in this case the inhibitor film did not recover after being damaged at a velocity of 30 m s -1 (between 48 and 64 h) and subsequent decrease of velocity. Moreover, when the inhibitor film lost its protection, the efficiency was not reduced as drastically as at 70°C and some protection remained. From the results of the experiment carried out at 70°C, it was expected that the inhibitor would regain its protection when the temperature was increased to 70°C. Protection was not regained, however. This effect is not well understood. A speculative explanation may be given by the surface roughening which occurs during the period in which the inhibitor film is damaged. Initially, corrosion will roughen the surface, increasing the surface shear forces significantly. However, at a certain roughness, the effect of the impact will begin to reduce, because the droplets will break-up sooner after the impact, consequently reducing pressure build-up and actual liquid velocity. Considering the corrosion rates during failure, it is likely that during the time of failure at 45°C, roughening of the surface has been less than in the case of the 70°C experiment, hence the effect of the impact is less in the latter case at a similar impact velocity. At 45°C, inhibitor B did not fail at all. Even at an impact velocity of 35 m s -1 , there was no sign of failure. The general corrosion rate was reduced from an initial 11 to about 0.25 mm y-1. Apparently, inhibitor B performed extremely well under these impact conditions. However, similar experiments carried out at 70°C clearly showed again that, in general, a performance of an inhibitor under one set of conditions is certainly not a guarantee for similar behaviour under different conditions. In this case the effect of temperature was even more dramatic. The inhibitor which was very resistant against the liquid droplet impact at 45°C, now performed poorly. After injection of the inhibitor at an impact velocity of 10 m s t, the corrosion rate dropped significantly, but remained always higher than 2 mm y-~, which is clearly unacceptable. These results illustrate that for making a selection from several candidate inhibitors for the actual operations, the inhibitors should be screened at a range of conditions, temperature being an important parameter. Critical erosion velocities can be obtained by using the jet and wheel technique for conditions which are not addressed by the API RP-14E formula. The behaviour of inhibitors under droplet impact conditions can be studied. In a similar manner the mechanical stability of corrosion product layers can be established.
Morphology The morphology of the samples after being exposed to the repetitive impact was intriguing. As can be seen from Fig. 8(a) and (b), from experiments 4 and 6 respectively, inhibitor A has failed preferentially in the area where high pressure occurs; the outer region being well protected. The performance of inhibitor B, on the contrary, proved to be better in the impact region, whilst the film completely failed on the outside. It appears that inhibitor A is less susceptible for high shear under the test conditions, where inhibitor B is less susceptible to fatigue. This phenomenon cannot be explained as yet, but it clearly indicates the complexity of the mechanisms involved in the liquid droplet events. According to
1310
J.H. GERRETSENand A. VISSER
Thomas, 7 the highest pressure should have occurred at a distance d = Rv/q, thus with a diameter, R, of the jet being 2 mm and an impact velocity of 20 m s-1 and assuming the velocity of sound in the liquid, cb being equal to those in water (1500 m s-l), d = 0.027 mm. Figure 8 shows that this region is about 2 mm, however. Such a discrepancy has been observed by others as well. 9 In our experiment, cushioning of the impact by the continuous liquid film at the surface may have increased the distance outside of which high liquid shear exists. Distortion of the jet by the aerodynamics in the vessel may also increase the effective distance. Another factor involved is instability of the jet. The distance, d, will also increase when oscillation of the jet occurs. CONCLUSIONS (1) The corrosion rate under liquid droplet impact conditions can be higher than expected from the De Waard and Milliams nomogram. (2) Inhibitor performance under simulated liquid droplet conditions depends strongly on temperature. (3) Recovery of inhibitor performances, once adversely affected, does not always occur instantaneously. Temperature was shown to be an important parameter here as well. (4) The morphology of corrosion under liquid droplet impact depends on the chemical composition of the inhibitor; one inhibitor may be susceptible to high shear, whereas another may be susceptible to damage by pressure. (5) The jet and wheel test facility allows for inhibitor screening tests for gas production facilities. Critical velocities for protection against droplet impact corrosion as function of inhibitor concentrations can be established. REFERENCES 1. API RECOMMENDED PRACTICE FOR DESIGN AND INSTALLATIONOF OFFSHORE PRODUCTION PLATFORM PIPING SYSTEMS,API, Washington D.C. 2. J. C. SMART, Corrosion NACE, paper 468 (1991). 3. Proceedings 5th, 6th, 7th Int. Conf. Erosion by Liquid and Solid Impact, Cavendish Laboratories, Cambridge (1979, 1983, 1987). 4. M. B. LESSER and J. E. FIELD, Proc. 6th. Int. Conf. Erosion by Liquid and Solid Impact, Cavendish Laboratories, Cambridge, paper 17 (1983). 5. J. E. FIELD, M. B. LESSERand J. P. DEAR, Proc. R. Soc. Lond. A 401,225-249 (1985). 6. P. H. PIDSLEY, Proc. 6th Conf. Erosion by Liquid and Solid Impact, Cavendish Laboratories, Cambridge, paper 18 (1983). 7. G. P. THOMASand J. H. BRUNTON,Proc. R. Soc. Lond. A 314, 549-565 (1970). 8. C. DE WAARD, U. LOTZ and D. MILLIAMS, Corrosion NACE, paper No. 577 (1991). 9. N. L. HANCOXand J. H. BRUNTON,Phil. Trans. R. Soc. Lond. A 260, 121-139 (1966).
0010-938X/93 $6.0tl + 0.00 © 1993 Pergamon Press Ltd
INHIBITOR PERFORMANCE U N D E R LIQUID DROPLET IMPINGEMENT CONDITIONS IN CO2-CONTAINING ENVIRONMENT J. H. GERRETSEN and A . VISSER Shell Research Arnhem, Billiton Research BV, P.O. Box 40, 6800 AA, Arnhem, The Netherlands Abstract--Oil and gas production and transport facilities are often constructed from relatively cheap carbon steel and, usually, inhibitors are then applied to mitigate corrosion, CO~ being a particular threat. A proper inhibitor formulation should reduce the corrosion rate by a factor of at least 10, under actual operating conditions. Severe corrosion may occur when film forming inhibitors are removed from the steel by high liquid shear forces. In gas production, impingement of liquid drolets carried in wet, CO2-containing gas may cause such erosion-corrosion. Current knowledge about the mechanism of flow induced removal of such protective inhibitors films is limited. The performances of commercial inhibitors under impact conditions are generally unknown and appropriate test methods for establishing critical conditions are lacking. The API RP-14E formula is often used for estimating critical erosion-corrosion velocities in uninhibited corrosive systems, but this formula is not applicable to inhibited systems. This paper focuses at correlating droplet impact velocity with performances of inhibitor films. Impact is simulated by collision of rotating steel coupons, mounted on a disc, against a liquid jet. To avoid drying of the samples between two collisions, the samples are also continuously wetted during rotation. A marked difference in performance between two commercial inhibitors was found. One inhibitor gave protection outside, but severe corrosion within the impact zone. The other inhibitor gave exactly the opposite behaviour. The equipment was shown to be capable of demonstrating differences in inhibitor performance as function of type and concentration. Critical impact velocities have been obtained for different inhibitors. The performance of the inhibitors tested were shown to be temperature and concentration dependent. The potential of an inhibitor to regain its protective properties, after failure above a critical velocity, could also be studied. This appeared to be a function of temperature for the two inhibitors tested. In principal, the equipment may be used for establishing critical velocities for the onset of erosioncorrosion under liquid droplet impact conditions for corrosion product films.
INTRODUCTION THE LIMITING p r o d u c t i o n v e l o c i t y b e y o n d w h i c h e r o s i o n - c o r r o s i o n in h i g h g a s p r o d u c t i o n o c c u r s , is u s u a l l y c a l c u l a t e d w i t h t h e A P I R P - 1 4 E f o r m u l a : 1 C Vlimiting erosio n -- ~ / ~ ,
(1)
where C may vary from 100 to 400, depending on the policy of the oil companies. The API formula is known to be inaccurate in many cases, usually underestimating t h e l i m i t i n g v e l o c i t y . T h i s is n o t s u r p r i s i n g as t h e f o r m u l a w a s o r i g i n a l l y d e r i v e d for predicting limiting velocities in the absence of inhibitor films or corrosion product scales. Because of the lack of another predictive tool, the formula has been used for Manuscript received 25 September 1992. 1299
1300
J.H. GERRETSENand A. VISSER
other operating conditions. The validity of the formula has recently been discussed by Smart, 2 who suggested that the physical relevance of the formula is given by a change in flow pattern at the critical velocity, namely from stratified multiphase flow to annular mist flow. In the latter case, damage may occur when the annular mist flow is disturbed and liquid droplets impinge on the surface. Flow disturbances can be caused by many geometries, such as bends, elbows, fittings, but also rough surface films may introduce significant turbulence. Current knowledge about inhibitor films and corrosion product performance under liquid droplet impact conditions is limited. When flow is completely smooth, the liquid shear stress is not likely to exceed the critical strain for failure of surface films and therefore is not high enough to cause damage. However, when flow is severely disturbed, damage to films may occur either by increased shear stress or by increased mass transport. Less severe flow conditions may readily remove films containing defects, prevent corrosion product films from forming, or inhibitor films from adhering. The latter is clear in the case of the application of oil soluble inhibitors and the occurrence of water drop-out. In the case of liquid droplet impact, the mechanism of which is briefly discussed in the next section, local shear stress can be extremely high. In addition to damage done by high local shear, damage may also occur by a fatigue mechanism in the case of repetitive impact. To combat erosion-corrosion in high production facilities, often constructed from carbon steel, inhibitors are applied. Today, inhibitor suppliers design the chemical composition of their inhibitors in such a way that they give a suitable protection against general corrosion under anticipated operating conditions. Their corrosion protective properties are usually judged in simple tests, like wheel/bottle tests. Rotating cylinder tests and even sometimes flow loop (smooth pipe flow) tests are increasingly being used to better establish their performance. Although an inhibitor is most likely to fail at places where flow disturbances occur, namely at bends, elbows and fittings, their persistence under more severe flow conditions is usually not thoroughly examined. Sometimes candidate inhibitors are also tested in the field. These field tests focus on general corrosion, while localized corrosion is the real threat, however. To judge the performance of candidate inhibitors for gas production operations where high flow velocities may occur they should specifically be screened on their performance at places where extreme conditions exist. For the specific conditions of annular mist flow a test facility has been built to examine inhibitor performance. In the test facility liquid droplet impact is being simulated by repetitive collisions of horizontally rotating coupons, mounted on the circumference of a disc, against a continuous vertical water jet. Conditions like temperature, ratio oil/water, type and concentration of inhibitor and water chemistry can be varied.
Droplet impact The theory of liquid droplet impact has been studied extensively in the past and excellent overviews can be found in literature. 3 Basically, the impact event resulting in damage can be divided in two parts. Initially the droplet is compressed only and there is no outward liquid flow. The contact area between droplet and the wall increases with a velocity greater than the compression waves. 4'5 Consequently the liquid in contact with the solid is com-
Inhibitor performance under liquid droplet impingement
1301
impact velocity
T ../ ['
\
,
', i
!
/
I q~ high pressure
v edge
steel
FIG. 1. Liquid droplet impact.
pressed. W h e n the speed of sideways m o v e m e n t b e c o m e s less than the compression wave velocity, pressure is relieved and o u t w a r d flow over the surface begins (Fig. 1).6 It has been pointed out that the liquid actually does not simply spread out over the surface but micro-jetting occurs, when the compression is released and the compressed liquid can expand. 7 T h e highest pressure occurs at the contact ring of the droplet before pressure is released. T h e pressure at the surface in the ring m a y easily be as high as 3 times the water h a m m e r pressure, whereas the velocity of the jets may be as high as 10 times the actual impact velocity. Especially at the edges of the contact ring, fatigue m a y occur. Outside the ring, d a m a g e is mainly due to high shear stresses. T h e water h a m m e r pressure is given by: Pw = plqv
(2)
where t~ (kg m -3) is the density of liquid, q (m s -1) is the velocity of sound in the liquid, v (m s -1) is the impact velocity. U n d e r operating conditions, the m a x i m u m pressure will not exceed the yield strength of C-steel (350 MPa). H e n c e , if metal erosion occurs it can only be caused by a fatigue mechanism. W h e n the angle, 0, between the edge of the compressed droplet and the metal is only a few degrees, the high pressure at the edge is reached, and jetting occurs (Fig. 1). E X P E R I M E N T A L METHOD Impact of liquid particles, as may occur in annular mist flow, is simulated in an experimental set-up, where horizontally rotating coupons hit a vertical water jet (Fig. 2). The samples are positioned in the vertical plane and mounted at the circumference of a disc. The samples collide against a continuous vertical water jet which has a diameter of 2 mm. In this way a cylindrical impact event is studied rather than single droplet impact. The mechanism leading to damage is considered to be similar, however. Preliminary experiments have shown that the samples dried immediately after an impact event, due to the centrifugal forces. Hence, after erosion by impact, corrosion could only take place for a fraction of the rotation time. Therefore a construction has been made to allow for continuous wetting during rotation. The construction consists of a reservoir mounted on the disk. A continuous water supply is arranged to the reservoir. Through nozzles liquid is spread over the samples during rotation. Drying of the samples by
1302
J . H . GERRETSENand A. VISSER
centrifugal forces is thus avoided. The water film over the surface of the samples will cushion the impact. In practice, a continuous liquid film will also always be present on the metal surface. The wheel rotates in a steel vessel which is continuously purged with either N2 or CO 2 containing less than 2 ppm O 2. The O 2 content is measured continuously, using an Orbisphere Oxygen sensor, mounted in the outlet gas stream. The steel vessel is placed on top of another vessel. The upper one is funnel shaped to avoid liquid hold-up. The lower vessel is used for preconditioning the liquid before starting an experiment and it supplies a liquid buffer (25 l) to avoid fast contamination with Fe ions and simultaneously a rapid pH increase. The tests can be operated with a two phase liquid (water/oil) flow. Kerosene, type Shellsol K, has been used as an oil phase. NaCl (1%) was added to the total amount of liquid. The liquid is recirculated continuously. Inhibitor can be injected directly to the liquid feed to the samples. The water/oil phases are separated in the lower vessel and, by means of two pumps, re-mixed just before entering the upper vessel. The actual water/oil ratio can be adjusted via the power of the two pumps. The water can be measured in a by-pass. The impact velocity can be varied between 9 and 50 m s -1. Corrosion is monitored in situ by means of pH and sensitive He measurements. During the experiments liquid samples are taken to measure the Fe concentration by ICP measurements. The pH measurements are carried out in the water phase, before re-mixing. H 2 monitoring takes place at the outlet side of the gas. The actual gas throughput is controlled by means of a Brooks mass flow controller. Samples were machined from steel 52-3, of which the chemical composition is given in Table 1. The samples were ground to 320 grit and degreased with alcohol and weighed before mounting. To avoid any galvanic action, the samples are isolated from the disc by PVDF washers. In view of a construction for in situ electrochemical measurements, not described in this paper, the contact angle between sample and
motor ] I
gas crtttIet to H2 ,fl 02 monitarin~
liquid feed gas inlet co2 / ~2
I sample
I" to conditioning vessel FIG. 2.
Schematic presentation of the jet and wheel test rig.
Inhibitor performance under liquid droplet impingement
1303
liquid jet is not exactly 90 ° but 88 ° . Consequently, the tests discussed here do not represent the worst case, the trends, however, are indicative. At the end of an experiment both the jet and the continuous water feed are shut off. The samples are then 'spin' dried for a few minutes. After releasing the pressure to ambient the samples are taken out and rinsed with kerosene and alcohol. After visual examination, corrosion products are removed using Clark's solution and weight loss is measured. The samples are then microscopically examined. The tests discussed in this paper are summarized in Table 2. Two different inhibitors were subjected to the tests. Inhibitor A is oil soluble and water dispersible. Inhibitor B is oil soluble and is not carried over to the water phase. The inhibitors were tested at 45 and 70°C. Also, base case experiments without inhibitors have been carried out. These tests were carried out at 27 and 45°C, to avoid any disturbances of formation of the solid corrosion products.
TABLE 1. Elements %
CHEMICALCOMPOSITIONOF STEEL 52-3
C
Si
Mn
P
S
Ni
Cr
Cu
0.19
0.44
1.41
0.013
0.027
0.08
0.17
0.15
Supplier: MCB+ Valkenswaard, The Netherlands.
pH
J
Tlme (h)
I
I
I
I
I
I
I
+
10
20
30
40
50
60
70
80
FI6.3.
Increase of pH during experiment 2.
1304
J.H. GERRETSENand A. VtssEa mg/i or mm/y 70.O0 60.00 50.00 40.O0 30.00
Corrosion/ pH
'
" .........
10.O0
.
0.00 3.50
3.70
3.90
4.10
4.30
4.50
4.70
pH
4.90
Fl~. 4. Calculatedrelation between pH and Fe2+ concentration (Ill) (mr V I), 45oc, 2 bar CO2, 1% NaCI. Corrosion rate predicted by the nomograms (o) and measured corrosion rate (A) (ram y-l).
accumufi~te~ Fe (rag)
[~[~
800
X
o~~
t
700 600 500 400
l
.~
200
100 0 4 0
~.20
~
40
60
80 i
"---0--- I.C~
i
meas ]
100
120 Time
1-12
FIG. 5. Fez~- accumulation, calculated from H2 measurements and Fez+ concentration obtained from ICP measurements (el, experiment 3. P = measured corrosion rate. Th = corrosion rate accordingto nomogram.
Inhibitor performance under liquid droplet impingement
mm/y or mg/I 45.00
1305
|7O C / 4 bar CO2/1% NaClal
35,00 30,00 25,00 20,00 i 15,00
.
.
.
.
.
.
.
.
.
.
.
.
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.
.
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.
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.
10,00 5,00
.
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.
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.
pH rclation~
0.00
I
3.5
at"
3.7
d
I
I
i
3.9
4.1
4.3
4.5
[
oH 4.7
FZG. 6. CalcuLated relation between pH and Fe 2+ concentration (Ill) (mg 1-1), 70°C, 4 bar CO2, l% NaC1. Corrosion rate predicted by the nomogram s (o) and measured corrosion rate (A) (mm y-l).
accumulated F e (rag) 600
[~~
500
Ichangeto 70 C
I
4OO
200100
~
0 ~
I t0
~
I ol 20
o
P
I 30
4
. Th126 . ~
, 40 o
I.C.P. - -
50
I 60 82 meas
'l 70
! 80 Time (h)
FIG. 7. Fe 2÷ accumulation, calculated from H2 measurements and Fe2÷ concentration obtained from ICP measurements (o), experiment 4. P = measured corrosion rate, Th = corrosion rate according to nomogram, without scaling factor.
90
1306
J.H. GERRETSENand A. VISSER TABLE 2.
Exp.
TEST CONDITIONSAND RESULTS
Impact Inhibitor Water velocity Temp. Inhibitor c o n c . PCO2 cut (m s-l) (°C) type (ppm) (bar) (%)
1 2 3
15 30 10-30
27 45 70
A
100
1 2 4
4 5 6
10-30 10-35 10-30
45-70 45 70
A B B
100 100 100
4 4 4
Remarks
100 Little effect on corrosion rate 100 Little effect on corrosion rate 90 Failurebetween 15 and 20 m s-1, regaining protection at lower velocity 90 Failure,no regaining of protection 90 No failure 90 Bad performance
EXPERIMENTAL RESULTS AND DISCUSSION The test conditions and results are summarized in Table 2. First, tests have been carried out without inhibitors and without kerosene. The samples did not show any preferential attack after these tests. Figure 3 shows the increase of p H during an experiment at 45°C. A change in p H can be correlated with a corrosion rate by calculating the relation between Fe 2+ concentration and p H thermodynamically. This relation is given in Fig. 4. The corrosion rate obtained from the measured p H shift and the corrosion rates as predicted by the 'de W a a r d and Milliams' n o m o g r a m , 8 are also shown in Fig. 4. The corrosion rates showed to be somewhat higher than the corrosion rates expected from the n o m o g r a m , thus possibly indicating a small effect of impact velocity on corrosion rate. A good agreement was obtained between the average corrosion rate calculated from weight loss and from the increase in p H , 12 and 13 m m y-X, respectively. Optically the impact region appeared to be somewhat darker. There was no evidence of metal erosion, as was expected from the impact forces and yield strength of the metal. The amount of dissolved iron remained below the calculated solubility limit of FeCO3. Hence erosion of a product film and subsequent enhanced corrosion could not occur under these conditions. The corrosion rate in experiment 1 at 27°C was also a little higher than expected from the nomogram. Although corrosion rates predicted by the n o m o g r a m have usually been regarded as representing worst case, it seems that corrosion rates can b e c o m e higher than predicted under extreme flow conditions. In the experiments in which inhibitors were studied, the samples were allowed to corrode freely for about 1 or 2 h before injection of inhibitor. In test 3, Fig. 5, carried out at 70°C, the initial corrosion rate was approx. 20 m m y - l , which is below the corrosion rate predicted by the n o m o g r a m , using it without a scale factor (Fig. 6). On injection of the inhibitor, the corrosion rate decreased immediately from about 20 to 0.2 m m y-1. The impact at 10 m s -1 collision speed did not apparently prevent the inhibitor from protecting the steel. W h e n the impact velocity was increased to 3 0 m s -1, the inhibitor failed. The average corrosion rate, determined from H2 measurements, increased to 22 m m y-1. Decreasing the impact velocity resulted in a decrease of the corrosion rate. Apparently this inhibitor is capable of regaining its protection, once damaged. In the next stage of the experiment the velocity was increased step-wise. Figure 5 shows failure of the inhibitor above 20 m s -1. The corrosion rate at 15 m s -1 remained low, therefore it appears that indeed a critical erosion velocity can be obtained for this inhibitor; the corrosion rate being constant
MorphoLogy of samples after Liquid droplet impact, experiments
Continuous water feed
I00 ppm inhib. A 45/70 ° C V impact = variable r
]
profi Le
I00 ppm inhib. B 70oC V impact -- variable i
i
profile FIG. 8.
Photomicrographs of samples from experiment 4(a) and 7(b).
1307
Inhibitor performanceunder liquid droplet impingement
1309
under a critical velocity, above it the efficiency dropping to virtually zero. Other experiments had shown that the critical velocity is dependent on inhibitor concentration. At marginal concentration the inhibitor failed at a much lower velocity. A similar experiment was carried out at 45°C (Fig. 7). It was interesting to observe that in this case the inhibitor film did not recover after being damaged at a velocity of 30 m s -1 (between 48 and 64 h) and subsequent decrease of velocity. Moreover, when the inhibitor film lost its protection, the efficiency was not reduced as drastically as at 70°C and some protection remained. From the results of the experiment carried out at 70°C, it was expected that the inhibitor would regain its protection when the temperature was increased to 70°C. Protection was not regained, however. This effect is not well understood. A speculative explanation may be given by the surface roughening which occurs during the period in which the inhibitor film is damaged. Initially, corrosion will roughen the surface, increasing the surface shear forces significantly. However, at a certain roughness, the effect of the impact will begin to reduce, because the droplets will break-up sooner after the impact, consequently reducing pressure build-up and actual liquid velocity. Considering the corrosion rates during failure, it is likely that during the time of failure at 45°C, roughening of the surface has been less than in the case of the 70°C experiment, hence the effect of the impact is less in the latter case at a similar impact velocity. At 45°C, inhibitor B did not fail at all. Even at an impact velocity of 35 m s -1 , there was no sign of failure. The general corrosion rate was reduced from an initial 11 to about 0.25 mm y-1. Apparently, inhibitor B performed extremely well under these impact conditions. However, similar experiments carried out at 70°C clearly showed again that, in general, a performance of an inhibitor under one set of conditions is certainly not a guarantee for similar behaviour under different conditions. In this case the effect of temperature was even more dramatic. The inhibitor which was very resistant against the liquid droplet impact at 45°C, now performed poorly. After injection of the inhibitor at an impact velocity of 10 m s t, the corrosion rate dropped significantly, but remained always higher than 2 mm y-~, which is clearly unacceptable. These results illustrate that for making a selection from several candidate inhibitors for the actual operations, the inhibitors should be screened at a range of conditions, temperature being an important parameter. Critical erosion velocities can be obtained by using the jet and wheel technique for conditions which are not addressed by the API RP-14E formula. The behaviour of inhibitors under droplet impact conditions can be studied. In a similar manner the mechanical stability of corrosion product layers can be established.
Morphology The morphology of the samples after being exposed to the repetitive impact was intriguing. As can be seen from Fig. 8(a) and (b), from experiments 4 and 6 respectively, inhibitor A has failed preferentially in the area where high pressure occurs; the outer region being well protected. The performance of inhibitor B, on the contrary, proved to be better in the impact region, whilst the film completely failed on the outside. It appears that inhibitor A is less susceptible for high shear under the test conditions, where inhibitor B is less susceptible to fatigue. This phenomenon cannot be explained as yet, but it clearly indicates the complexity of the mechanisms involved in the liquid droplet events. According to
1310
J.H. GERRETSENand A. VISSER
Thomas, 7 the highest pressure should have occurred at a distance d = Rv/q, thus with a diameter, R, of the jet being 2 mm and an impact velocity of 20 m s-1 and assuming the velocity of sound in the liquid, cb being equal to those in water (1500 m s-l), d = 0.027 mm. Figure 8 shows that this region is about 2 mm, however. Such a discrepancy has been observed by others as well. 9 In our experiment, cushioning of the impact by the continuous liquid film at the surface may have increased the distance outside of which high liquid shear exists. Distortion of the jet by the aerodynamics in the vessel may also increase the effective distance. Another factor involved is instability of the jet. The distance, d, will also increase when oscillation of the jet occurs. CONCLUSIONS (1) The corrosion rate under liquid droplet impact conditions can be higher than expected from the De Waard and Milliams nomogram. (2) Inhibitor performance under simulated liquid droplet conditions depends strongly on temperature. (3) Recovery of inhibitor performances, once adversely affected, does not always occur instantaneously. Temperature was shown to be an important parameter here as well. (4) The morphology of corrosion under liquid droplet impact depends on the chemical composition of the inhibitor; one inhibitor may be susceptible to high shear, whereas another may be susceptible to damage by pressure. (5) The jet and wheel test facility allows for inhibitor screening tests for gas production facilities. Critical velocities for protection against droplet impact corrosion as function of inhibitor concentrations can be established. REFERENCES 1. API RECOMMENDED PRACTICE FOR DESIGN AND INSTALLATIONOF OFFSHORE PRODUCTION PLATFORM PIPING SYSTEMS,API, Washington D.C. 2. J. C. SMART, Corrosion NACE, paper 468 (1991). 3. Proceedings 5th, 6th, 7th Int. Conf. Erosion by Liquid and Solid Impact, Cavendish Laboratories, Cambridge (1979, 1983, 1987). 4. M. B. LESSER and J. E. FIELD, Proc. 6th. Int. Conf. Erosion by Liquid and Solid Impact, Cavendish Laboratories, Cambridge, paper 17 (1983). 5. J. E. FIELD, M. B. LESSERand J. P. DEAR, Proc. R. Soc. Lond. A 401,225-249 (1985). 6. P. H. PIDSLEY, Proc. 6th Conf. Erosion by Liquid and Solid Impact, Cavendish Laboratories, Cambridge, paper 18 (1983). 7. G. P. THOMASand J. H. BRUNTON,Proc. R. Soc. Lond. A 314, 549-565 (1970). 8. C. DE WAARD, U. LOTZ and D. MILLIAMS, Corrosion NACE, paper No. 577 (1991). 9. N. L. HANCOXand J. H. BRUNTON,Phil. Trans. R. Soc. Lond. A 260, 121-139 (1966).