P R E P R I N T – ICPWS XV Berlin, September 8–11, 2008

A Fundamental Study of Flow-Accelerated Corrosion in Feedwater Systems

D.H. Listera, L. Liua, A. Feichta, M. Khatibia, W. Cooka, K. Fujiwarab, E. Kadoic, T. Ohirac, H. Takiguchic and S. Uchidad

aUniversity of New Brunswick, Fredericton, New Brunswick, Canada

bCentral Research Institute of the Electric Power Industry, Yokosuka, Japan cJapan Atomic Power Co., Tokyo, Japan

dJapan Atomic Energy Agency, Tokai, Japan Email: lihui@unb.ca

A once-through water loop operating under typical power reactor feedwater conditions is being used to determine the effects of water chemistry and flow on flow-accelerated corrosion (FAC). Electrochemical corrosion potential (ECP) and corrosion rate are measured with on-line probes and mechanisms are indicated by detailed surface analyses. The program so far has investigated FAC at 140°C in both neutral and ammoniated water over a range of flow rates. In neutral wa-ter, FAC was correlated well with fluid shear stress at the surface and was stifled with oxygen concentrations of about 40µg/kg (ppb). Chromium in the steel reduced FAC significantly. Cor-roded surfaces developed thin magnetite films and some overlaid typical “scalloped” textures. Scallops were influenced by the microscopic oxide structures that developed from underlying metal grains; in particular, pearlite grains formed lamellar oxides that predominated on scallop crests. In ammoniated water at a room temperature pH of 9.15-9.35, FAC was extremely sensi-tive to traces of oxygen and apparently proceeded by a “front” of protective oxide, based on haematite, moving downstream as the less protective magnetite was progressively oxidised. The subsequent reduction of a more oxidised, stifling film back to magnetite and the resumption of FAC when oxygen levels were reduced also proceeded from the upstream ends of probes.

Introduction

Flow-accelerated corrosion (FAC) of carbon steel is a common problem in many types of steam-raising plant. The catastrophic failure of a suction line to the main feedwater pump at the Surry-2 PWR in 1986 and of a feedwater line at the Mi-hama-3 PWR in 2004 resulted in fatalities; similar accidents at fossil plants have been reported [1,2]. Not surprisingly, further insights into the mecha-nisms of FAC are being sought in the search for mitigating techniques. To that end, a collaborative research program between Japan and Canada is investigating the details of how dissolved oxygen reacts with oxide films on carbon steel undergoing FAC under PWR feedwater conditions. In a labora-tory loop at the University of New Brunswick (UNB) in Canada, on-line probes indicate electro-chemical corrosion potentials and corrosion rates while analyses of exposed specimens are carried out at UNB and at the Central Research Institute of the Electric Power Industry (CRIEPI) in Japan. The results are reviewed by the whole team, which includes representatives from the Japan Atomic

Power Company (JAPCo) and the Japan Atomic Energy Agency (JAEA). The mechanisms leading to a threshold concentration of oxygen, above which FAC is inhibited or “stifled”, are being sought in order to specify a practical value that may be used in operating power plants.

Experiments

The experimental water loop has been described before [3]. It is made mostly of Hastelloy-C and stainless steel and can operate at temperatures up to 310°C and pressures up to 11MPa. Although de-signed as a once-through system, its coolant is recirculated from the pressure-reducing valve to the positive-displacement pump via a cooler, ion-exchange columns and a controlled-atmosphere tank where the chemistry is adjusted with ammonia, hydrazine, etc. The pump can deliver up to 3.5 L/min of coolant to the test section, which is fitted with by-pass piping to enable probes to be inserted or removed without shutting down the whole loop. Oxygen is added as required via a pump injecting solutions in water to the low-pressure coolant return line before the tank, where an on-line Orbisphere

Figure 1. Schematic diagram of FAC loop

measures its concentration continuously. A sche-matic diagram of the loop is presented in Figure 1.

Rates of FAC are monitored on-line with tubu-lar probes made of the carbon steel of interest. The tubes are typically 90 mm long with bores of 1.6 mm, 2.4 mm or 3.2 mm and they have the middle length of about 12 mm turned down to a wall thick ness that gives an electrical resistance that can be measured accurately; changes in resistance of 10 µΩ (±4%) can be detected reliably (see Figure 2, which shows a typical resistance probe with its restraint to avoid disassembly under pressure). The rate of increase of resistance as measured with a multimeter then gives a measure of the FAC rate as the wall thins. Electrochemical corrosion potential (ECP) of a resistance probe relative to a high-temperature Ag/AgCl reference electrode is also measured. Two or three probes are typically in-stalled in series in the test section and tubes of simi-lar sizes are installed downstream to be removed as required for surface analysis.

Stainless steel probes designed solely to meas-ure ECP were installed in TS 1. It was noted that during the commissioning of the loop, when the coolant initially was saturated in air, ECP finally levelled off at about – 215 mV (SHE), but only

about three days after measured oxygen had fallen to zero.

Three runs are described here. Run 1 used a carbon steel (CS) with 0.019 wt % Cr, the other two

Figure 2. Diagram of resistance probe

runs used a carbon steel specially made with 0.001 wt% Cr. All runs were at 140 °C, the temperature of the feedwater line that ruptured at Mihama 3 and close to that known to produce the maximum rate of FAC. Runs 1 and 2 were at neutral chemistry

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conditions while Run 3 had ammonia added to give a pH25°C between 9.15 and 9.35.

In Run 1, probes with three internal diameters of 2.4, 1.6 and 3.2 mm (in order downstream) were installed in each of Test Sections (TSs) 1, 2 and 3. Those in TS 1 were resistance probes to measure FAC rate on-line, those in TS 2 were for surface analysis at the end of the run and those in TS 3 were for surface analysis after each change in loop parameter such as flow rate (the TS 3 probes were replaced at each change). The effects of flow rate were studied.

In Run 2, probes of the lower-Cr steel with in-ternal diameters of 2.4 and 1.6 mm were installed in TSs 1, 2 and 3. The effects of concentration of added oxygen on FAC were studied and indications of flow-rate effects were obtained to compare with the results from the higher-Cr steel in Run 1.

In Run 3, probes of the lower-Cr steel with in-ternal diameters of 2.4 and 1.6 mm were again installed in TSs 1, 2 and 3. At an average pH25°C of 9.2 with ammonia, the effects of concentration of added oxygen on FAC rate were studied.

Results

Since the steels used for the probes were sepa-rately calibrated for electrical resistivity over a

range of temperatures using a temperature-controlled furnace, the measurements of electrical resistance were readily converted to wall thickness and the inner radius calculated. The slopes of plots of radius versus time indicated FAC rate. Similar plots were obtained from the other runs.

The radius plots from Run 1 shown in Figure 3 indicate that the response to changing flow rate was rapid. In some experiments under similar condi-tions, where FAC rates are high, the plots tend to become less steep with time as the probe bores increase and the coolant linear velocity decreases. This effect is not evident in Figure 3. The effect of coolant temperature is evident, however, since shifts in the data occurred with inadvertent heating or cooling. The slopes of the plots – and therefore the indicated FAC rates – were unaffected. The effect of added oxygen is also evident in Figure 3, which shows the stifling of FAC for about two days when the ion-exchange column was replaced with a fresh, air-saturated one at Day 14.

Measurements of ECP were less clearly defined. Ingress of oxygen into the loop, particularly during runs under ammoniated conditions, controlled the ECP. The ammonia affected many of the polymer seals in the loop; they degraded and became perme-able to oxygen and had to be replaced with more resistant materials. Under neutral conditions, oxy-

Figure 3. Plots of probe radius versus time for Run 1.

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Figure 4. SEM pictures of probe inner surfaces after Run 1.

gen ingress was much easier to control. Seals were effective and the high FAC rates, driven by high iron solubility, appeared to act as a sink for oxygen. On the CS resistance probes, even when the oxygen may have been too dilute to affect the FAC, it had a measurable effect on the ECP. In general, the ECP plots for the CS probes varied in concert and fell with increasing FAC rate – as expected. Values between -610 and -950 mV (SHE) were measured in Run 1.

The examination of the probe surfaces after Run 1 revealed a distinct effect of coolant flow on the morphology (see Figure 4). Oxide “stringers” were aligned axially and scallops were evident – the latter in particular on the more corroded surface of the 1.6 mm probe. The oxides were identified as magnetite with laser-Raman microscopy.

A more detailed scrutiny of the oxides indicated that the FAC mechanisms had promoted unusual oxide formations. As indicated in Figure 5, “coral-like” growths were associated with the stringers and with the scallop crests.

Metallography of a probe cross-section and its oxide showed that the growths occurred over pear-lite grains in the metal, so it is assumed that the

cementite and ferrite lamellae in the pearlite form oxides of different composition. As FAC proceeds by the accepted mechanism of the protective mag-netite film’s being thinned by dissolution in turbu-lent coolant undersaturated in dissolved iron, so that it reaches a steady-state thickness at a constant FAC

rate, the oxide over the ferrite lamellae will dissolve at about the same rate as that over the surrounding ferrite metal grains, leaving the oxide over the ce-mentite to form the “coral”. Presumably, as these oxide features develop, the flow patterns of the coolant at the surface are modified so that the local turbulence perpetuates the scallops. The magnetite

Figure 5. Oxide on scallop crest in Run 1

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in the scallop troughs was very fine-grained (see Figure 6). On average, the base oxide was 0.5 – 1.0 µm thick, which is customary for FAC situations.

Figure 6. Oxide in scallop trough in Run 1

Energy-dispersive X-ray analysis (EDX) indi-

cated that trace elements from the underlying metal were concentrated in the oxides on the probes ex-posed in Run 1. In particular, Cr was concentrated to a ratio with Fe of 3.35×10-3 in the crest oxide and to 2.01×10-3 in the valley oxide; by contrast, the underlying metal had a Cr:Fe ratio of 1.92×10-4 (EDX of such thin oxides has an inherent uncer-tainty because of interference from the underlying metal; however, the interference is expected to produce low values if anything, so that the concen-tration effect should be real). Note that most of the Cr in the oxides probably came from the underlying metal, since the main loop surfaces that might re-lease Cr to the coolant (i.e., the stainless steel and Hastelloy components) should also release Ni, yet no (or only small amounts of) Ni were found on the probes. Chromium is known to impart protec-tive properties to carbon steel in FAC environments [4], presumably by making the oxide more compact and less soluble.

In Run 2, the two probes (2.4 and 1.6 mm – in order downstream) produced radius plots similar to those of the higher Cr steel in Run 1, except that the FAC rate was about 2.5 times higher. This com-paratively rapid increase in radius lowered the lin-ear velocity of the coolant in the probes signifi-cantly during the run, so that the FAC rate declined with time. An average FAC rate for a period at an average velocity is therefore quoted here. This allowed velocity/diameter effects to be examined in more detail for this run than would otherwise be possible with only two nominal probe sizes and a constant volumetric flow rate. Part-way through the run, oxygen was injected in increments to find the concentration at which the FAC was stifled. At about 40 ppb, the upstream probe stifled to be fol-lowed by the other, downstream, probe about four hours later. This apparent movement of an oxidized front moving downstream was seen in Run 3 under

ammoniated conditions described later. The onset of stifling and the resulting removal of an oxygen sink caused the oxygen concentration to jump to 120 ppb. It should be noted that oxygen had to be completely removed – to zero concentration – for a day or so before FAC could be resumed.

This hysteresis in the effect of oxygen no doubt reflects the interaction with the oxide film on the probe surface. It is postulated that for stifling of FAC to occur, the oxygen has to penetrate through the pores of the magnetite to the metal-oxide inter-face, where it must achieve some minimum concen-tration in order to react with the Fe2+ ions as they are produced and form a more protective oxide based on Fe2O3 (maghaemite or haematite) rather than Fe3O4 (magnetite). In diffusing through the magnetite, it will also react with the Fe2+ in the oxide surface and may never reach the metal if the driving force is too low. Conversely, when a sur-face is stifled with a passive metal-oxide interface protected with Fe2O3 and the coolant becomes de-pleted in oxygen, the oxide has to be reduced to some extent before the production of Fe2+ ions can recommence at the metal. Laser-Raman studies of the surface of a probe removed during the period of oxygen injection but before stifling (when the con-centration was about 20 ppb) in fact indicated the presence of haematite.

During Run 2, the ECP varied roughly as the oxygen concentration, with the 2.4 mm probe showing the largest variation (between about -1 V vs. SHE at zero oxygen to about -170 mV at the peak oxygen level); a somewhat higher potential of the 1.6 mm probe was attributed to its different location in the loop relative to the reference elec-trode.

Curiously, although the FAC rates during the non-oxidising phases were higher than in Run 1, scallops were not seen on the probes. Moreover, the oxides showed little tendency to align with the flow direction, though SEM examination at high-magnification showed coral-like oxides similar to those associated with underlying pearlite grains that were prevalent in the scallop crests in Run 1 (see Figure 7).

This lack of tendency to produce scallops can-not be attributed to a different metal grain structure of the low-Cr steel, since metallography indicated that it was similar to that of the higher-Cr steel. In particular, the pearlite grains were distributed simi-larly, although those of the lower-Cr steel in Run 2 were slightly more elongated. In any case, as de-scribed later, localised severe scalloping was found

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on this material exposed under ammoniated condi-tions, when general FAC rates were lower.

Figure 7. Oxide feature on low-Cr steel in Run 2

In Run 3, under ammoniated chemistry at pH25°C

of 9.2 with no oxygen, the FAC rates of the two resistance probes (2.4 and 1.6 mm diameter – in order downstream) of the lower-Cr steel were about half of those under neutral conditions.

A flow correlation that has been applied before [3] to Runs 1 and 2 has been applied to Run 3 (see Figure 8). Note that only two probes were installed in this run, so only two flow conditions were ob-tained. However, since the correlation is expected to pass through (or close to) the origin, the zero

point has been included. By contrast, Run 1 had 32 flow conditions and Run 2 had four.

The correlation is based on the premise that the FAC of carbon steel is mass-transfer controlled, which in turn is based on the postulate that there are two processes in series – the corrosion at the metal-oxide interface creating the oxide film there by half the iron that enters solution precipitating as magnet-ite (the rest diffuses through the film and is lost to the bulk coolant), followed by the dissolution of the film at the oxide-coolant interface by coolant un-dersaturated in dissolved iron.

This leads to the equation for FAC rate R [5]:

( )d

d

khChk

R2

2+

Δ= (1)

where ΔC is the undersaturation in iron, kd is the dissolution rate constant and h is the mass transfer coefficient. If mass transfer controls, h is small compared with kd and Equation (1) reverts to:

ChR Δ= (2) so that R varies as h for a given chemistry condition.

For a given flow condition, the differences in R for different materials (e.g., the lower- and higher-Cr steels in these experiments) are then caused by the different equilibrium concentrations of dis-solved iron (i.e., solubilities of the oxide films) in

Figure 8. Flow correlation for Runs 1, 2 and 3

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the ΔC term. The effect of Cr on creating more insoluble oxides was commented on earlier.

Correlations for mass transfer usually relate the dimensionless quantities Sherwood Number and Reynolds Number (Re) to link the mass transfer to the flow characteristics. For constant physical properties of the system, this leads to the expression for FAC rate:

pReARd = (3) where d is pipe diameter and A and p are constants. For many applications, the Re exponent p falls between 0.6 and 0.9 [6,7]. However, the data for Runs 1 and 2 gave values of 1.2 and 1.3 with corre-lation coefficients of 0.83 and 0.98, respectively. These exponents seem rather high. The three points (including the zero) for Run 3 gave a very high exponent of 4.8.

An alternative approach leading to Figure 8 considered the Reynolds analogy of transport pa-rameters in turbulent flow. The Stanton number for mass transfer (Sh/Re/Sc, where Sh is the Sherwood number and Sc is the Schmidt number) is equated to the friction factor (τ/(ρ.u2), where τ is the fluid shear stress and ρ is the fluid density). If the (constant) property terms are neglected and it is again as-sumed that R is proportional to h, Equation (4) is obtained:

τBRu = (4) where B is a constant.

The least-squares regressions for the data in Figure 8 indicate that for Runs 1 and 2 the relation-ships are virtually linear, with exponents of shear stress of 1.06 and 0.97, respectively (the corre-sponding correlation coefficients are 0.98 and 1.00). Equation (4) apparently holds for FAC under these conditions of neutral chemistry, when magnetite solubility is high and corrosion is rapid. For Run 3, the exponent of 1.52 is rather high to attribute the correlation to Equation (4). This might suggest that the assumption of mass-transfer control is doubtful for high-pH chemistry, when magnetite solubility is comparatively low (about 14 ppb at 140°C under ammoniated conditions at pH25°C of 9.2, in compari-son with about 119 ppb under neutral conditions [8]). A similar suggestion might arise from the fitting of the Run 3 data to the Reynolds number correlation, which gave an exponent of 4.8 (men-tioned earlier). However, with only two data points, such a conclusion remains speculative.

The sensitivity of the loop oxygen control under ammoniated chemistry was mentioned earlier. This was also reflected in the probe behaviour when oxygen was injected in Run 3. That FAC is stifled under ammoniated conditions at low concentrations

of oxygen is well-established [9]; stifling concen-trations here were within 1-2 ppb for both the 1.6 and the 2.4 mm probe, determined during additions to 10 ppb over four weeks. The upstream, 2.4 mm, probe stifled about two days before the downstream one. After the resistance probes had stifled, down-stream surface analysis probes were removed from the loop and examined. It was significant that the more upstream probe exhibited a “front” of haema-tite-rich oxide part-way along its length. Upstream of the front the probe surface had a thick, red oxide while downstream the oxide was the characteristic shiny black of FAC-induced magnetite. Laser-Raman microscopy and SEM confirmed that the upstream, red oxide was haematite, comprising a thick layer of plates and needles several µm across. Presumably, these were Fe2O3 crystals, and they were on top of what appeared from their octahedral shape to be Fe3O4 crystals, again up to several µm in size. These formations are in Zone 1 in Figure 9. Downstream, in the FAC zone (Zone 4 in Figure 9), the by-now familiar coral-like formations of oxide associated with pearlite grains in the metal ap-peared with the surrounding thin layer of magnetite.

Figure 9. Inside surface of 2.4 mm probe after

exposure to oxidizing chemistry in Run 3 These observations of surface features are con-

sistent with an oxidation front that moves down-stream as the corroding surface becomes progres-sively protected by the conversion of its magnetite film to haematite or maghaemite by the dissolved oxygen. The thickness of the oxide on the stifled portion of the probe surface suggests that the plates and needles precipitated on top of the magnetite, which itself oxidised rather slowly. Also, the front must have allowed magnetite crystals to form as haematite precipitated on top. Presumably, the oxygen at the surface varied from close to measured bulk concentration in Zone 1 to below the threshold level for stifling in Zone 4. Further downstream, the concentration probably approached zero as oxygen interacted with the FAC process (the “sink”).

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Figure 10. Upstream (2.4 mm) probe after exposure to oxygenated then low-oxygen coolant in Run 3

(left-to-right: upstream end, middle, downstream end)

Figure 11. Downstream (1.6 mm) probe after exposure to oxygenated then low-oxygen coolant in Run 3

(left-to-right: upstream end, middle, downstream end)

This observation of an oxidising front moving downstream opens the possibility of adding oxygen to the upstream end of a coolant system undergoing FAC in amounts sufficient to passivate the surface progressively and to be consumed to the threshold stifling level at the downstream end. The passivated state could then be maintained with oxygen concen-trations close to the threshold.

In order to find out how long the stifled probes would take to resume FAC, Run 3 was continued under deoxygenated conditions (zero-0.5 ppb, with an occasional spike to about 20 ppb when a fresh ion-exchange column was valved in). After run-ning in this mode for about five months, with no resumption of FAC, the two resistance probes were removed from the loop, sectioned and analysed with SEM and laser-Raman microscopy in the same way as surface-analysis probes were examined.

Figures 10 and 11 present low-magnification SEM pictures of the inside surfaces. Both were matt-black, although the upstream (2.4 mm) probe had a tinge of brown/red and the 1.6 mm probe seemed to have a thicker film. Surprisingly, both probes displayed severe corrosion and scalloping at the upstream end, visible to the naked eye, in which the metal glinted through the oxide. A rough com-parison indicates that the inlet of the 1.6 mm probe corroded to a diameter similar to that of the inlet of the corroded 2.4 mm probe. Laser-Raman analysis along the lengths of both probes indicated mainly magnetite with an admixture of haematite that in-

creased towards the downstream end. The oxide film on the downstream 1.6 mm probe gave a stronger signal than that on the other probe, sug-gesting it was in fact thicker, although there seemed to be more haematite on the upstream probe. The reason why neither probe had indicated the resump-tion of FAC was clear; the severely corroded and scalloped length at the upstream end (about 4 cm for both probes) had not extended far enough into the machined and monitored length in the middle of the tube to affect the electrical resistance.

These observations are significant. They indi-cate that FAC resumed on both probes but only to a certain distance downstream from the entrance. Presumably, the fluid-flow turbulence due to the entrance effect augmented the effect of the more- reducing chemistry under low-oxygen conditions to modify the passivating film based on haematite and make it less protective. This might indicate that the reduction of the oxide is mass-transfer controlled; however, both the 1.6 mm and the 2.4 mm probe had the section of resumed FAC extending about the same distance downstream, so the reduction “front” would have progressed at the same rate downstream, even though the coolant velocity and Reynolds number in the 1.6 mm probe were higher than in the 2.4 mm probe. In any case, the mecha-nism cannot be likened to the one that generated the oxidation front that was described earlier for the oxygen addition experiment, where there was a reactant (dissolved oxygen) that was transported

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and depleted downstream by reaction with surfaces (separate calibration studies measured gradients in dissolved oxygen around the loop; see [3], for ex-ample). Rather, it would have been the slow con-version of a protective, ferric-based oxide at the metal-oxide interface, presumably controlled by the very low solubility of that oxide. More experi-ments are required to identify the mechanism pre-cisely.

Summary and conclusions

Experiments on the flow-accelerated corrosion (FAC) of carbon steel have been carried out using on-line electrical resistance probes installed in se-ries in a laboratory loop operating under simulated PWR feedwater conditions at 140 °C. Results have indicated the effects of chromium content of the steel, of coolant flow, of ammonia additions at pH25°C of 9.2 and of oxygen additions to the coolant, on FAC rate.

At neutral chemistry, steel containing 0.001% Cr corroded at about 2.5 times the rate of steel containing 0.019% Cr. FAC was apparently con-trolled by mass transfer, although standard mass-transfer correlations linking the rate with Reynolds number did not describe the results as well as one derived from the Reynolds analogy linking the rate with fluid shear stress. Scalloping, influenced by curious oxide formations that were associated with the grain structure of the metal, was seen only on the higher-Cr steel. An oxygen concentration of 40 ppb stifled FAC of the lower-Cr steel.

Adding ammonia approximately halved the FAC rate of the lower-Cr steel. Mass-transfer-based relations did not describe the FAC rates un-der these conditions very well, although the data were too few for the conclusion to be definitive. The high-pH chemistry made the loop very sensi-tive to dissolved oxygen. A concentration of 1 – 2 ppb was sufficient to stifle FAC, apparently via a front of oxide based on haematite (rather than mag-netite, which is commonly found on steel undergo-ing FAC) that progressed downstream. The subse-quent resumption of FAC when dissolved oxygen was removed occurred only at the inlet sections of probes, where pronounced corrosion and scalloping were seen.

Acknowledgement

The authors would like to thank their respective organisations for supporting this research program. In addition, they are grateful to the people – too

numerous to mention individually – who have con-tributed to the laboratory work and to the ongoing technical discussions.

References

[1] Czajkowski, C., “Metallurgical evaluation of an 18-inch feedwater line failure at the Surry Unit 2 power station”, NUREG/CR-4868. Brookhaven Nat. Lab. (1987). [2] NISA, “Final report on Mihama-3 secondary system piping failure”, Nuclear and Industrial Safety Agency, Tokyo, Japan (2005). [3] Lister,D.H.,Feicht,A.,Cook,W.,Khatibi,M.,Liu,L., Ohira,T., Kadoi,E., Takiguchi,H. Fujiwara,K, and Uchida,S. “Effects of dissolved oxygen on flow-accelerated corrosion in feedwater

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[4] Bouchacourt, M. “Identification of key variables: EdF studies”EPRI Workshop on Erosion-Corrosion of Carbon Steel Piping, Washington, DC, USA (1987 April). [5] Berge, P., Ducreux, J. and Saint-Paul,P., “Effects of Chemistry on Erosion Corrosion of Steels in Water and Wet Steam”, Proc. 2nd. BNES Internl. Conf. on Water Chem. of Nucl. Reactor Systems, Bournemouth, UK. (1980). [6] Berger, F.P. and Hau, K-F. F-L., “Mass Transfer in Turbulent Pipe Flow Measured by the Electrochemical Method”, Internl. J. Heat and Mass Trans., 20, 1185 (1977). [7] Sydberger, T. and Lotz, U., “Relation Between Mass Transfer and Corrosion in a Turbulent Pipe Flow”, J. Electrochem. Soc. 129 (2), 276 (1982). [8] Tremaine, P.R. and LeBlanc, J.C., “The

Solubility of Magnetite and the Hydrolysis and Oxidation of Fe2+ in Water to 300oC”, J. Solution Chem., Vol. 9 No. 6 (1980).

[9] Woolsey, I.S., Bignold, G.J., De Whalley, C.H. and Garbett, K., “The Influence of Oxygen and Hydrazine on the Erosion-Corrosion Behaviour and Electrochemical Potentials of Carbon Steel under Boiler Feedwater Conditions”, Proc. 4th BNES Internl. Conf. on Water Chem. of Nucl. Reactor Systems, Bournemouth, UK. (1986).

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