Comprehensive nuclear materials 5 06 assisted cracking of carbon and low alloy steels

Comprehensive nuclear materials 5 06 corrosion and environmentally assisted cracking of carbon and low alloy steels Comprehensive nuclear materials 5 06 corrosion and environmentally assisted cracking of carbon and low alloy steels Comprehensive nuclear materials 5 06 corrosion and environmentally assisted cracking of carbon and low alloy steels Comprehensive nuclear materials 5 06 corrosion and environmentally assisted cracking of carbon and low alloy steels Comprehensive nuclear materials 5 06 corrosion and environmentally assisted cracking of carbon and low alloy steels Comprehensive nuclear materials 5 06 corrosion and environmentally assisted cracking of carbon and low alloy steels

5.06 Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels H.-P Seifert Paul Scherrer Institut, Villigen PSI, Switzerland J Hickling Independent Technical Consultant, Prastio-Avdimou, Cyprus D Lister University of New Brunswick, Fredericton, NB, Canada ß 2012 Elsevier Ltd All rights reserved 5.06.1 Introduction 107 5.06.2 5.06.2.1 5.06.2.2 5.06.2.2.1 5.06.2.2.2 5.06.2.2.3 5.06.3 5.06.3.1 5.06.3.2 5.06.3.2.1 5.06.3.2.2 5.06.3.2.3 5.06.3.2.4 5.06.3.2.5 5.06.4 5.06.4.1 5.06.4.2 References Uniform and Flow-Accelerated Corrosion Uniform Corrosion Flow-Accelerated Corrosion Controlling factors Mechanisms and models Service experience and mitigating actions Localized Corrosion and Environmentally Assisted Cracking Pitting Environmentally Assisted Cracking Basic types of EAC and major factors of influence Corrosion fatigue and strain-induced corrosion cracking Stress corrosion cracking EAC mechanisms and models Service experience and mitigation actions Conclusions Uniform and Flow-Accelerated Corrosion Localized Corrosion and Environmentally Assisted Cracking 109 109 111 111 114 118 120 120 122 122 123 128 132 136 139 139 139 140 Abbreviations AC AGR ANL ASME ASME BPV ASME III ASME XI ASTM BWR BWRVIP BWRVIP-60 Content of Cr, Mo and Cu in alloy in EPRI ‘CHECWORKS’ FAC-Code Advanced gas-cooled reactor Argonne National Laboratory, USA American Society of Mechanical Engineers ASME Boiler and Pressure Vessel Code Section III of ASME BPV Code Section XI of ASME BPV Code American Society of Testing and Materials Standards Boiling water reactor Boiling Water Reactor Vessel and Internals Program Basis document for SCC disposition lines for low-alloy steels CANDUW CF CRDM CS DCPD DH DL DO DSA EAC EC ECP ECPcrit CANada Deuterium Uranium, PHWR developed by Atomic Energy of Canada Ltd Corrosion fatigue Control rod drive mechanism (housing) Carbon steel (Reversed) direct current potential drop crack length measurement method Dissolved hydrogen (concentration) Disposition line Dissolved oxygen (concentration) Dynamic strain ageing Environmentally assisted cracking Erosion corrosion Electrochemical corrosion potential Critical cracking potential (e.g., for SICC) 105 106 Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels EPRI Electric Power Research Institute, USA F & A model EAC model for CS & LAS developed by P Ford and P Andresen (GE GR) FAC Flow-accelerated corrosion FRAD Film rupture anodic dissolution EAC mechanism HAEAC Hydrogen-assisted EAC mechanism HAZ Heat-affected zone of weldment HCF High-cycle fatigue HT High temperature HWC Hydrogen water chemistry JSME Japanese Society of Mechanical Engineers LAS Low-alloy steel LCF Low-cycle fatigue LEFM Linear elastic fracture mechanics LWR Light water reactor MT Mass transfer in EPRI ‘CHECWORKS’ FAC-Code NDT Nondestructive testing NMCA Noble metal chemical addition NRC Nuclear Regulatory Commission, USA NWC Normal water chemistry PHWR Pressurized heavy water reactor PWHT Postweld heat treatment PWR Pressurized water reactor PWSCC Primary water stress corrosion cracking (in PWRs) QỵT Quench and temper heat treatment RPV Reactor pressure vessel SCC Stress corrosion cracking SEM Scanning electron microscope SHE Standard hydrogen electrode SICC Strain-induced corrosion cracking SS Stainless steel SSR(T) Slow strain rate (test) SSY Small-scale yielding UTS Ultimate tensile strength VGB German Association of Large Power Plant Operators YS Yield strength Ceq Symbols KI KI,i C Cb Concentration of Fe(II) species at the oxide–coolant interface Concentration of Fe(II) species in the bulk coolant CODLL d D da/dN da/dNAir da/dNCF da/dtAir ¼ da/dNAir/ DtR da/dtCF ¼ da/dNCF/ DtR da/dtSCC da/dtSICC dCODLL/dt de/dt de/dtcrit dKI/dt EA Fen G h kc kd Thermodynamic equilibrium concentration of Fe(II) species Crack-opening displacement at load line in precracked fracture mechanics specimen Pipe diameter Diffusivity Crack advance per fatigue cycle Crack advance per fatigue cycle in air Corrosion fatigue crack advance per fatigue cycle in hightemperature water Time-based fatigue crack growth rate in air Time-based corrosion fatigue crack growth rate in hightemperature water SCC crack growth rate SICC crack growth rate Crack-opening displacement rate in slow rising load or displacement test Strain rate (sometimes locally at crack-tip) Critical strain rate (e.g., for SICC) Stress intensity factor rate in slow rising load or displacement test Arrhenius activation energy of thermally activated process Environmental correction factors, ratio of fatigue life in air at room temperature to that in water at service temperature Geometry factor in EPRI ‘CHECWORKS’ FAC-Code Mass transfer coefficient for Fe(II) species from the oxide–coolant interface to the bulk environment by convection Geometry factor in Siemens-KWU ‘WATHEC’ FAC-Code Dissolution reaction rate constant of magnetite at the oxide–coolant interface Stress intensity factor (LEFM) Stress intensity factor at the onset of SICC crack growth in slow rising load tests with precracked fracture mechanics specimens Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels KIJ Stress intensity factor at the onset of ductile crack growth m Paris law exponent for fatigue and corrosion fatigue n Frequency exponent for corrosion fatigue pH % log[Hỵ] pH value R Load ratio of minimum to maximum load Flow-accelerated corrosion rate RFAC Dissolution rate of magnetite at the Rd oxide–coolant interface Formation rate of magnetite at the Rg metal/oxide interface Mass transport rate of Fe(II) Rm species from the oxide–coolant interface to the bulk environment by convection Re ¼ dur/m Reynolds number Sc ¼ m/( rD) Schmidt number Sh ¼ hd/D Sherwood number t Time T Temperature u Coolant velocity Z Reduction of area in tensile test [X] Concentration of element/species X in water or in alloy a Steam void fraction in EPRI ‘CHECWORKS’ FAC-Code DC ¼ Ceq À Cb Undersaturation in dissolved Fe(II) species Dd Wall thinning DK ¼ KI,max À Stress intensity factor range of KI,min fatigue cycle Upper DK threshold for corrosion DKCF,H fatigue Lower DK threshold for corrosion DKCF,L fatigue DK threshold for fatigue in air DKth, Air Decline time (down-ramp) of DtD fatigue cycle Hold time at maximum load of DtH fatigue cycle Rise time (up-ramp) of fatigue cycle DtR e Mechanical strain Critical strain (e.g., for SICC) ecrit k Specific electrical conductivity m Viscosity of coolant n Loading frequency Upper critical frequency for ncrit,H corrosion fatigue ncrit,L r s scrit t 107 Lower critical frequency for corrosion fatigue Density Mechanical stress Critical stress (e.g., for SCC) Fluid shear stress at pipe wall 5.06.1 Introduction Carbon and low-alloy steels (CS & LAS, Table 1) and their associated weld filler metals are widely used for pressure vessels and piping in both the primary and secondary coolant circuits of watercooled reactors (light water reactors (LWRs) and CANDUs – pressurized heavy water reactors (PHWRs)), as well as in service water systems.1 The main reasons for the use of CS & LAS are their combination of relatively low cost, good mechanical strength and toughness properties in thick sections (hardenability), and good weldability, as well as their good stress corrosion cracking (SCC) resistance in primary coolant environments Compared with austenitic stainless steels and nickel-base alloys, ferritic CS & LAS exhibit only moderate corrosion and irradiation resistance They also show a ductile-to-brittle transition in toughness properties at lower temperatures CS & LAS components in the primary circuit of pressurized water reactors (PWRs) are clad (usually with austenitic stainless steel) and thus not generally come into direct contact with the reactor coolant This is also the case for the reactor pressure vessel (RPV) in boiling water reactors (BWRs), although the RPV head is sometimes left unclad and the cladding has been removed from the blend radius of many RPV feedwater nozzles In BWRs of German and of newer General Electric designs, extensive use is also made of unclad LAS and CS in both the feedwater and steam lines, as well as in the condensate system The primary coolant piping in conventional CANDUs is made exclusively of unclad CS In secondary coolant systems, the steam generator pressure vessel shell is unclad, as are the feedwater, drain, and steam lines CS & LAS pressure-boundary components, in particular in the primary circuit such as the RPV, are very critical systems with regard to plant safety and lifetime (extension) Minimizing corrosion improves plant availability and economics and is also fundamental for safe operation over extended periods of 50–60 years 108 Typical CS & LAS piping and pressure vessel materials in Western LWRs (US designation, according to Section II of ASME BPV Code)  Designation Type Product form Cmax (%) Mn (%) Pmax (%) Smax (%) Simin (%) Cumax (%) Nimax (%) Crmax (%) Momax (%) Vmax (%) YS25 C (MPa) Heat treatment Microstructure SA 106 Gr B CS C–Mn Pipe drawn 0.30 0.29 1.06 0.035a 0.035a 0.10 0.40a 0.40b 0.40b 0.15b 0.08b ! 240 300–400c Normal Ferriticpearlitic SA 333 Gr CS C–Mn Pipe drawn 0.30 0.29 1.06 0.025a 0.025a 0.10 – – – – – ! 240 300–400c Normal Ferriticpearlitic SA 516 Gr 70 CS C–Mn Vessel plate 0.27d 0.79 1.30 0.03a 0.035a 0.13 0.45 – – – – – ! 260 300–400c Normal Ferriticpearlitic SA 533 B Cl.1 LAS Mn–Mo–Ni (R)PV plates 0.25 1.07 1.62 0.12e (0.35) 0.15e (0.35) 0.13 0.45 0.10e 0.37 0.73 – 0.41 0.64 0.05 ! 345 450–550c Q&T Bainitic SA 508 Gr Cl LAS Mn–Mo–Ni (R)PV forging 0.25 1.20 1.50 0.12e (0.25) 0.15e (0.25) 0.15 0.40 0.10e 0.40 1.00 0.25 0.45 0.60 0.05 ! 345 450–550c Q&T Bainitic SA 508 Gr.2 Cl LAS Ni–Mo–Cr (R)PV forging 0.27 0.50 1.00 0.12e (0.25) 0.15e (0.25) 0.15 0.40 0.10e 0.50 1.00 0.25 0.45 0.55 0.70 0.05 ! 345 450–550c Q&T Bainitic a In modern steels, these values are less than 0.015% Combination shall not exceed 1.0% c Typical range d Carbon varies with thickness up to 0.31% e Requirement for core belt region YS ¼ yield strength; Normal ¼ normalized; Q & T ¼ quenched and tempered b Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels Table Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels Consideration of both uniform and flowaccelerated corrosion (FAC) behavior for all unclad surfaces is important for corrosion product transport and deposition (e.g., crud formation on fuel elements) but – together with the assessment of resistance to localized corrosion phenomena such as pitting and environmentally assisted cracking (EAC) – is obviously also required for integrity reasons In the case of EAC, however, safety considerations furthermore require that possible defects extending through the cladding be taken into account, so that the susceptibility of the RPV must be assessed as if no cladding were present Sometimes, thick pads of Alloy 182 have been welded directly onto the RPV to act as attachment points for internal structures; the higher yield strength of Alloy 182, the thicker section and its known SCC susceptibility raise special concerns for these areas In such cases, it is possible that SCC or thermal fatigue of the austenitic alloy will occur such that the crack tip propagates to the interface between the austenitic and ferritic alloys Furthermore, leakage of coolant from the primary circuit in PWRs poses a special hazard for CS & LAS components, since the boric acid it contains can concentrate and lead to uniform corrosion, or ‘wastage,’ of external surfaces This chapter covers both the uniform and localized corrosion behavior of CS & LAS pressureboundary components in the primary (BWR, PWR, and CANDU) and secondary (PWR and CANDU) coolant systems of Western reactors, whereby the discussion is focused on relevant US nuclear codes and rules together with material standards in this area Special emphasis in Sections 5.06.2 and 5.06.3 is placed on FAC and on EAC, both of which have resulted in serious pipe ruptures (FAC) or leaks (EAC) during both nuclear and fossil service in the past In Section 5.06.2, the uniform and boric acid corrosion behavior of CS & LAS, as well as the nature of the protective oxide film on these materials, are summarized first, followed by a condensed review of the FAC behavior of these steels The major factors controlling FAC, the underlying mechanism and predictive models, as well as the relevant service experience and possible mitigation actions are discussed After a brief overview of pitting in CS & LAS in the first part of Section 5.06.3, crack initiation susceptibility conditions and crack growth behavior are discussed in detail for the different types of EAC and compared with the relevant design codes and crack growth disposition curves for CS & LAS This is followed by a review of the mechanistic understanding of EAC and of existing EAC models LWR service 109 experience and mitigation actions with regard to EAC are then summarized and compared with this experimental and theoretical background knowledge Finally, Section 5.06.4 summarizes the major conclusions of this review 5.06.2 Uniform and FlowAccelerated Corrosion 5.06.2.1 Uniform Corrosion Uniform or general corrosion does not normally cause a problem for the structural integrity of CS or LAS components in nuclear coolant systems Corrosion rates in typical circuits are generally of the order of a micrometer per year (1 mm yearÀ1) or less – higher than those of stainless steel or nickelbased alloys, for example, but quite acceptable Around 300  C, uniform corrosion rates of CS & LAS are minimal at a slightly alkaline pH300  C of $6–6.5 (neutral high-purity water has a pH300  C of 5.7) and intermediate dissolved oxygen levels Under some shutdown conditions, however, LWR primary coolant can be aggressive to these materials, in particular in conjunction with increased oxygen levels (e.g., through oxygen ingress from air); below $100  C, corrosion rates may be high Compact, defect-free oxide films grown at higher temperatures during service are kinetically quite stable at lower temperatures and usually provide sufficient protection against uniform corrosion during short shutdown periods Nevertheless, reactor vessels and LAS piping in PWRs are clad with stainless steel, which helps reduce the build-up of crud on fuel and of radiation fields by ensuring a high degree of water purity with a low level of dissolved iron A particular concern in PWRs arises from the leakage of borated coolant from joints such as gasketed flanges and its impingement on components such as flange studs Up to 2001, some 140 leaks had been reported publicly.2 Solid boric acid at room temperature and dilute, deaerated boric acid solutions regardless of temperature have little effect on CS & LAS, but as the boric acid concentrates, corrosion rates up to about mm yearÀ1 may be reached Aerated solutions can be much more aggressive, with the attack increasing with acid concentration Note that as hot coolant escapes to the environment, its boric acid content (which may be nominally 2000 ppm (1 ppm ¼ mg kgÀ1; ppb ¼ mg kgÀ1) or more as elemental boron) concentrates by evaporation At temperatures in the neighborhood of 100  C, which 110 Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels are attained by surfaces impacted by coolant flashing to steam, corrosion rates can reach $250 mm yearÀ1.2 In some situations, flow effects can exacerbate the attack, as described in Section 5.06.2.2 The resistance of CS and LAS to corrosion is dependent upon the protective properties of the oxide film Environments such as boric acid that dissolve or erode the oxide then promote corrosion The predominant oxide on CS and LAS in coolant circuits operating above about 130  C is magnetite – Fe3O4 In deoxygenated alkaline water, the magnetite forms a double layer that has been well characterized in terms of materials performance in boiler systems at temperatures of about 300  C.3 This morphology is found on CS in CANDU primary circuits, and would be found on pressure-vessel steel exposed to PWR primary coolant in the absence of high-alloy cladding The layers are formed by the simple oxidation of the steel by water: Fe ỵ 2H2 O ẳ FeOHị2 þ H2 ½IŠ The nascent hydrogen is absorbed by the metal and diffuses to the exterior Roughly half of the ferrous species (often as the dissolved hydroxide – depending on the pH) are precipitated oxidatively at the metaloxide interface as small crystallites of magnetite, each a few tens of nanometers across, also releasing hydrogen to the coolant: 3FeOHị2 ẳ Fe3 O4 ỵ 2H2 O ỵ H2 ẵII The precise fraction precipitated is determined by the density of the oxide relative to that of the metal, since the inner layer occupies the volume of metal corroded.3 The remainder of the dissolved iron diffuses through the oxide to the bulk coolant and precipitates according to eqn [II] as an outer layer of magnetite crystals, each several micrometers across, again releasing hydrogen to the coolant If metal species other than those of iron originate from alloy components elsewhere in a circulating system, they may coprecipitate and modify the locally formed magnetite An example of double-layer formation is shown in Figure The concentration of dissolved iron in the coolant governs the oxide formation If the coolant is significantly undersaturated in iron, the outer layer cannot precipitate and the inner layer may even dissolve at the oxide–coolant interface In nonisothermal systems, temperature gradients create solubility differences and transport iron around the circuit, modifying the oxide films accordingly (the same phenomenon transports different oxides around circuits containing other Coolant flow Precipitation Dissolution Outer oxide o/s interface Inner oxide m/o interface Corroding metal Figure Schematic of double layer oxide formation on carbon steel in high-temperature water materials, such as the nickel-base alloys in PWRs) Thick films may also spall and release oxide particles to be distributed by the coolant In circuits connected to the reactor core, oxide transport may create deposits on the fuel, impeding heat transfer and leading to increased radiation fields around out-of-core components (note that the nickel-base alloys and stainless steel in PWRs can produce deposits derived from nickel ferrite, NiFe2O4; on high-burnup fuel undergoing subcooled boiling, these can harbor boron from the coolant and provoke shifts in the neutron flux, as well as affect radiation fields) Evolved hydrogen also affects magnetite solubility (by the one-third power of the concentration – as indicated by eqn [II]) Such increased solubility at the metal–oxide interface has been invoked as the reason for the lack of precipitation within pores as iron diffuses to the oxide–coolant interface.4 Magnetite films formed on steel surfaces that are pressure boundaries, where the hydrogen evolved by eqn [I] continuously effuses through the metal, tend to have a more adherent inner layer of larger crystallites than those formed on totally immersed surfaces such as test coupons, where the evolved hydrogen can only diffuse through the oxide to the bulk coolant once the metal becomes saturated.5 Under neutral oxidizing conditions, magnetite is still the predominant base oxide formed on steels.6 However, since dissolved oxygen becomes the oxidant rather than water, hydrogen generation is suppressed and the basic oxidation reactions become: 2Fe ỵ 2H2 O ỵ O2 ẳ 2FeOHị2 6FeOHị2 ỵ O2 ẳ 2Fe3 O4 ỵ 6H2 O ẵIII ẵIV The oxide layers – especially the outer one – then tend to contain the more-oxidized forms maghemite and/or hematite (both of formula Fe2O3), particularly in BWR circuits.7 Under reactor coolant conditions, corrosion rates and oxide solubilities under Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels oxidizing conditions are generally substantially lower than those under reducing conditions At high oxygen levels, however, the risk for pitting and EAC increase significantly (see Section 5.06.3) The forms of oxide that are thermodynamically stable under various conditions in coolant circuits are indicated by Pourbaix diagrams, which plot the equilibrium potentials of the oxidizing–reducing reactions against pH; the higher the potential, the more oxidizing the environment For dissolved species, the equilibria and therefore the lines in the diagram are dependent upon the concentration; when illustrating corrosion situations, a concentration of 10À6 M or less is often assumed It should be borne in mind, therefore, that such diagrams are mainly indicative in nature and illustrate the possibilities of species formation without taking account of reaction kinetics Figure 2, adapted from Beverskog and Puigdomenech,8 is an example for species pertinent to steel at 310  C, where a species concentration of 10À8 M is representative The hydrogen line in the figure represents the equilibrium: 2H2 O ỵ 2e ẳ 2OH ỵ H2 5.06.2.2 ẵV Flow-Accelerated Corrosion 5.06.2.2.1 Controlling factors Flow-accelerated (or -assisted) corrosion (FAC), sometimes called erosion–corrosion (EC) in older literature, −0.2 Fe2O3 −0.4 Fe(OH)4− −0.6 Fe3O4 E(v) Fe(OH)+ Fe(OH)2 −0.8 Hydr ogen line Fe(OH)3− −1 Fe −1.2 −1.4 6.5 7.5 pH310 ЊC 8.5 9.5 Figure Pourbaix diagram for iron at 10À8 m at 310  C Reproduced from Beverskog, B.; Puigdomenech, I Corros Sci 1996; 38(12): 2121–2135 111 is essentially the dissolution and erosion of the normally protective oxide film on CS (or LAS with a Cr-content < $0.2 wt%), exacerbated by fluid flow effects, resulting in excessive corrosion rates and substantial pipe wall thinning Nowadays, the term EC implies the involvement of a significant mechanical component as an abrasive (e.g., by dispersed solid particles in the liquid phase) or cavitation-induced (mechanical) removal of surface material; it should therefore be differentiated from FAC, which is primarily caused by a flow-induced increase in the mass transfer of dissolving and reacting (corrosive) species at high-flow or highly turbulent locations, although fluid shear stress on the oxide film at the material surface may also make substantial contributions to the damage in some situations FAC is a pervasive problem in most types of steam-raising system and has caused feedwater line ruptures, occasionally with fatal consequences, in both fossil and nuclear plants.9,10 In primary coolant systems also, less serious (though costly) FAC occurs chronically in the CS outlet feeders of conventional CANDUs,11 and flow effects are implicated in the corrosion of PWR pressure-vessel steel by borated coolant leaking through cracked penetrations in the RPV head.12 FAC thus occurs in the regions of high turbulence in both single and two-phase flows, but never in systems with dry steam FAC depends on hydrodynamics (mainly steam quality, flow rate, fluid shear stress at the wall, turbulence intensity, and mass transfer coefficient), environmental factors (mainly temperature, pH, dissolved oxygen, hydrogen, and iron concentrations) and material parameters (metal composition – Mo, Cu and, in particular, Cr content).9 The critical parameter combinations for the occurrence of FAC in feedwater systems and the main parameter effects are schematically summarized in Figure The conditions leading to increased FAC rates are usually related to regions with turbulent flow, to low electrochemical corrosion potentials ECP (i.e., to chemically reducing conditions), and to low iron concentrations in the water (Figures and 4) Depending on the pH, the maximum FAC rates occur at about 130  C in single-phase flow, and at about 180  C in two-phase flow (in the latter, it is the condition in the liquid layer at the steel surface that controls the FAC rate, but this is difficult to measure or predict) Note that FAC can still be a problem at other temperatures, even though rates are lower For example, feeder FAC in CANDU primary coolants occurs at 300–310  C at the core outlet, and FAC is also significant in feedwater systems at the low Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels Log FAC rate 112 pH Hydrodynamics Hydrodynamics Hydrodynamics pH • Shear stress at surface • Flow rate • Turbulent intensity • Mass transfer coefficient Flow Critical conditions for high FAC risk in feedwater ∼150 °C Material Environment • [Cr] in metal < 0.2% Log FAC rate Log FAC rate pH • Low [Fe] • pH < 9.2 • 120 °C < T < 180 °C • [O2] < 2–40 ppb pH ∼0.2% Cr pH T ∼40 ppb Log FAC rate pH ∼1−2 ppb pH Log FAC rate pH [Cr] in metal ∼130 °C pH pH [O2] Figure Critical parameter combinations for flow-accelerated corrosion (derived from Uchida, S et al In: Proceedings of the 13th International Conference on Environmental Degradation of Materials in Nuclear Power Systems, CD-ROM Whistler, British Columbia, Canada, 19–23 August, King, P., Allen, T., Busby, J., Eds.; Toronto, ON: The Canadian Nuclear Society, 2007) and major parameter effects on flow-accelerated corrosion under feedwater conditions temperature of condensate extraction Specific geometries like elbows, bends, protruding weld roots, orifices, and valves cause local turbulence, which significantly increases FAC rates at, or immediately downstream of, the location concerned Systems such as the moisture-separator/reheater drain lines, where steam has condensed and relatively iron-free water is flowing, are particularly susceptible In primary coolant systems, there is the desire to keep iron concentrations low to prevent crud build-up and radiation transport problems, hence the frequent use of highalloy materials that are resistant to FAC as cladding However, it must be recognized that a recirculating system will always tend toward equilibrium; in other words, dissolved iron concentrations on average will vary around solubility values, depending upon oxide dissolution and precipitation kinetics, temperature gradients around the circuit, and the capacity of sinks such as the purification circuit Most studies of FAC have been performed under feedwater conditions, which generate high rates of attack that can reach several millimeters per year in some situations Neutral chemistry, low-oxygen conditions at about 140  C, as may be found in BWR feedtrains, can give high FAC rates, so dual-cycle PWRs or PHWRs routinely add an amine such as ammonia to raise the pH in the secondary coolant circuit The actual pH employed depends upon the materials of construction; for all-ferrous feedtrains, a pH25  C from 9.3 to 9.6 is usually specified, but the value is kept below 9.2 to avoid excessive corrosion of copper-base alloys, if these are present Also, to achieve a more even distribution of additive around the circuit, an amine (such as morpholine) with a coefficient of distribution between the steam and liquid phases closer to unity than that of ammonia may be used Oxygen dissolved in the coolant is also a powerful inhibitor of FAC; it has been added routinely to feedwater systems in BWRs and certain fossil boilers for some time Depending upon the rate of attack, levels of oxygen from a few ppb to several tens of ppb are sufficient to stifle FAC completely Maintaining a dissolved oxygen content >$30 ppb, which raises the corrosion potential ECP in the feedwater system above the Fe3O4/Fe2O3 phase boundary in the Pourbaix diagram in Figure 2, is particularly crucial in BWRs operating on hydrogen water chemistry Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels 113 St37.2 (A414 Gr B-carbon steel) Flow (kg h−1) 983 491 15Mo3 (A161 Gr T1–0.5% Mo) 907 756 605 378 302 227 15NiCuMoNb5 13CrMo44 (A213 Gr T12–1% Cr, 0.5% Mo) 10CrMo910 (A213 Gr T22–2.2% Cr, 1% Mo) 3000 4.0 pH = 9.04 pH = 7.0 2.0 Increasing flow Loss rate (mm year−1) 3.0 1.0 Specific material wear rate (μg cm−2 h−1) 1000 300 100 30 10 3.0 1.0 0.3 90 (a) 100 110 120 130 140 150 160 170 Temperature (ЊC) 0.1 50 (b) 100 150 200 250 300 Temperature (ЊC) Figure Effect of temperature, flow rate (a) (data from Bignold, G J et al In: Proceedings of the International Specialist’s Meeting on Erosion-Corrosion of Steels in High-Temperature Water and Wet Steam, Les Renardie´res, France, 11–12 May; EDF: France, 1982) and material (b) (data from Heitmann, H G.; Schub, P In: Proceedings of the Third Meeting on Water Chemistry of Nuclear Reactors, pp 243–252, Bournemouth, UK, October; British Nuclear Engineering Society (BNES): London, UK, 1983) on single-phase flow-accelerated corrosion under different flow and chemistry conditions Reproduced from Dooley, R B Power Plant Chem 2008, 10(2), 68–89 (HWC) with high rates of hydrogen injection into the feedwater If HWC is combined with noble metal chemical addition (NMCA), the FAC risk is reduced, since much lower hydrogen injection rates are then adequate to mitigate SCC in stainless steel recirculation piping and reactor internals (Recombination of hydrogen and oxygen to lower the ECP requires the radiation fields present in the RPV.) Oxygen levels significantly above $50 ppb may increase the risk of strain-induced corrosion cracking and corrosion fatigue in CS & LAS feedwater piping (see Section 5.06.3) Furthermore, the deliberate addition of oxygen to feedwater systems in dual-cycle reactors may pose problems, since residual oxygen entering the steam generators can provoke SCC of the high-alloy steam-generator tubes Nevertheless, severe FAC of components in the feed train of advanced gas-cooled reactors (AGRs) has been successfully mitigated since the early 1980s by oxygen additions.13 Material properties have a significant impact on FAC rates, but typically the plant operator has no control over this (unless a replacement of piping is an option) Certain elements in the steel can act to retard FAC, as mentioned earlier; for example, chromium is particularly effective and a concentration of 0.1% in the metal reduces FAC in 180  C ammoniated water and water–steam mixtures at pH25  C by about 70%.14 Moreover, under CANDU primary coolant conditions of 310  C and pH25  C 10.5 (adjusted with lithium), increasing the chromium content of SA-106 Grade B CS from 0.019% to 0.33% reduces FAC by about 50%.11 114 Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels 5.06.2.2.2 Mechanisms and models As with uniform corrosion (discussed in Section 5.06.2.1), FAC is governed by the ability of the oxide film to protect the metal Magnetite forms on the steel at the metal–oxide interface and is degraded at the oxide–coolant interface by fluid flow effects and by dissolution according to the general equation [VI] (which indicates the dependence of the dissolved species on pH under reducing conditions and which is equivalent to eqn [II] for b ¼ 2) The turbulence in the coolant and the solubility of the magnetite are then paramount in determining the severity of the attack Fe3 O4 ỵ 32 bịHỵ ỵ H2 ẳ 3ẵFeOHịb 2bịỵ ỵ 3bịH2 O ẵVI with b ¼ 0, 1, 2, or Mass transfer is often assumed to control the mechanism.15 This derives from the postulate that the magnetite film attains a steady-state thickness as it dissolves at the rate Rd at its outer surface in coolant undersaturated in dissolved iron and forms continuously at the metal–oxide interface at the same rate Rg Since the magnetite formation at the metal–oxide interface accounts for only about half of the corroded metal, the other half diffuses through the magnetite to the oxide–coolant interface, and with the iron from the magnetite dissolution is transported to the bulk coolant at the rate Rm The FAC rate RFAC is thus twice the dissolution rate Rd of the magnetite at the oxide–coolant interface This concept of two processes in series – dissolution Rd and mass transfer Rm– leads to the equation for the steady-state FAC rate RFAC¼ dm/dt ¼ Rm¼ 2Rd with all the variables in equivalent units of iron per unit surface and time Steady-state assumption for the serial process: Rg ẳ Rd ẳ 0:5Rm ẵ1 Dissolution rate of magnetite at the oxide–coolant interface according to eqn [VI] (assuming firstorder kinetics): Rd ¼ 0:5 Á dm=dt ¼ kd Ceq Cị ẵ2 where kd is the dissolution reaction rate constant, C is the concentration of Fe(II) species at the oxide–coolant interface, and Ceq is their equilibrium concentration according to eqn [VI], which corresponds to their maximum solubility in the coolant Transport of Fe(II) species from the oxide–coolant interface to the bulk environment by turbulent mass transfer: Rm ẳ h C Cb ị ẵ3 where Cb is the concentration of Fe(II) species in the bulk coolant and h is the mass transfer coefficient, which is dependent on flow conditions and geometry From eqns [1]–[3] it follows that: RFAC ¼ h Á kd Á DC ð0:5 Á h ỵ kd ị ẵ4 where DC ẳ Ceq Cb Þ is the undersaturation in iron Models of FAC are based on the principles behind eqn [4] We expect that kd strongly increases with temperature according to an Arrhenius law for a thermally activated process (although there are no data to confirm this over the temperature ranges of interest), whereas h only shows a moderate increase through the temperature dependence of the properties in eqn [6] If mass transfer controls, h is small compared with kd (h ( kd) and the equation reverts to: RFAC ẳ hDC ẵ5 For a coolant of constant conditions containing little or no dissolved iron (i.e., Cb % 0), the driving force DC approaches a constant value – the solubility of the oxide, Ceq – and RFAC varies as the mass transfer coefficient (which increases with increasing flow rate and turbulence) The mass-transfer model then implies that the effects of materials composition and coolant chemistry on FAC rate are brought about by their effects on oxide solubility (Figure 5) According to eqn [VI], the saturation concentration or solubility Ceq depends on temperature, pH and H2 concentration by simple chemical equilibrium thermodynamics Accordingly, the effect of chromium in the steel can be attributed to the relative stability of mixed oxides containing chromium (iron chromite, FeCr2O4, for example, is virtually insoluble in reducing coolant and accounts for the protection afforded by stainless steel and similar alloys) As corrosion proceeds and the magnetite dissolves, chromium is not leached out in concert but continually concentrates in the film It is interesting to note that the inhibition occurs immediately at the start of exposure and continues at about the same level, suggesting that the mechanism is the rapid formation at the metal–oxide interface of a more protective layer of oxide that is maintained throughout exposure.16 It appears that the higher the chromium content of the steel, the more protective that Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels 100 10−1 10−2 101 Δtrise = 14 s R = 0.95 0.018% S 100 10−1 10−2 43 s −6 N 20 e = s ca t ise e Δ r od 65 C = R HWC, 274/288 ЊC 64 s N- 14 e = s e ca Δt ris de 95 Co = R e rv 10−3 ME AS X I ‘W et, ME ’R XI Ն0 ‘W 65 et, ’R Յ 0.2 101 N-643, EAC-curve for S > 0.013 % S, R = 0.65, ΔtR = 100 s N-643, EAC-curve for S > 0.013 % S, R = 0.65, ΔtR > s N-643, non-EAC curve, R = 0.65 ASME XI ‘Wet,’ R = 0.65 ASME XI ‘Air,’ R = 0.65 AS da/dN (μm per cycle) 102 da/dNCF (μm per cycle) 128 ir’ E XI ‘A M AS Δtrise = 200–2000 s R = 0.3–0.7 0.004–0.018% S cu AC -E on N 10−3 10−4 10 100 10 100 (a) Stress intensity factor amplitude ΔK (MPa m1/2) (b) Stress intensity factor amplitude ΔK (MPa m1/2) Figure 17 ASME XI ‘Air’ and ‘Wet’ curves and Code Case N-643 for high and low-sulfur steels (a) and comparison of cyclic corrosion fatigue crack growth rates under hydrogen water chemistry conditions for different loading conditions with the corresponding curves (b) Reproduced from Seifert, H P.; Ritter, S Corros Sci 2008, 50, 1884–1899 where environmental effects on fatigue crack growth can be neglected or excluded are not defined in the present ASME Section XI Code.38,47 A more specific Code Case N-64357 for fatigue crack growth in CS & LAS exposed to PWR primary environments has been developed since 2000 and may be used as an alternative to the ASME XI wet reference fatigue crack growth rate curves for this specific environment (Figure 17) Depending on system conditions, the Code Case N-643 procedure predicts either lower or higher crack growth rates than the general Section XI approach The main advantage of this newer Code Case is that it contains criteria for the onset/cessation of EAC and that it better reflects the experimentally observed cracking behavior in PWR environments, since it considers steel sulfur and frequency–loading-rate effects to a certain extent However, this approach has not (yet) found general acceptance, primarily because of the difficulty of defining appropriate rise times for actual plant transients and the complications involved in including these in component analyses In a similar way to that discussed earlier for fatigue life design, conservatism in fatigue flaw tolerance evaluations may arise from (a) the fatigue flaw tolerance evaluation procedures themselves (e.g., by the use of design transients) and/or (b) the fatigue crack growth rate curves As discussed in Seifert and Ritter,38,47 the current ASME XI wet curves conservatively cover the CF crack growth rate laboratory data under most combinations of loading, environmental, and material parameters Even under highly oxidizing BWR/NWC conditions, they are only significantly exceeded under some very specific (but plantrelevant) circumstances (Figures 14 and 17), which have caused some isolated CF cracking incidents in the past The Section XI curves might therefore be regarded as an adequate, general bounding approach under most system conditions, but they not realistically describe and reflect the experimentally observed CF crack growth behavior of CS & LAS in oxygenated high-temperature water The curves predict crack growth rates which are either significantly too high (e.g., n 10À2 Hz and ECP < À200 mVSHE) or too low (e.g., n 10À2 Hz and ECP > mVSHE) Furthermore, system conditions or thresholds (e.g., n > Hz), where environmental effects on fatigue crack growth can be neglected, or even excluded, are not defined in ASME XI 5.06.3.2.3 Stress corrosion cracking 5.06.3.2.3.1 conditions Initiation and susceptibility Table shows an assessment scheme according to Hickling,58 based on both laboratory and field experience, for SCC initiation susceptibility and crack growth in CS & LAS at normal strength levels under BWR/NWC conditions Initiation of propagating SCC cracks from smooth, defect-free surfaces under static load in high-purity water is only observed for the following conjoint conditions: stresses at the water-wetted surface above the high-temperature yield strength, quasi-stagnant flow conditions, and dissolved oxygen contents !$0.2 ppm Furthermore, if complete exhaustion of low-temperature creep is allowed to occur before the specimens are exposed to high-purity, high-temperature water, no SCC is observed, thus indicating ‘nonclassical’ SCC behavior and confirming the importance of slow dynamic surface straining Assessment scheme for SCC susceptibility of CS & LAS O2 (ppm) Operating medium: HT water or steam condensate with T > 170  C Flow conditions k (mS cmÀ1) Crack initiation by SCC? Derivation Crack growth by SCC? Derivation 10 mm year–1) at KI < 60 MPa m½ T = 274/288 ЊC NWC 200 Continuous operation allowed Prompt shut-down EPRI action Level limit –400 EPRI action Level limit –200 HWC Corrosion potential ECP (mVSHE) No SCC (‘EPRI action level limit’ of EPRI BWR water chemistry guidelines64) or load transients not covered by fatigue evaluation procedures These curves are currently undergoing minor revisions The conservative character of the BWRVIP-60 disposition line for SCC crack growth in CS & LAS under NWC conditions has been confirmed for temperatures in the range of 274–288  C and 132 Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels both base and weld filler/heat-affected zone (HAZ) materials (Vickers hardness < 350 HV5, 0.02 wt% S) if the water chemistry is maintained within current BWR/NWC operational practice ( 0.95) or during chloride transients (! EPRI action level limit), even at fairly low stress intensity values around 30 MPa m1/2 (see Figure 19) On the other hand, ‘line 2’ seems to cover even very severe sulfate transients above the EPRI action level limit.38,60 At low ECP, that is, in the case of BWR/HWC or PWR conditions, line appears conservatively to cover the SCC crack growth in LAS, even under the otherwise critical conditions mentioned earlier for more oxidizing, environments.38,60 FRAD-EAC-mechanism 5.06.3.2.4 EAC mechanisms and models 5.06.3.2.4.1 EAC crack growth mechanism For the case of EAC in CS & LAS in hightemperature water, the film rupture/anodic dissolution (FRAD), also referred to as slip dissolution/ oxidation, and hydrogen-assisted EAC (HAEAC) cracking mechanisms have been proposed in the literature (see Figure 20) Either of these may be superimposed on pure mechanical fatigue In the FRAD mechanism, the protective oxide film formed on CS & LAS in high-temperature water is ruptured by plastic straining at the crack-tip The crack-tip then advances by anodic dissolution of the bare metal matrix Anodic dissolution is slowed down and finally stopped by the nucleation and reformation of the oxide film (‘repassivation’) Thus, continued crack advance will depend on a further oxide rupture process due to the action of a strain rate at the cracktip The crack propagation rate is controlled by both anodic dissolution/repassivation kinetics and the frequency of oxide film rupture at the strained crack-tip The first part is governed by the chemical composition of the local crack-tip electrolyte and the material The second part is determined by the fracture strain of the oxide film and the crack-tip strain rate.63 Hydrogen-assisted EAC mechanism syy Crack surface covered by an oxide film Bare surface of metal generated by oxide film rupture and anodic dissolution of the metal Elastic–plastic stress distribution Acidic oxygen free water Oxide layer Void 8 MnS MnS Fe2+ Local anodic reaction Repassivation: Oxide nucleation and growth ad Fe → Fe2+ + 2e Hydrolysis and generation of H+ Fe2+ + H2O → FeOH+ + Η+ Liquid phase transport Local cathodic reaction H+ + e → H MnS + 2H+ →H2S + Mn2+ Hydrogen absorption Hydrogen transport in lattice Hydrogen trapping on inclusions Hydrogen-induced cracking Linkage of microcracks to main fracture Figure 20 Schematic illustrations of film rupture/anodic dissolution (left) and of hydrogen-assisted environmentally assisted cracking mechanisms for carbon and low-alloy steel in high-temperature water Adapted from Ford, F P J Pressure Vessel Technol 1988, 110, 113128; Haănninen, H et al Corros Sci 1983, 23, 663–679 Corrosion and Environmentally-Assisted Cracking of Carbon and Low-Alloy Steels Oxide film rupture–repassivation events at the strained crack-tip are also involved in the HAEAC model, but here, hydrogen-induced microcrack formation ahead of the crack-tip and linkage of these microcracks to the main crack are the prime sources of EAC crack growth, and result in discontinuous crack propagation The hydrolysis of metal cations from anodic dissolution is an important source of hydrogen, but – in contrast to the FRADmechanism – anodic dissolution does not contribute significantly to crack advance Hydrogen transport in the electrolyte and in the metal lattice is believed to be fast Therefore, the generation of bare metal surface by film rupture and film reformation may be the rate-controlling steps and explain the strain rate dependence of EAC.65,66 The FRAD and HAEAC mechanisms may be simultaneously active and controlled by the same rate-limiting steps (e.g., oxide film rupture rate or repassivation kinetics) Both mechanisms are able to explain the observed, dominant effect of strain rate and of MnS inclusions on EAC The EAC behavior of CS & LAS in high-temperature water can best be rationalized by a superposition/combination of the FRAD and HAEAC mechanisms At lower temperatures (800 MPa/>350 HV5) and high strain rates (>10À3 sÀ1), hydrogen effects are more HCl H2SO4 NaCl Role of MnS inclusions Sulfur-anions as HSÀ, S2À, and SO2À may signifi4 cantly retard repassivation after oxide film rupture and therefore increase crack advance by anodic dissolution in the FRAD model.38,63 Retarded repassivation of the film-free surface and adsorbed HSÀ, S2À, or H2S also increase hydrogen absorption into the metal lattice and therefore favor HAEAC.38,65,66 Furthermore, the dissolution of MnS is a further potential source of hydrogen, and MnS inclusions in the region of maximum hydrostatic stress ahead of the crack-tip may act as strong hydrogen traps and thus HAEAC microcrack initiation sites.38,65,66 The effects of steel sulfur content are synergistic with environmental variables, such as (sulfur-) anionic impurities in the bulk environment, ECP (dissolved oxygen content), and flow rate (Figure 21).38,67 This is believed to be due to the creation of a sulfur-rich crack-tip environment responsible for EAC, which arises both from the dissolution of MnS intersected by the growing crack and by the transport of sulfuranions by migration/diffusion/convection within the crack enclave.38 SSRT, 1ϫ10–6s–1 hpw, 288 ЊC Low-flow autoclave –100 Corrosion potential ECP (mVSHE) Corrosion potential ECP (mVSHE) 5.06.3.2.4.2 Filled symbols TGSCC Open symbols no SCC –100 –200 pronounced At high temperatures (!150  C) and/or lower yield strength/hardness levels and slow strain rates (