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

Carbon and Low-Alloy Steels

University of New Brunswick, Fredericton, NB, Canada

ß 2012 Elsevier Ltd All rights reserved.

5.06.3.2.2 Corrosion fatigue and strain-induced corrosion cracking 123

5.06.4.2 Localized Corrosion and Environmentally Assisted Cracking 139

Abbreviations

AC Content of Cr, Mo and Cu in alloy in

EPRI ‘CHECWORKS’ FAC-Code

AGR Advanced gas-cooled reactor

ANL Argonne National Laboratory, USA

ASME American Society of Mechanical

Engineers

ASME BPV ASME Boiler and Pressure Vessel

Code

ASME III Section III of ASME BPV Code

ASME XI Section XI of ASME BPV Code

ASTM American Society of Testing and

Materials Standards

BWR Boiling water reactor

BWRVIP Boiling Water Reactor Vessel and

Internals Program

BWRVIP-60 Basis document for SCC

disposition lines for low-alloy

steels

CANDUW CANada Deuterium Uranium,

PHWR developed by Atomic Energy

DCPD (Reversed) direct current potential

drop crack length measurement method

DH Dissolved hydrogen (concentration)

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

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)

Q þT 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

SSR(T) Slow strain rate (test)

SSY Small-scale yielding

UTS Ultimate tensile strength

VGB German Association of Large Power

Plant Operators

Symbols

C Concentration of Fe(II) species at

the oxide–coolant interface

C b Concentration of Fe(II) species in

the bulk coolant

per fatigue cycle in temperature water da/dtAir¼

Dt R

Time-based corrosion fatigue crack growth rate in high- temperature water da/dt SCC SCC crack growth rate da/dt SICC SICC crack growth rate dCOD LL /dt Crack-opening displacement rate

in slow rising load or displacement test

d e/dt Strain rate (sometimes locally at

crack-tip)

d e/dt crit Critical strain rate (e.g., for SICC)

dKI/dt Stress intensity factor rate in slow

rising load or displacement test

E A Arrhenius activation energy of

thermally activated process

F en Environmental correction factors,

ratio of fatigue life in air at room temperature to that in water at service temperature

‘CHECWORKS’ FAC-Code

h Mass transfer coefficient for Fe(II)

species from the oxide–coolant interface to the bulk environment

by convection

k c Geometry factor in Siemens-KWU

‘WATHEC’ FAC-Code

k d Dissolution reaction rate constant

of magnetite at the oxide–coolant interface

K I Stress intensity factor (LEFM)

K I,i Stress intensity factor at the onset

of SICC crack growth in slow rising load tests with precracked fracture mechanics specimens

K IJ Stress intensity factor at the onset

of ductile crack growth

m Paris law exponent for fatigue and

R FAC Flow-accelerated corrosion rate

R d Dissolution rate of magnetite at the

oxide–coolant interface

Rg Formation rate of magnetite at the

metal/oxide interface

R m Mass transport rate of Fe(II)

species from the oxide–coolant

interface to the bulk environment

DK th, Air DK threshold for fatigue in air

Dt D Decline time (down-ramp) of

e crit Critical strain (e.g., for SICC)

k Specific electrical conductivity

s crit Critical stress (e.g., for SCC)

t Fluid shear stress at pipe wall

Carbon and low-alloy steels (CS & LAS, Table 1)and their associated weld filler metals are widelyused for pressure vessels and piping in both theprimary and secondary coolant circuits of water-cooled reactors (light water reactors (LWRs) andCANDUs – pressurized heavy water reactors(PHWRs)), as well as in service water systems.1The main reasons for the use of CS & LAS aretheir combination of relatively low cost, goodmechanical strength and toughness properties inthick sections (hardenability), and good weldability,

as well as their good stress corrosion cracking (SCC)resistance in primary coolant environments Com-pared with austenitic stainless steels and nickel-basealloys, ferritic CS & LAS exhibit only moderatecorrosion and irradiation resistance They also show

a ductile-to-brittle transition in toughness properties

at lower temperatures

CS & LAS components in the primary circuit ofpressurized water reactors (PWRs) are clad (usuallywith austenitic stainless steel) and thus do not gener-ally come into direct contact with the reactor coolant.This is also the case for the reactor pressure vessel(RPV) in boiling water reactors (BWRs), althoughthe RPV head is sometimes left unclad and the clad-ding has been removed from the blend radius of manyRPV feedwater nozzles In BWRs of German and ofnewer General Electric designs, extensive use is alsomade of unclad LAS and CS in both the feedwater andsteam lines, as well as in the condensate system Theprimary coolant piping in conventional CANDUs ismade exclusively of unclad CS In secondary coolantsystems, the steam generator pressure vessel shell isunclad, as are the feedwater, drain, and steam lines

CS & LAS pressure-boundary components, in ticular in the primary circuit such as the RPV, arevery critical systems with regard to plant safety andlifetime (extension) Minimizing corrosion improvesplant availability and economics and is also funda-mental for safe operation over extended periods of50–60 years

par-Table 1 Typical CS & LAS piping and pressure vessel materials in Western LWRs (US designation, according to Section II of ASME BPV Code)

form

Cmax(%)

Mn (%)

Pmax(%)

Smax(%)

Simin(%)

Cumax(%)

Nimax(%)

Crmax(%)

Momax(%)

Vmax(%)

YS25C(MPa)

Heat treatment

structure

Micro-SA 106 Gr B CS Pipe drawn 0.30 0.29 0.035a 0.035a 0.10 0.40a 0.40b 0.40b 0.15b 0.08b  240 Normal

a In modern steels, these values are less than 0.015%.

b 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.

Consideration of both uniform and

flow-accelerated 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

obvi-ously 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

susceptibil-ity 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

loca-lized corrosion behavior of CS & LAS

pressure-boundary 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 inSections 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

pre-dictive models, as well as the relevant service

experi-ence and possible mitigation actions are discussed

After a brief overview of pitting in CS & LAS in the

first part ofSection 5.06.3, crack initiation

suscepti-bility conditions and crack growth behavior are

dis-cussed 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

understand-ing of EAC and of existunderstand-ing EAC models LWR service

experience and mitigation actions with regard toEAC are then summarized and compared with thisexperimental and theoretical background knowledge.Finally, Section 5.06.4 summarizes the major con-clusions of this review

Flow-Accelerated Corrosion5.06.2.1 Uniform CorrosionUniform or general corrosion does not normallycause a problem for the structural integrity of CS

or LAS components in nuclear coolant systems.Corrosion rates in typical circuits are generally ofthe order of a micrometer per year (1mm year1) orless – higher than those of stainless steel or nickel-based alloys, for example, but quite acceptable.Around 300C, uniform corrosion rates of CS &LAS are minimal at a slightly alkaline pH300C of

6–6.5 (neutral high-purity water has a pH300Cof5.7) and intermediate dissolved oxygen levels Undersome shutdown conditions, however, LWR primarycoolant can be aggressive to these materials, in par-ticular in conjunction with increased oxygen levels(e.g., through oxygen ingress from air); below

100C, corrosion rates may be high Compact,defect-free oxide films grown at higher temperaturesduring service are kinetically quite stable at lowertemperatures and usually provide sufficient protectionagainst uniform corrosion during short shutdown peri-ods Nevertheless, reactor vessels and LAS piping inPWRs are clad with stainless steel, which helps reducethe build-up of crud on fuel and of radiation fields byensuring a high degree of water purity with a low level

of dissolved iron

A particular concern in PWRs arises from theleakage of borated coolant from joints such as gas-keted flanges and its impingement on componentssuch as flange studs Up to 2001, some 140 leaks hadbeen reported publicly.2 Solid boric acid at roomtemperature and dilute, deaerated boric acid solu-tions regardless of temperature have little effect

on CS & LAS, but as the boric acid concentrates, sion rates up to about 1 mm year1 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

corro-2000 ppm (1 ppm¼ 1 mg kg1; 1 ppb¼ 1 mg kg1) ormore as elemental boron) concentrates by evaporation

At temperatures in the neighborhood of 100C, which

are attained by surfaces impacted by coolant flashing

to steam, corrosion rates can reach250 mm year1.2

In some situations, flow effects can exacerbate the

attack, as described inSection 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 130C 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 300C.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:

Feþ 2H2O¼ FeðOHÞ2þ H2 ½I

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

metal-oxide interface as small crystallites of magnetite, each

a few tens of nanometers across, also releasing

hydro-gen to the coolant:

3FeðOHÞ2¼ Fe3O4þ 2H2Oþ H2 ½II

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

dif-fuses through the oxide to the bulk coolant and

pre-cipitates 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 inFigure 1

The concentration of dissolved iron in the coolant

governs the oxide formation If the coolant is

signifi-cantly 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

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 toincreased radiation fields around out-of-core compo-nents (note that the nickel-base alloys and stainlesssteel in PWRs can produce deposits derived fromnickel ferrite, NiFe2O4; on high-burnup fuel under-going subcooled boiling, these can harbor boron fromthe coolant and provoke shifts in the neutron flux, aswell as affect radiation fields)

Evolved hydrogen also affects magnetite solubility(by the one-third power of the concentration – asindicated byeqn [II]) Such increased solubility at themetal–oxide interface has been invoked as the reasonfor the lack of precipitation within pores as iron dif-fuses to the oxide–coolant interface.4Magnetite filmsformed on steel surfaces that are pressure boundaries,where the hydrogen evolved byeqn [I]continuouslyeffuses through the metal, tend to have a more adher-ent inner layer of larger crystallites than those formed

on totally immersed surfaces such as test coupons,where the evolved hydrogen can only diffuse throughthe oxide to the bulk coolant once the metal becomessaturated.5

Under neutral oxidizing conditions, magnetite isstill the predominant base oxide formed on steels.6However, since dissolved oxygen becomes the oxi-dant rather than water, hydrogen generation is sup-pressed and the basic oxidation reactions become:

2Feþ 2H2Oþ O2¼ 2FeðOHÞ2 ½III6FeðOHÞ2þ O2¼ 2Fe3O4þ 6H2O ½IVThe oxide layers – especially the outer one – thentend to contain the more-oxidized forms maghemiteand/or hematite (both of formula Fe2O3), particu-larly in BWR circuits.7 Under reactor coolant con-ditions, corrosion rates and oxide solubilities under

Coolant flow

Corroding metal Inner oxide

Outer oxide

Dissolution

m/o interface o/s interface Precipitation

Figure 1 Schematic of double layer oxide formation on carbon steel in high-temperature water.

oxidizing conditions are generally substantially lower

than those under reducing conditions At high oxygen

levels, however, the risk for pitting and EAC increase

significantly (seeSection 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

equi-librium potentials of the oxidizing–reducing reactions

against pH; the higher the potential, the more

oxidiz-ing the environment For dissolved species, the

equili-bria and therefore the lines in the diagram are

dependent upon the concentration; when illustrating

corrosion situations, a concentration of 106M 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,8is an example for

species pertinent to steel at 310C, where a species

concentration of 108M is representative The

hydro-gen line in the figure represents the equilibrium:

2H2Oþ 2e¼ 2OHþ H2 ½V

5.06.2.2 Flow-Accelerated Corrosion

5.06.2.2.1 Controlling factors

Flow-accelerated (or -assisted) corrosion (FAC),

some-times called erosion–corrosion (EC) in older literature,

is essentially the dissolution and erosion of the normallyprotective oxide film on CS (or LAS with a Cr-content

< 0.2 wt%), exacerbated by fluid flow effects, ing in excessive corrosion rates and substantial pipe wallthinning Nowadays, the term EC implies the involve-ment of a significant mechanical component as an abra-sive (e.g., by dispersed solid particles in the liquid phase)

result-or cavitation-induced (mechanical) removal of surfacematerial; it should therefore be differentiated from FAC,which is primarily caused by a flow-induced increase inthe mass transfer of dissolving and reacting (corrosive)species at high-flow or highly turbulent locations,although fluid shear stress on the oxide film at thematerial surface may also make substantial contribu-tions to the damage in some situations

FAC is a pervasive problem in most types ofsteam-raising system and has caused feedwater lineruptures, occasionally with fatal consequences, inboth fossil and nuclear plants.9,10In primary coolantsystems also, less serious (though costly) FAC occurschronically in the CS outlet feeders of conventionalCANDUs,11 and flow effects are implicated in thecorrosion of PWR pressure-vessel steel by boratedcoolant leaking through cracked penetrations in theRPV head.12FAC thus occurs in the regions of highturbulence in both single and two-phase flows, butnever in systems with dry steam

FAC depends on hydrodynamics (mainly steamquality, 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 criticalparameter combinations for the occurrence of FAC

in feedwater systems and the main parameter effectsare schematically summarized inFigure 3

The conditions leading to increased FAC ratesare usually related to regions with turbulent flow, tolow electrochemical corrosion potentials ECP (i.e.,

to chemically reducing conditions), and to low ironconcentrations in the water (Figures 3 and 4).Depending on the pH, the maximum FAC ratesoccur at about 130C in single-phase flow, and atabout 180C in two-phase flow (in the latter, it is thecondition in the liquid layer at the steel surface thatcontrols 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 cool-ants occurs at 300–310C at the core outlet, and FAC

is also significant in feedwater systems at the low

Figure 2 Pourbaix diagram for iron at 108m at 310C.

Reproduced from Beverskog, B.; Puigdomenech, I Corros.

Sci 1996; 38( 12 ): 2121–2135.

temperature of condensate extraction Specific

geo-metries 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

cool-ant systems, there is the desire to keep iron

concen-trations low to prevent crud build-up and radiation

transport problems, hence the frequent use of

high-alloy 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

con-ditions at about 140C, as may be found in BWR

feedtrains, can give high FAC rates, so dual-cyclePWRs or PHWRs routinely add an amine such asammonia to raise the pH in the secondary coolantcircuit The actual pH employed depends upon thematerials of construction; for all-ferrous feedtrains,

a pH25Cfrom 9.3 to 9.6 is usually specified, but thevalue is kept below 9.2 to avoid excessive corrosion ofcopper-base alloys, if these are present Also, toachieve a more even distribution of additive aroundthe circuit, an amine (such as morpholine) with acoefficient of distribution between the steam andliquid phases closer to unity than that of ammoniamay be used

Oxygen dissolved in the coolant is also a powerfulinhibitor of FAC; it has been added routinely to feed-water systems in BWRs and certain fossil boilers forsome time Depending upon the rate of attack, levels

of oxygen from a few ppb to several tens of ppb aresufficient to stifle FAC completely Maintaining adissolved oxygen content >30 ppb, which raisesthe corrosion potential ECP in the feedwater sys-tem above the Fe3O4/Fe2O3 phase boundary in thePourbaix diagram inFigure 2, is particularly crucial

in BWRs operating on hydrogen water chemistry

[Cr] in metal Log FAC rate ∼0.2% Cr

Critical conditions for high FAC risk

in feedwater Flow

Figure 3 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.

(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

recircula-tion piping and reactor internals (Recombinarecircula-tion 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-cooledreactors (AGRs) has been successfully mitigated sincethe 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 concentra-tion of 0.1% in the metal reduces FAC in 180Cammoniated water and water–steam mixtures atpH25C9 by about 70%.14Moreover, under CANDUprimary coolant conditions of 310C and pH25C10.5(adjusted with lithium), increasing the chromium con-tent of SA-106 Grade B CS from 0.019% to 0.33%reduces FAC by about 50%.11

pH = 9.04

10CrMo910 (A213 Gr T22–2.2% Cr, 1% Mo) 13CrMo44 (A213 Gr T12–1% Cr, 0.5% Mo) 15NiCuMoNb5

15Mo3 (A161 Gr T1–0.5% Mo)

St37.2 (A414 Gr B-carbon steel) Flow (kg h −1)

983

756 605

491 378 302 227 907

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

Fe3O4þ 3ð2  bÞHþþ H2¼ 3½FeðOHÞbð2bÞþ

with b ¼ 0, 1, 2, or 3

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

inter-face, and with the iron from the magnetite

dissolu-tion is transported to the bulk coolant at the rate Rm

The FAC rate RFACis thus twice the dissolution rate

Rdof the magnetite at the oxide–coolant interface

This concept of two processes in series – dissolution

Rdand 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

1 Steady-state assumption for the serial process:

2 Dissolution rate of magnetite at the oxide–coolant

interface according to eqn [VI] (assuming

first-order kinetics):

Rd¼ 0:5  dm=dt ¼ kdðCeq CÞ ½2

where kdis 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

3 Transport of Fe(II) species from the oxide–coolantinterface to the bulk environment by turbulentmass transfer:

where Cbis the concentration of Fe(II) species in thebulk coolant and h is the mass transfer coefficient,which is dependent on flow conditions and geometry.Fromeqns [1]–[3]it follows that:

no data to confirm this over the temperature ranges

of interest), whereas h only shows a moderateincrease through the temperature dependence of theproperties in eqn [6] If mass transfer controls, h issmall compared with kd (h kd) and the equationreverts to:

For a coolant of constant conditions containing little or

no dissolved iron (i.e., Cb  0), the driving force DCapproaches a constant value – the solubility of theoxide, Ceq – and RFAC varies as the mass transfercoefficient (which increases with increasing flow rateand turbulence) The mass-transfer model thenimplies that the effects of materials composition andcoolant chemistry on FAC rate are brought about bytheir effects on oxide solubility (Figure 5) According

toeqn [VI], the saturation concentration or solubilityCeqdepends on temperature, pH and H2concentration

by simple chemical equilibrium thermodynamics.Accordingly, the effect of chromium in the steelcan be attributed to the relative stability of mixedoxides containing chromium (iron chromite, FeCr2O4,for example, is virtually insoluble in reducing coolantand accounts for the protection afforded by stainlesssteel and similar alloys) As corrosion proceeds andthe magnetite dissolves, chromium is not leached out

in concert but continually concentrates in the film It

is interesting to note that the inhibition occurs diately at the start of exposure and continues at aboutthe same level, suggesting that the mechanism is therapid formation at the metal–oxide interface of a moreprotective layer of oxide that is maintained through-out exposure.16 It appears that the higher the chro-mium content of the steel, the more protective that

imme-oxide layer; hence, the immunity of stainless steel and

certain LAS like 13CrMo44 or 10CrMo910 These

observations cannot be fully explained with theDC

term in the simple mass-transfer model of eqn [5],

which would indicate an increasing inhibition with

time as the chromium concentrates in the film One

mode of inhibition of FAC by additives in the coolant

may be via incorporation in the magnetite to make it

less soluble From the experience with chromium,

whatever species is added should be available to affect

the metal–oxide interface consistently, presumably by

being kept permanently in solution in the coolant

So far, titanium has shown promise as a

coolant-borne inhibitor of FAC under CANDU primary

cool-ant conditions Its effectiveness has been attributed

to its ability to form a mixed oxide with iron –

ulvo¨spinel – along with the magnetite that forms

on corroding CS.17 An in-plant demonstration of

titanium addition to a CANDU primary system is

described inSection 5.06.2.3

The simple mass-transfer model also indicates

that temperature should affect FAC partly through

its influence on magnetite solubility In ammoniated

water at pH25C 9.0, there is a strong temperaturedependence and the maximum FAC rate occursbetween 130 and 140C, depending on flow rate.18The solubility of magnetite under the same condi-tions increases from the range 5–15 ppb at 25C to

a maximum of about 30 ppb that persists over therange 110–150C,9,19while mass transfer coefficients

at the same mass flow should approximately doublebetween 25 and 140C Similarly, in neutral water, themaximum attack for several materials occurs at about

150C,20while the magnetite solubility increases fromabout 70 ppb at 25C to a maximum of about 140 ppb

at 120–130C The rough correspondence betweenthe temperatures of maximum FAC rate and of maxi-mum magnetite solubility, as well as the effect oftemperature itself on solubility, indicate the stronginfluence of oxide film dissolution on the FAC mech-anism It is likely that at low temperatures dissolutionrates of magnetite are low enough for kdto have aneffect througheqn [4]and lower the flow dependenceaccordingly.18

The inhibiting effect of amines and high pH atfeedwater temperatures should also be realizedmainly through the solubility of magnetite Thus, inneutral water at 140C, the solubility of magnetite isabout 119 ppb, but if the pH25Cis raised with ammo-nia to 9.2, the value falls to the range 14–26 ppb.9,19This would suggest that a reduction in FAC rate by afactor of 8.5–4.5 might be expected from ammoniat-ing the coolant to pH25C9.2; however, experimentsindicate a reduction by a factor of only about 2.16

It is also instructive to consider the mass transferimplications of the model according toeqn [5] Masstransfer in pipe flow in aqueous systems can bedescribed via a correlation of the mass transfer coef-ficient h with dimensionless numbers:

Typically, experiments on mass transfer of solved species yield values between about 0.6 and0.9 for the exponent p.21,22Recent experiments in awater loop on FAC under neutral conditions at

dis-140C, however, indicated that the FAC rate RFACcorrelated rather weakly with Re.1.2,23An alternative

0.2 8.90 0.3 9.00

2.0 9.60

0.5 9.20 1.0 9.40 0.1 8.75

Figure 5 Solubility of magnetite/iron as a function of

temperature at various ammonia concentrations.

Reproduced from Dooley, R B Power Plant Chem 2008,

10( 2 ), 68–89.

mass transfer analysis gave an excellent correlation

with fluid shear stress at the pipe wall,t:

where P is a constant (see Figure 6) Thus, a steel

containing 0.019% chromium gave the correlation

RFACu ¼ 0.07t, while a steel containing 0.001%

chro-mium in parallel experiments gave RFACu ¼ 0.18t,

where RFACis in units of millimeters per year, u is

in meters per second, andt is in pascals.16

The predominance of mass transfer in developing

such correlations depends upon the dissolution rate

constant, kdineqn [4], being large enough to make

the mass transfer coefficient, h, controlling This

would seem to be valid under neutral chemistry

con-ditions, where the solubility of magnetite is high, but

under high-pH conditions, where the solubility is

reduced, kd may be reduced also and its influence

may become significant However, although recent

indications24 are that FAC in 140C ammoniatedwater at pH 9.2 is not correlated well by the simplemass-transfer model leading toeqn [6], those experi-ments also indicated a greater dependence on flowrate or shear stress, viz.,t raised to the power 1.5–2.0.This cannot be attributed to an increasing influence

of kdineqn [4]; apparently, a different mechanism isinvolved

Surface texturing usually accompanies FAC Insteam–water mixtures, ‘tiger-striping’ is caused bythe streaming pattern of the liquid film on the sur-face, while in single-phase water, ‘scalloping’ sculptsthe attacked surface with grooves, flutes, or shallowdepressions (Figure 7(a) and 7(b)) However, inexperiments in neutral water at 140C, in which cor-rosion rates of several millimeters per year were seen

in tubular test sections, a low-chromium steel oped no scallops, even though it corroded at more thantwice the rate of a higher chromium steel that devel-oped distinct scallop patterns.23 The scalloping thatwas seen was approximately related to the pipe flowvia a characteristic ‘scallop Reynolds number’:

in which the characteristic dimension is the averagescallop spacing While the scallops were formed by thecorrosion of the metal, it was significant that distinctoxide forms developed and were related more to scal-lop crests than to valleys Those forms, shown in

Figure 7(c), occurred over pearlite grains in themetal and may be described as ‘coral-like.’ They pro-vide further confirmation of the importance of oxidedissolution in the mechanism, since they are no doubtformed by the different solubilities of the differentcompositions of oxide overlaying the lamellae ofcementite and ferrite in the pearlite As the magnetite

(b)

3 μm

Diameter of piping (c)(a)

Figure 6 Correlation for flow-accelerated corrosion at

140C in neutral water: carbon steel with 0.019% Cr.

generally dissolves, that over the cementite lamellae is

less soluble and is left standing proud It was noted in

the experiments that the ‘coral’ oxide concentrated

about 50% more chromium than the surrounding

oxide on the ferrite grains, possibly because the

under-lying pearlite contained more chromium as carbide

associated with the cementite

Loop studies using tubular test sections of the

material of interest16 under reactor feedwater

condi-tions establish the basis for adding oxygen with

mini-mal residual concentrations left at the end of piping

systems At 140C in neutral water, about 40 ppb of

dissolved oxygen are required to stifle FAC, whereas at

pH25C9.2 with ammonia, only 1–2 ppb are required

The concentrations required for stifling are related to

the measured FAC rates and it is clear that – as oxygen

is progressively added to the corroding system – the

cathodic reaction of water being reduced to hydrogen

is replaced by oxygen reduction; at the stifling

concen-tration, the oxygen sink disappears and with continuing

addition its concentration in the loop jumps sharply

However, although there is an obvious relationship

between the FAC rate at stifling and the stoichiometric

flux of oxygen by mass transfer to the surface, a

straightforward linear correspondence may not apply.13

While several mechanistic models of FAC in

feed-water systems based mostly on the principles behind

eqn [4] have been developed, empirical models have

been applied extensively for some time In the 1980s, for

example, parametric studies at the laboratories of the

then Siemens-KWU led to the formulation of a

corre-lation between pipe wall thinningDd and the system

variables u (flow velocity), T (temperature), pH, O2

(oxygen concentration), M (materials composition –

Cr, Mo, and Cu), and t (exposure time):

Dd ¼ kcf ½u; T ; pH; O2; M; t ½9

where kc is a geometry factor The correlation

was developed initially from data for single-phase

water flow, but was adapted to two-phase steam-water

flows, with the bulk velocity u substituted by the mean

velocity of the annular film of water covering the pipe

wall The resulting computer code, ‘WATHEC,’ was

restricted to steels with the content of Cr plus Mo less

than 5% and exposure times greater than 200 h The

predictions of wall thinning for a large number of

situations were equal to or greater than the measured

values in 85% of the cases – in other words, the code

was considered to be suitably conservative.25Later, the

data management tool ‘DASY’ was added to the code

The EPRI-sponsored computer code

‘CHEC-WORKS™’ combined an empirical equation, which

had some basis in mechanisms such as that leading to

eqn [4], with a comprehensive data managementscheme.26 The data management includes analysis

of ultrasonic test data, calculation of critical wallthickness for components at risk, and organization

of pertinent databases The FAC rate RFACis written

as a function of the system variables:

RFAC ¼ f ½T; AC; MT; O2; pH; G; a ½10where T is temperature; AC is alloy content of Cr, Mo,and Cu; MT is mass transfer; O2 is concentration ofdissolved oxygen; G is a geometry factor; and a is thesteam void fraction The factors ineqn [10]are interre-lated and the equation is nonlinear While the absolutepredictions of RFACin CHECWORKS™ are not gen-erally of high precision, iterations incorporating plantmeasurements can identify the locations of risk andcan rank components in the order of vulnerability.27The FAC of CS is most pronounced under feed-water conditions, but it also occurs at higher tem-peratures in the primary coolant systems of PHWRs.The phenomenon was identified in the late 1990s atthe Point Lepreau CANDU in New Brunswick,Canada, where surfaces of affected outlet feeders of

CS were scalloped and the wall thinning rates plottedagainst coolant velocity indicated a dependence onthe velocity raised to the power 1.5.28

Regions ofhigh turbulence, such as the tight-radius bends close

to the reactor face, were more severely affected

It was also noted that the coolant at the core outletwas unsaturated in dissolved iron, since it entered thecore at 265C saturated after its passage through thesteam generators of nickel alloy and the inlet feeders

of CS; as its temperature rose in the fuel channelsthe solubility at the high pH rose in concert (theCANDU core contains no iron-bearing alloys, so itcannot act as a source of dissolved iron)

Although the turbulence (and therefore mass-transfer) regions are again the most affected inprimary coolant FAC, it is unlikely that the mechanism

high-in primary coolant is straightforward mass-transfercontrol based on eqn [11] First of all, the velocitydependence is too high (the power of 1.5 rather than0.6–0.9 as expected from correlations such aseqn [6]).Second, measurements of the dissolution rate of mag-netite under chemistry conditions close to those ofCANDU coolant29have given a value of kdineqn [4]

very much lower than the mass transfer coefficient h,which would put the mechanism squarely under dis-solution control with no velocity effect at all Thealternative theory proposed is that dissolution of mag-netite works in synergy with fluid shear stress at the

surface to degrade the oxide Thus, the loosening of

the magnetite crystallites in the film makes them

sus-ceptible to removal by the fluid forces and as they are

eroded away the film becomes less protective A

mathe-matical model developed on this principle was able to

predict quite well the thinning of the walls of outlet

feeders at an operating plant in terms of the

develop-ment of the oxide film, the pattern of attack around

representative bends, and the corrosion potential ECP

and velocity-dependence of FAC rate in individual

feeders.30The model was adapted for predicting

corro-sion under conditions when the coolant is saturated in

dissolved iron and gave reasonable predictions of oxide

film growth and general corrosion in CANDU inlet

feeders, where corrosion rates are quite low

It is probably more than a coincidence that FAC

under these primary coolant conditions, when

magne-tite solubility is low, seems not to be controlled directly

by mass transfer, while similar indications apply under

feedwater conditions at high pH, when solubilities

are also low The parallel between the two situations

could be clarified if measurements of kdunder

feed-water conditions were available and the measurements

under primary coolant conditions were verified

The high rate of general corrosion of CS caused

by aerated concentrated solutions of boric acid

origi-nating from leaking PWR coolant was described in

Section 5.06.2.1 Some of the studies that quantified

the attack were done with dynamic systems, such as

evaporating sprays, and it became clear that flow has

an effect.2Of immediate concern is the corrosion of

RPV steel caused by borated coolant leaking through

cracked penetrations housing control rod drive

mechanisms At the Davis Besse PWR in 2002, such

corrosion had threatened the integrity of the vessel

The sequence of events that can lead to cavity

forma-tion next to a nozzle was postulated12to be in three

phases: initially, slow seepage of coolant into the

exter-nal annulus (crevice) in the head would be

accompa-nied by low corrosion rates; next, when the crevice had

opened enough and the crack had lengthened to give

substantial leak rates, an evaporating coolant jet would

accelerate the attack through flow effects; finally,

leak-age into a cavity would create a turbulent evaporating

pool, extending the attack sideways

An extensive testing program sponsored by the

Electric Power Research Institute (EPRI), Palo Alto,

California, investigated the phases of boric acid

attack at Davis Besse The second phase, which

expe-rienced substantial flow effects, was simulated with

laboratory experiments in which a flashing jet of

borated coolant was directed onto a heated sample

of pressure-vessel steel and the damage assessed

in terms of system parameters – notably, coolantchemistry and flow rate.31 Volumetric (or massive)metal loss was correlated with volumetric coolantflow and seemed to behave differently from metalpenetration, which was correlated with jet velocity.FAC was in evidence through miniature scallops inthe damage craters that formed around (but somedistance away from) the points of jet impact Metalloss rates attained about 3 cm3year1at a flow rate of

200 ml min1with a boric acid concentration lent to 1500 ppm [B] and pH300Cof 6.9 adjusted withlithium; the rate depended on the volumetric flow

equiva-in the jet raised to the power 0.84 Under thesame chemistry conditions, the penetration ratereached 200 mm year1at a jet velocity of 140 m s1and the two were correlated via the velocity raised tothe power 4.3 It was notable that neither pH300Cnorthe boron concentration was the controlling chemis-try parameter; rather, it was the ratio [B]/[Li].5.06.2.2.3 Service experience and mitigatingactions

Many incidences of feedwater pipe thinning by FACfrom two-phase coolant were reported in the 1980s

In 1985 March, a line downstream of a level controlvalve for a feedwater heater at the Haddam NeckPWR actually ruptured because of FAC induced byflow-impingement However, the first major incident

in a nuclear plant was the catastrophic pipe break atthe Surry Unit 2 PWR in December 1986, which led

to five deaths and several injuries The 0.46 m ter line thinned and ruptured at an elbow, 0.3 m from

diame-a 0.6 m hediame-ader, diame-as diame-a section of the pipe wdiame-all 0.6was blown out Until then, FAC by steam–water mix-tures had been considered to be more serious thanFAC by single-phase coolant Six months later, exces-sive thinning of a feedwater line was reported at theTrojan plant and, in September 1988, Surry Unit 2reported 20% wall loss in the suction line to a feed-water pump over a 1.2-year period

Reports of serious thinning of feedwater pipingcontinued after the Surry incident, even thoughplant inspections had generally become more rigorousand chemistry control had tightened In May 1990,the Loviisa Unit 1 WWER (Eastern type PWR) inFinland suffered a break in a 0.3 m diameter line inthe turbine hall, releasing about 50 m3of steam andwater into the building, and in February 1993, asimilar incident occurred in Unit 2

The latest major FAC incident in a nuclear plantwas the rupture of a feedwater line at the Mihama

Unit 3 PWR in 2004, which led to four deaths and

seven serious injuries.32The thinning of the pipe wall

from 10 to 0.6 mm by FAC caused a large section to

peel back after rupture, allowing the coolant at

140C to flash to steam on release The damagedpipe is shown in Figure 8 Maximum thinning rateswere observed just downstream of an orifice plate,where turbulence intensity was high It should benoted that chemistry had been maintained at high

pH with ethanolamine and that hydrazine was used toscavenge oxygen However, the location had not beeninspected since the plant start-up in 1976 The possi-bility of adding oxygen to the feedwater is beingconsidered and inspection procedures have beenrevised extensively

In 1997, at the Point Lepreau CANDU PHWR

in New Brunswick, Canada, the outlet feeder pipes

of CS that carry the heavy water coolant from thecore to the steam generators were found to be cor-roding excessively The same problem has since thenbeen seen at all CANDUs in operation before 2000

Figure 9shows the arrangement of pipes at a reactorface The feeders are about 76 mm diameter and carry

Figure 8 The ruptured feedwater line at Mihama-3.

10

4 3

9

1 Reactor outlet header

3 Reactor outlet header

2 Reactor inlet header

4 Reactor inlet header

5 Feeder tube upper supports

6 Calandria end shield face

coolant at velocities between about 12 and 22 m s1.

At the coolant temperature of 310C and pH 10.6

(adjusted with lithium), the solubility of magnetite is

relatively low – about 1.7 ppb – and the FAC rates

accordingly attain only about 120mm year1 The

attack is only a fraction of that observed under

feedwater conditions and has caused no safety issues,

but it means that feeder integrity may be

life-limiting and has necessitated replacements at some

reactors Mitigating actions have been taken by

reducing the primary coolant alkalinity to the

bot-tom of the recommended pH range, where the

mag-netite solubility is close to the minimum, while

plants in service since 2000 employ feeders with

a relatively high chromium content (0.3% in

contrast with the 0.02% in earlier reactors)

One trial of a coolant additive was made at the

Darlington Unit 3 900 MW CANDU in 2002,

when a titanium dioxide slurry was added to one

channel to give a concentration of ten or so

micro-gram per kilomicro-gram at the outlet feeder33; significant

reductions in FAC rate were recorded, but it

was decided not to undertake the further

develop-ment needed to proceed to the next stage of

full-plant addition Meanwhile, feeder replacement

has become a manageable – if costly – operation at

severely affected plants

The boric-acid corrosion at the Davis Besse PWR

in Ohio, which in 2002 was found to have a large

wastage cavity in the RPV head adjacent to a

pene-tration housing a control rod drive mechanism, was

postulated to be partly due to FAC (seeFigure 10)

About 2040 cm3 of the pressure-vessel steel had

corroded away and about 106 cm2 of stainless steel

cladding were exposed at the bottom Substantial

quantities of solid boric acid had deposited close by

Subsequent investigations12determined that coolant

had leaked from a crack in the adjacent Alloy 600

nozzle into the surrounding annulus and in time

had widened it The crack was an example of primary

water stress corrosion cracking (PWSCC), of which

numerous incidences have been recorded in PWRs

No simple means of mitigation have been proposed

for existing plants, since the coolant boron level is

fixed by reactivity considerations Long-term

pre-vention entails avoiding operating with coolant

leaks (as is, in fact, the regulation in some

jurisdic-tions), for example, through minimizing the

possibil-ity of PWSCC by using less-susceptible Alloy 690

material for the CRDM penetrations In the

mean-time, more rigorous inspection regimes are being

implemented

Environmentally Assisted Cracking5.06.3.1 Pitting

In CS & LAS, a shallow form of pitting can occur inthe complete absence of anionic water impurities asthe electrochemical corrosion potential (ECP) at thesteel surface is raised, for example, through oxygenand other oxidizing species Such corrosion pits often,but not exclusively, initiate at MnS inclusions whichintersect the steel surface.Figure 11, originally com-piled by Hickling, shows the critical boundariesbetween uniform surface attack and shallow pitting

in high-purity water at low flow rates as a function oftemperature and dissolved oxygen level The criticaloxygen concentration for pitting drops with decreas-ing temperature and is further reduced by a simulta-neous mechanical straining of the surface, or byincreased sulfate and chloride impurity levels Fur-thermore, pitting in CS & LAS is favored by highsteel sulfur contents and quasi-stagnant flow condi-tions In the absence of impurities, increasing the flowrate of water across the metal surface mitigates theaforementioned form of pitting corrosion Alkalization

Figure 10 Cavity in the reactor pressure vessel head at the Davis Besse pressurized water reactor.

shifts this boundary to much higher values, as does the

introduction of buffering and passivating species (e.g.,

on the secondary side of steam generators).34–37

Some degree of pitting corrosion is inevitable

after long-term exposure of unclad CS & LAS

sur-faces to water in LWR systems and is not usually a

threat to either coolant purity or to structural

integ-rity Shallow pitting has been observed primarily in

specific piping systems with residual water because of

incomplete draining during nonoperational periods(shutdown corrosion) This can be avoided by ade-quate wet or dry preservation techniques If pittinghappens during normal plant operation at high tem-peratures, however, it indicates conditions underwhich EAC may also occur (since this is controlled

by similar parameters, seeSection 5.06.3.2) and caneven be directly implicated in the initiation of EAC(Figure 12)

Dissolved oxygen content DO (ppb)

lzumiya, Tanno Videm

Mizuno et al.

Ford

General surface attack

Coupon specimens

Pitting

Pitting without straining

General surface attack pitting

Pitting with straining

Figure 11 Boundaries between uniform corrosion and pitting attack in carbon and low-alloy steel in quasi-stagnant high-temperature water Compiled from The general and localized corrosion of carbon and low-alloy steels in

oxygenated high-temperature water EPRI-NP-2853; EPRI: Palo Alto, CA, 1983; http://my.epri.com/ ; Electric Power Research Institute BWR environmental cracking margins for carbon steel piping EPRI Report NP-2406; EPRI: Palo Alto, CA, 1982; http://my.epri.com/; Indig, M et al Rev Coat Corros 1982, 5, 173–225; Videm, K In Proceedings of the 7th

Scandinavian Corrosion Congress, Trondheim, Norway, May 26–28, 1975; pp 444–456.

20 mm

Acc.V 20.0 kV 5.9 1696x SE 11.0 5/1;250C;HigN Spot Magn Det WD 20 mm

Figure 12 Strain-induced corrosion cracks initiating from a corrosion pit or a (dissolved) MnS inclusion at the surface of a low-alloy and carbon steel in high-temperature water in slow strain rate experiments Adapted from Congleton, J et al Corros Sci 1985, 25, 633–650; Atkinson, J D et al In Proceedings of the 13th International Conference on Environmental Degradation of Materials in Nuclear Power Systems, Whistler, British Columbia, Canada, Aug 19–23; King, P., Allen, T., Busby, J., Eds.; The Canadian Nuclear Society: Toronto, Canada, 2007; CD-ROM.

5.06.3.2 Environmentally Assisted

Cracking

5.06.3.2.1 Basic types of EAC and major

factors of influence

EAC is used here as a general term to cover the

full spectrum of corrosion cracking ranging from

stress corrosion cracking (SCC) under static load to

corrosion fatigue (CF) under cyclic loading

condi-tions (Table 2).38,39Strain-induced corrosion

crack-ing (SICC) involvcrack-ing slow, dynamic straincrack-ing with

localized plastic deformation of material, but where

obvious cyclic loading is either absent, or is restricted

to a limited number of infrequent events such plant

startup and shutdown, is increasingly used as an

appropriate term to describe the area of overlap

between SCC and CF.38,39

Under critical parameter combinations, EAC is

observed in all wrought and welded CS & LAS in

high-temperature water The EAC crack path is

usually perpendicular to the direction of maximum

tensile stress and transgranular in nature, with aquasicleavage appearance showing a feathery mor-phology at high magnifications The general fractureappearance is similar for SCC, SICC, and even CF(at least for strong environmental acceleration offatigue crack growth), thus confirming that EAC isgoverned by the same basic process for all threeloading modes In the case of cyclic loading at fre-quencies103Hz, the fracture surface also usu-ally contains both ductile and brittle fatigue striations,which are perpendicular to the local crack-growthdirection.38

EAC initiation and growth in CS & LAS aregoverned by a complex interaction of environmental,material, and loading parameters, and most influen-cing factors are both interrelated and synergistic Themajor parameters of influence, which have beenidentified so far, are summarized in Table 3.38,39The effect of these parameters on EAC initiationand crack growth (including key thresholds) is

Table 2 Basic types of EAC in CS & LAS and relevant nuclear codes

Mechanism Environmentally assisted cracking (EAC)

condition

Thermal fatigue, thermal stratification,

Start-up/shut-down, thermal stratification

Transient-free, steady-state power operation

Characterization of

crack growth

ASME XI Code Case N-643 (PWR)

High-sulfur line of F & A model as upper bound

BWRVIP-60 disposition lines Characterization of

Hickling, J et al PowerPlant Chem 2005, 7, 31–42.

Table 3 Major influencing factors for EAC in C & LAS

Environmental parameters Material parameters Loading parameters Corrosion potential, dissolved oxygen

Cl, SO 4  , S2, HS Susceptibility to dynamic strain ageing,

Concentration of interstitial C and N

Type of loading Flow rate Hardness/yield stress if >350HV5/800 MPa Residual stress

Source: Seifert, H P.; Ritter, S Research and service experience with environmentally-assisted cracking of carbon & low-alloy steels in temperature water SKI-Report 2005:60; SKI: Stockholm, Sweden, 2005; ISSN 1104-1374 http://www.stralsakerhetsmyndigheten.se/ Hickling, J et al PowerPlant Chem 2005, 7, 31–42.

high-discussed in detail in Seifert and Ritter38 and an

interpretation of their synergism is given both there

and inSection 5.06.3.2.4

5.06.3.2.2 Corrosion fatigue and

strain-induced corrosion cracking

5.06.3.2.2.1 Initiation and susceptibility

conditions

Slow strain rate (SSR)38,40 and low-cycle fatigue

(LCF) tests40–42with smooth specimens have clearly

shown that CF and SICC can occur in CS & LAS in

oxygenated, high-purity, high-temperature water if

the following conjoint threshold conditions are

simultaneously satisfied:

Water temperature >150C In LCF

experi-ments, susceptibility then increases with

tempera-ture up to320C SSR tests, on the other hand,

usually indicate a maximum of susceptibility

between 200 and 270C, depending on strain rate

Corrosion potential > ECPcrit¼ 200 mVSHEor

dissolved oxygen content >30 ppb Above this

threshold, EAC susceptibility then generally

increases with increasing ECP/oxygen content,

but saturates at very high levels

Loading which leads to (local) macroscopic strains

at the water-wetted surface above the elastic limit

The susceptibility then increases strongly with the

degree of plastic strain SSR experiments with

tapered specimens, and LCF tests with different

waveforms, indicate a minimum critical strain of

0.15–0.2%, which is in a similar range to typical

oxide film rupture strains on CS & LAS (0.05–

0.2%) in high-temperature water

Positive strain rates below 103s1 The EAC

susceptibility then increases with decreasing strain

rate de/dt In most LCF investigations, saturation

of the decrease in fatigue life is observed below

a strain rate of 105s1, but SSR tests indicate a

maximum of susceptibility between 105 and

107s1, depending on ECP and temperature

EAC susceptibility increases with increasing steel

sulfur content and a lower threshold is often quoted

at0.003 wt%, but experimental evidence for the

latter is weak Distinct material susceptibility

to dynamic strain aging (DSA) in the critical

temperature/strain-rate range, or a low yield

stress, may also favor crack initiation by SICC

If one or other of these conjoint threshold conditions

is not satisfied, SICC initiation is extremely unlikely

and no, or only a minor, environmental reduction of

fatigue life is observed in high-temperature water.Furthermore, in high-purity water, a high flow ratemay completely suppress SICC susceptibility and sig-nificantly retard CF crack initiation (in particular,for small strain amplitudes or slow strain rates) com-pared to quasi-stagnant conditions, since the risk forthe formation of an aggressive, occluded water chem-istry within small surface defects is significantlyreduced by convection Note, however, that high levels

of chloride or sulfate may extend the range of tibility to less severe conditions (e.g., to lower ECP andstrain levels)

suscep-The range of system conditions where EAC crackgrowth from incipient cracks may occur is sig-nificantly extended compared to the initiation sus-ceptibility conditions specified earlier For example,

CF crack growth has been observed in high-purityPWR water at ECPs below500 mVSHEunder cer-tain cyclic loading conditions (102to 10 Hz).38Apart from local stress raisers such as weldingdefects, which may help overcome the strain threshold

in the field, the effect of initial surface condition face roughness, cold work, residual stress, oxide film,and preoxidation) on SICC and CF initiation is muchless pronounced than for (high-cycle) fatigue in air, orwith SCC of stainless steels or Ni-base alloys SICCcracks usually, but not exclusively, initiate at MnSinclusions or corrosion pits.38,40,43,44Pitting, particu-larly if occurring actively, therefore facilitates SICCinitiation (Figure 12) CF cracks, on the other hand,initiate mainly along slip bands, carbide particles, or atthe ferrite–pearlite phase boundary, and less frequently

(sur-at micropits or MnS inclusions.40–42 The effect ofpitting and MnS inclusions on CF initiation is thusmoderate, but may become more pronounced in thecase of deep, high-aspect-ratio pits, mild environmen-tal conditions, or at small strain amplitudes.40

5.06.3.2.2.2 SICC initiation and crack growth fromincipient cracks

In high-purity water in the absence of any significantfatigue contribution, CS & LAS show distinct SICCsusceptibility only in highly oxidizing environments.For example, it is almost impossible to initiate relevantSICC crack growth in precracked specimens in slowrising-load tests with constant load rate at KIvalues

<70 MPa m1/2

in high-purity water at an ECP of

<100 mVSHE Even under highly oxidizing tions (ECP þ50 mVSHE), KIvalues of25 MPa m1/2have to be exceeded to initiate SICC in slow, rising-load experiments in high-purity water A maximum

condi-in SICC condi-initiation susceptibility (i.e., a mcondi-inimum condi-in

of the decrease in fatigue life is observed below

a strain rate of 10 5< /sup>s1, but SSR tests indicate a

maximum of susceptibility between 10 5< /sup> and. .. that high levels

of chloride or sulfate may extend the range of tibility to less severe conditions (e.g., to lower ECP andstrain levels)

suscep-The range of system conditions where...

107s1, depending on ECP and temperature

EAC susceptibility increases with increasing steel

sulfur content and a lower threshold is often quoted

at0.003 wt%, but