Materials Selection and Design (2010) Part 12 doc

The subject of designing for high-temperature service is outside the scope of this article see the article "Design for High-Temperature Applications" in this Volume and Ref 58, 73, and

Table 4 Reference electrodes for use in anodic protection

Calomel Sulfuric acid

Silver-silver chloride Sulfuric acid, Kraft solutions, fertilizer solutions

Mo-MoO 3 Sodium carbonate solutions

Bismuth Ammonium hydroxide solutions

Type 316 stainless steel Fertilizer solutions, oleum

Hg-HgSO 4 Sulfuric acid, hydroxylamine sulfate

Pt-PtO Sulfuric acid

Fig 17 Potential control in the passive region for anodic protection Source: Ref 5

Localized Attack

Dissimilar-Metal Corrosion. If two metals, say aluminum and copper, are immersed in aerated chloride solutions, they will individually attain corrosion potentials of about -1.0 and -0.25 V (vs SCE), respectively, as indicated in the galvanic series (Fig 13) If these metals are electrically connected, then electrons will flow through the connection from the aluminum to the copper, and the aluminum will become anodically polarized and copper will be cathodically polarized with respect to their original (unconnected) corrosion potentials The extent of this polarization and the resultant corrosion rates on the two metals will be dependent on the relative areas of the two electrodes and the conductivity of the solution For instance, referring to the schematic Evans polarization diagrams for two metals M and N (Fig 18), it is seen

that their corrosion potentials change from the "unconnected" values EcorrM and EcorrN toward a common value Ecorr-couple, which is determined by the criterion that the total reduction current is balanced by the total oxidation current at the

various parts of both surfaces The result is that the corrosion current on the more active metal is increased from IcorrM to

IcorrMcouple The increase depends on the relative areas of the two metals; this important aspect can be illustrated by considering the aluminum/copper couple in aerated chloride solutions (Fig 19) If it is assumed that the area of the two metals are equal (say 1 cm2), then in the uncoupled condition the corrosion rate on the aluminum is about 0.7 A/cm2 (0.3 mil/year), but when connected to the copper, the corrosion potential will increase from -1.05 V (vs SCE) to approximately -0.8 V (vs SCE), and the corrosion rate on the aluminum will increase dramatically to 70 A/cm2 (30 mils/year) However, if the area of the copper is increased by a factor of ten, then the oxygen reduction current on the copper will increase by a factor of ten, and consequently, the current density on the original 1 cm2 of aluminum will be increased to a catastrophic 700 A/cm2 (300 mils/year)

Fig 18 Schematic potential/current curves for metals M and N, illustrating the fact that the corrosion current

on M (icorrM) increased to icorrMcouple when connected to metal N This current is determined by the criterion that the total oxidation and reduction currents be equal

Fig 19 Prediction of galvanic corrosion rates of aluminum/copper couples and effect of aluminum/copper

surface area ratio Source: Ref 5

The other important factor in this phenomenon is the ionic and electronic resistivity of the overall circuit For instance, as

the ionic resistance increases, the resultant "iR" drop in the solution will eventually cause the corrosion potentials to

revert to the uncoupled values However, it is incorrect to assume that all areas of a given metal surface are at the same

corrosion potential (which is created by the local metal-environment interface) because of local polarization (e.g., areas of

the aluminum adjacent to the copper) or chemistry variations (e.g., oxygen consumption in crevices and cracks) Thus, if the metals are physically in contact, the area immediately adjacent to the joint will be polarized, causing the attack at dissimilar-metal joints to be localized near the joint itself

The obvious design factors in systems involving dissimilar metals are to: (a) select metals with the least difference in

"uncoupled" corrosion potentials, or (b) minimize the area of the more noble metal with respect to that of the active metal,

or (c) insert an electrically insulating gasket between the two dissimilar metals (Fig 20)

Fig 20 Flange insulated to eliminate a galvanic couple Source: Ref 10

Intergranular Attack. Metals discussed so far in this article have been regarded as homogeneous, but in fact, most engineering alloys are multiphased or contain distributions of solid-solution elements that have different chemical activities Such metallurgical inhomogeneities occur especially at grain boundaries, either due to grain-boundary segregation or intermetallic precipitation, and these can give rise to localized intergranular corrosion

Intergranular corrosion of austenitic Fe-Cr-Ni stainless steels offers an ideal example of the phenomenon and design methods to counteract it In this system, chromium carbide can precipitate at the grain boundary by a classical thermally

activated nucleation and growth process during heat treatment or welding This carbide precipitation, called sensitization,

occurs at about 510 to 790 °C (950 to 1450 °F) and is accompanied by chromium depletion in the adjacent metal matrix such that the chromium content can fall from 18% to less than 10% Cr in a band up to 10 m from the grain boundary This depleted zone will have markedly different corrosion properties from the adjacent high-chromium matrix For instance, it is seen (Fig 21) that if the chromium content falls much below 12%, then the corrosion rate in acidic solutions rises markedly in a given oxidizing potential range, and preferential corrosion will occur This behavior will be aggravated by the fact that this narrow depleted zone will have a lower corrosion potential than the larger area of passivated high-chromium alloy connected to it, and hence further corrosion will occur due to galvanic effects, as explained in the previous section, "Dissimilar-Metal Corrosion."

Fig 21 Effect of chromium content on anodic polarization of Fe-Ni alloys of 8.3 to 9.8% Ni in 2N H2 SO 4 at 90

°C Source: Ref 106

This type of attack depends on two conjoint conditions, and the problem can be resolved by attention to only one of these conditions First, grain-boundary chromium-carbide precipitates and their associated chromium-depleted zone are necessary These conditions can be avoided by attending to such factors as the temperature/time requirements for such

"sensitization," by annealing the structure to solutionize the carbide, and by making alloy compositional changes such as low carbon content (e.g., L-grade stainless steels with <0.03% carbon) or the addition of elements such as niobium or titanium, which form more stable carbides than chromium carbide Second, the intergranular corrosion will be observed only under oxidizing potential conditions; thus, avoidance of oxidizers such as oxygen, Fe3+ or Cu2+ (e.g., Eq 5a, 5b, 6, 7) may well alleviate the problem

Similar intergranular attack phenomena are seen in other passive systems, where the localized attack is associated with either active depleted zones (e.g., the copper-depleted zones in Al-Cu or Al-Zn-Mg-Cu alloys or the molybdenum-depleted zones in Ni-Cr-Mo alloys) or with active precipitates (e.g., Mg2Al3 in the Al-Mg alloys or MgZn2 in Al-Zn-Mg

alloys) In all instances, however, mitigation can be obtained by attention to the metallurgical conditions that give rise to the precipitation or to the specific environments in which galvanic attack occurs

Dealloying corrosion is associated with alloys whose constituents are elements having very different electrochemical activities For instance, zinc-copper, gray cast iron, and aluminum-tin alloys give rise to phenomena classified as

"dezincification," "graphitic corrosion," and "dealuminification," respectively In two-phase structures such as / brasses and graphitic iron, the mechanism is partially galvanic (or dissimilar-metal) attack, but it is believed that in some systems, the mechanism is enhanced by combined dissolution of both elemental components and then reprecipitation of the more noble element The resultant damage is a friable surface with large amounts of porosity Given these component parts to the mechanism, it is apparent that potential changes can either exacerbate or mitigate the problem

Pitting, Crevice Corrosion, and Differential Aeration. Pitting and crevice corrosion arise from the creation of a localized aggressive environment that breaks down the normally corrosion-resistant passivated surface of the metal This localized environment normally contains halide anions (e.g., chlorides) and is generally created because of differential aeration, which creates potential drops between most of the surface and occluded regions (e.g., pits, crevices, and inclusions) that concentrate the halide at discrete locations

In pitting, this localization may begin at microscopic heterogeneities such as scratches and inclusions (e.g., sulfides) Above a given potential, negatively charged anions (e.g., Cl-) accumulate on the metal surface and can cause breakdown

of the protective oxide The breakdown mechanism continues to be a topic of research (Ref 28) Catastrophic localized

breakdown occurs at a specific corrosion potential, Epit, that is a function of the material, chloride concentration, pH, and temperature (Fig 22) Once this breakdown occurs, pit propagation can progress rapidly, because:

• The environment within the pit is deaerated (i.e., at low potential), thereby setting up a potential drop between the pit and the higher-potential surface surrounding the pit This potential drop concentrates the aggressive anion in the pit

• Electroneutrality considerations dictate that the increase in negatively charged non-OH- anions be counterbalanced by an increase in cations, and these usually are hydrogen ions (i.e., the pH decreases)

• The combination of the two factors above leads to an increased metal dissolution rate within the pit

• The high concentration of metal cations produced in the pit leads to the precipitation of, for example, metal hydroxides near the mouth of the pit, which helps ensure that the severe localized environment is contained (Fig 23)

• The precipitation of noble ions on the metal surface, or within the alloy as second phases, further accelerates the corrosion rate

Fig 22 Schematic representation of critical pitting potential, Epit , due to anodic polarization Source: Ref 5

Fig 23 Schematic representation of processes occurring in an actively growing pit in iron Source: Ref 5

A very similar sequence of events occurs in crevice corrosion, the main difference being that the initiation event is associated with the creation of a localized aggressive environment in a macroscopic (often a designed-in, geometric) crevice As with pitting, however, the mechanism of this localization is associated with deaeration (low potential) inside the crevice, coupled with an aerated (high-potential) environment outside

Environmentally Assisted Cracking

Environmentally assisted cracking describes the initiation and subcritical crack propagation in structural metals due to the

combined action of tensile stress, material microstructure, and environment The conditions of these three parameters may

be specific to a given alloy/environment system, and if one of these conditions is not met, the problem does not occur

The topic is made more difficult, however, by the superposition of various mechanisms (slip dissolution, hydrogen embrittlement) and phenomena (stress corrosion, corrosion fatigue, hydrogen embrittlement) Part of this problem is addressed in the section "Life Prediction and Management" in this article The rest of this section addresses the understanding of the mechanisms of environmentally assisted cracking and how this can be used qualitatively to control cracking

Candidate Crack-Propagation Models. The basic premise for all of the proposed crack-propagation mechanisms for ductile alloys in aqueous solutions is that the crack tip must propagate faster than the corrosion rate on the unstrained crack sides If this were not true, the crack would degrade into a blunt notch (Ref 29, 30) Indeed, the suppression of both stress corrosion and corrosion fatigue in many systems can be explained in terms of blunting of cracks during the early propagation stage For instance, low-alloy steels will not exhibit stress corrosion in acidic or concentrated chloride solutions unless the general corrosion/blunting effect is counteracted with chromium or nickel alloying additions (Ref 31,

32) Similar blunting explanations can be proposed for the case of corrosion-fatigue crack initiation of aluminum in chlorides in comparison to hydroxides (Ref 33)

Numerous crack propagation mechanisms were proposed in the period 1965-1979 (Ref 34, 35, 36, 37, 38, 39, 40, 41, 42,

43, 44) With the advent in recent years of more sensitive analytical capabilities, however, many of the earlier cracking hypotheses have been shown to be untenable The candidate mechanisms for environmentally assisted crack propagation (for both stress corrosion and corrosion fatigue) have been narrowed down to slip dissolution, film-induced cleavage, and hydrogen embrittlement

Qualitative Prediction Methods for Environmentally Assisted Cracking in Ductile Alloy/Aqueous Environments. Qualitative predictions of cracking have centered around the observation that the rate-determining step

in all of the above cracking mechanisms is not necessarily the atom-atom rupture process itself, but is one (or a combination) of mass transport of species to and from the crack tip, passivation reactions, and the dynamic strain processes at the crack tip (Ref 29, 30) Thus, changes in cracking susceptibility for most ductile alloy/aqueous environment systems with, for instance, changes in temperature, electrode potential, stressing mode, or environmental composition, can be explained logically (Ref 29, 30) using a reaction-rate surface (Fig 24), regardless of the specific atom-atom rupture mechanism at the crack tip This fact can be reinterpreted in terms of the crack propagation rate/stress intensity (Fig 25) (Ref 29) relationship for a given alloy/environment system subjected to different loading histories in which the limiting and rate-controlling reactions can be defined (Ref 29)

Fig 24 Schematic reaction-rate surface illustrating the variation in crack-propagation rate with the

rate-controlling parameters in the slip-dissolution, film-induced cleavage and hydrogen-embrittlement mechanisms for environmentally assisted cracking in ductile alloy/aqueous environment systems Source: Ref 29, 30

Fig 25 Suggested variation (Ref 29) in environmentally controlled crack-propagation rate with stress intensity

for various crack-tip deformation rates COD Note the suggested rate-controlling parameters and the fact that

these relationships should be bounded by a maximum crack propagation and a minimum theoretical KIscc or KTH

The importance of passivation kinetics on crack propagation is well recognized (Ref 29, 45, 46, 47, 48, 49) Very slow passivation rates at the crack tip will promote crack blunting due to excessive dissolution on the crack sides, whereas very fast rates will minimize the amount of crack-tip penetration per oxide-rupture event Maximum susceptibility, with high-aspect-ratio cracks, will occur at intermediate passivation rates or in regimes of barely stable passivity (e.g., near the passivation potential) The effects of potential, anion (or cation) content, and alloying addition on cracking susceptibility can be quantitatively understood by this simple concept, regardless of whether the advancement mechanism is slip dissolution or hydrogen embrittlement For instance, cracking susceptibility in poorly passivating systems (e.g., austenitic stainless steel in caustic at high temperature, low-carbon steel in caustic or phosphate) will be increased by actions that promote passivation Thus in these systems, cracking susceptibility will be greatest in potential ranges adjacent to active/passive transitions on a polarization curve (Fig 26) In contrast, systems that exhibit strongly passivating behavior (e.g., aluminum alloys, austenitic stainless steels in neutral solutions) will crack most severely under potential conditions where incipient passivity breakdown occurs due to the presence of aggressive anions (e.g., chloride) (Fig 26b)

Fig 26 Schematic electrode-potential/current-density relationship for (a) "poorly" passivating and (b)

"strongly" passivating systems, indicating where severe cracking susceptibility in ductile alloy/aqueous

environment systems is commonly encountered

The fundamental importance of passivation on the crack-propagation process in ductile alloy/aqueous environment systems also indicates an analytical method of determining the potential ranges where cracking susceptibility may be severe An example is the direct measurement of passivation rate (Ref 50), by comparing the bare surface and fully passivated dissolution rates to determine whether a high-aspect-ratio crack is possible, or by performing potentiodynamic scans at various rates Such rapid prediction capabilities are of use in preliminary failure analyses or risk assessments, but this usefulness should be tempered by the realization that these techniques will indicate only the possibility of severe susceptibility

As indicated in Fig 24 and 25, the rate-determining step can change as the system parameters (e.g., corrosion potential, stressing frequency, temperature) lead to an increase in crack-propagation rate Ultimately the rate-determining step will often be liquid diffusion, either of solvating water molecules, anions, or solvated cations to and from the crack tip Under these conditions, the propagation rate will become independent of stress intensity (i.e., the stage II region in Fig 25) and will exhibit a temperature dependence associated with an activation enthalpy of 4 kcals/(g mole) associated with liquid diffusion It is important from a practical viewpoint to realize that this is a limiting condition, and that the activation enthalpy can change continuously between 4 and 30 kcals/(g mole) (symptomatic of passivation control) with corresponding changes in, for instance, loading rate (Ref 29), and temperature (Ref 29) Thus, mechanistic analyses based

on a specific value of activation enthalpy must be treated with caution, unless it is determined that a limiting value is being measured

References cited in this section

5 D.A Jones, Principles and Prevention of Corrosion, 2nd ed., Prentice Hall, 1996

10 M.G Fontana, Corrosion Engineering, 3rd ed., McGraw-Hill Book Co., 1986

26 R.F Steigerwald and N.D Greene, J Electrochem Soc., Vol 109, 1962, p 1026

27 H.H Uhlig and R.W Rene, Corrosion and Corrosion Control, 3rd ed., John Wiley & Sons, 1985, p 217

28 Z Szklarska-Smialawska, Pitting Corrosion of Metals, NACE, 1986

29 F.P Ford, "Mechanisms of Environmental Cracking Peculiar to the Power Generation Industry," Report NP2589, EPRI, 1982

30 F.P Ford, Stress Corrosion Cracking, Corrosion Processes, R.N Parkins, Ed., Applied Science, 1982

31 R.N Parkins, N.J.H Holroyd, and R.R Fessler, Corrosion, Vol 34, 1978, p 253

32 B Poulson and R Robinson, Corr Sci., Vol 20, 1980, p 707

33 J Congleton, "Some Aspects of Crack Initiation in Stress Corrosion and Corrosion Fatigue," paper presented at Corrosion 88, NACE, St Louis, 21-25 March 1988

34 Conf Proc., Environmental-Sensitive Mechanical Behavior (Baltimore, MD, June 1965), A.R.C

Westwood and N.S Stoloff, Ed., Gordon and Breach, 1966

35 R.W Staehle, A.J Forty, and D Van Rooyen, Ed., The Fundamental Aspects of Stress-Corrosion

Cracking, Ohio State University, Sept 1967

36 J.C Scully, Ed., Theory of Stress Corrosion Cracking, NATO, Brussels, March 1971

37 O Devereaux, A.J McEvily, and R.W Staehle, Ed., Corrosion Fatigue Chemistry, Mechanics and

Microstructure, University of Connecticut, Storrs, June 1971

38 M.P Bastein, Ed., L'Hydrogene dans les Metaux, Science et Industrie, Paris, 1972

39 L.M Bernstein and A.W Thompson, Ed., Hydrogen in Metals, L, American Society for Metals, 1973

40 R.W Staehle, J Hochmann, R.D McCright, and J.E Slater, Ed., Stress-Corrosion Cracking and Hydrogen Embrittlement of Iron-Base Alloys, NACE, 1977

41 A.W Thompson and I.M Bernstein, Ed., Proc Effect of Hydrogen on Behavior of Materials (Jackson

Lake, WY, Sept 1975), TMS, 1976

42 R.M Latanision and J.T Fourie, Ed., Surface Effects on Crystal Plasticity (Hohegeiss, Germany, 1975),

Noordhof-Leyden, 1977

43 P.R Swann, F.P Ford, and A.R.C Westwood, Ed., Mechanisms of Environment Sensitive Cracking of

Materials, The Metals Society, April 1977

44 Corrosion Fatigue, Met Sci., Vol 13, 1979

45 T.R Beck, Corrosion, Vol 30, 1974, p 408

46 R.W Staehle, in Theory of Stress Corrosion Cracking, J.C Scully, Ed., NATO, Brussels, March 1971

47 J.C Scully, Corros Sci., Vol 8, 1968, p 771

48 D.J Lees, F.P Ford, and T.P Hoar, Met Mater., Vol 7, 1973, p 5

49 J.R Ambrose and J Kruger, J Electrochem Soc., Vol 121, p 1974, p 599

50 F.P Ford and M Silverman, Corrosion, Vol 36, 1980, p 558

106 K Osozaawa and H.J Engell, Corros Sci., Vol 6, 1966, p 389

Design for Corrosion Resistance

F Peter Ford and Peter L Andresen, General Electric Corporate Research and Development Center; Peter Elliott, Corrosion and Materials Consultancy, Inc

Engineering Design Principles

The earlier sections of this article provide a background about the principles of corrosion This section focuses on engineering aspects of design that can, without due care and attention, precipitate unexpected premature failure More extensive texts relating specifically to design are available (Ref 51, 52, 53, 54), as are guides from various material suppliers and promoters (Ref 55, 56, 57) However, it is the fine details of engineering design, often compounded by human errors or poor communication (Ref 22, 58, 59, 60, 61, 62, 63, 64), that account for many unexpected failures, at times significant On occasion a poor design can cause premature failure of the most advanced corrosion-resistant materials

Design Considerations

Designing for corrosion control can only be effective if it is part of the overall design philosophy However, a designer is seldom a corrosion engineer, so it is necessary to convey corrosion knowledge to the designer Unlike conventional

engineering, the basic difficulty is that corrosion is not a tangible property; it is more a behavioral pattern Thus, to realize

safe, reliable designs, it is essential that there be a rigid control on materials and fabrication and an extensive effort to eliminate human errors or misunderstandings that result from poor communication

The results of a survey of chemical-process plants (Ref 65) showed that design faults ranked highest (58%) in the reasons for failure Of almost equal ranking was the incorrect application of protective treatments (55%), followed by categories that demonstrate a lack of knowledge about the operating conditions (52%), lack of process control (35%), and an unawareness that there was actually a corrosion risk (25%)

In an ideal world, designers would call for some corrosion assessment prior to preparing the detailed engineering design

Typically, schemes would permit some form of evaluation with respect to both function and the necessary action, for example from the proposal-to-production planning stages (Fig 27) (Ref 65) In the practical "real" world, however, communication of "agreed" reasons for failures may not always reach the designer Indeed, communication to contractors, who are closest to the application, is even poorer (Ref 65, 66) Studies have shown that, while management is always informed of the reasons for failure in the chemical-processing industry, site personnel are informed only 77% of the time, designers 55%, material suppliers 37%, and contractors only 11% of the time

Fig 27 Action steps during design Source: Ref 64

"Draftsmen's Delusions." A further complication in designing against corrosion relates to the general interpretation

of design drawings in what has been referred to as "draftsmen's delusions." For example, the draftsman might be considering a certain piece of equipment without knowledge of the fact that there may be unusual shapes, moving parts,

or environmental issues A lack of attention to design detail causes many premature failures by corrosion-related processes

All too often, the designer will have in mind one thing, which in reality becomes totally different For example, a simple cross-over line between two reactor vessels might, in practice, become an extended line with several turns, merely to position a shut-off valve at a more convenient and accessible position closer to ground level There are countless examples of this situation in "real-world" failure analysis, where the actual designs were not those originally intended; examples will be provided below

Quality assurance and control usually ensure that the requested material is what it should be, given rigorous inspection However, material controls at fabrication are sometimes less than perfect In one case, an inspector noticed a blemish on the outside of a steel pressure vessel Upon closer scrutiny (from the vessel interior) the problem was found to

be a steel bolt that had been inadvertently rolled into the steel plate during fabrication

While considering materials, it is important to avoid nonspecific descriptions or terms in reference to design drawings and specifications There are many instances where generic terms, such as "stainless steel," "bronze," "Hastelloy," or

"Inconel," are too vague and the ultimate choice is far from what was expected and required Wherever possible, and notably in high-risk areas, materials should be selected and tested according to code requirements (Ref 67, 68) Substitutes, if requested, should be properly evaluated before use

Reliability Engineering

The designer should play a significant role in reliability management The communication chain, or the "reliability loop" from the designer, to the manufacturer, then to the user and back to the designer, is a key factor (Ref 66, 69) When a risk

is well documented, it should be possible to overdesign or at least isolate the area to minimize risk to users Where a risk

of failure is high, the emphasis should be toward a "fail-safe" or "no-fail/replace" procedure Failures vary considerably: the design function can be partially or totally affected; the onset of failure can be gradual or sudden The combination of sudden and total failure represents the worst catastrophic situation (e.g., explosions, fires, and total structural collapse), many of which can be attributed to "a small design detail."

Corrosion Awareness

This article is intended to improve corrosion awareness, but clearly this is only a starting point To be effective, the user must also be willing to take action and the designer should insist on appropriate codes and/or recommended working practices Whether the necessary action will be taken is affected by financial, technical, safety, social, and/or political issues (Ref 69)

To prevent corrosion/degradation, the designer can:

• Avoid obvious design weaknesses (see examples below)

• Use more reliable materials, even if this entails greater cost

• Introduce additional precautions (inhibitors, cathodic protection, coatings)

• Establish efficient maintenance/repair teams having detailed procedures and including qualified surveyors, inspectors, and supervisors

• Ensure that standby products are available, fully labeled, and properly stored (using dessicants and noncorrosive packaging)

As noted above, the planned approach for reliable engineering design should include corrosion control and/or preventive measures, for which standards and specifications are available Actual (real-world) approaches vary from plant to plant (or component to component) and include the:

• "Avoid failure" approach

• "Keep-it-working" approach

• "Let-it-fail-then-repair" approach

Aspects germane to these approaches are summarized in Table 5 (Ref 66, 69)

Table 5 Design aspects for reliability

into account (typical)

"Avoid failure" approach

Control product duty Fail safe Visible instrumentation

Keep product simple Fewer functions Fewer joints

Predict reliability Maintain in fixed limits Overdesign (thicker), etc

Study variable interfaces Contact faces, bearings Lubrication (wear, friction) Seals Consult experts and test work

Instill care for familiar items

or designs

Complacency Negligence

Study innovation Consult experts and test work Records of operation Records of failure mode/frequency

Redundant items Stand-by units in good repair Storage Handling Vibration, impact, wear, etc

Recognize human limits Misreading of instructions Sequence of controls (hypnosis?) Adequate housekeeping User attitudes

Test product design Proving the design Variables Safety limits Pilot plant and real-life testing

Control malfunctions Human Automatic Mixture Important for software/hardware

"Keep it working" approach

Maintainability Access

Short-term items Replace routinely before failure Regular inspection Monitoring

Spares Identity correctly coded Location known Easy access replacement Avoid identical mating systems

Built-in adjustments Corrects for progressive deterioration Manual/automatic

"Let it fail then repair" approach

Defects Subtle or catastrophic

Spares Need quantified reliability (difficult for corrosion)

Minimize off period Keep repair time low Speedy fault diagnosis Speedy removal of failed part(s) Speedy replacement

Speedy check of assembly Experts, inspectors

Source: Ref 66

New plants and process equipment are commonly designed to prevent or reduce problems that occurred previously Updated and improved procedures are of little use, however, if no action is made until after the equipment fails (Ref 70) Situations have occurred where cathodic protection was ineffective because the anode material was not connected to the structure it was protecting; a monitoring signal (a bell) was ignored because of noise; online monitoring data were estimated (guessed at) so that the inspector could avoid travel in blizzard conditions; and a sentinel hole (weep hole) leaked for weeks without any investigation as to cause until the vessel exploded

Why Failures Occur

In the context of design, there are several factors that relate to material/component failure:

• Overload suggests a weakness in plant control instrumentation or operation

• Abnormal conditions can result from a lack of process control or variations in raw materials (alloy type

or chemical inhibitors, etc.)

• Poor fabrication may relate to inadequate instructions or inspection (e.g., excessive cold work,

overmachining, flame cutting, or excessive torque loading)

• Poor handling Scratches or machine marks can result from poor detailing or poor instruction Not to be

excluded are identity marks (incised codes) and inspection stamps

• Assembly, if incorrect (e.g., welds and fasteners) can seriously influence stress, flow, and compatibility

• Storage and transportation can significantly influence material performance, especially for items

shipped to/from tropical, humid climates where heavy rains, violent seas, storms, and cargo sweat may each contribute to material degradation unless adequate precautions are taken (Ref 71) Proper design

and effective corrosion control management normally should accommodate these aspects of material handling

• Accessibility Some structures or components may not be accessible for remedial work, even if a

corrosion risk is recognized Buried structures can be affected where soil and bacterial corrosion might apply (Ref 72)

Design and Materials Selection

Corrosion control measures are best initiated at the design stage (Fig 27) Materials are usually selected to perform a basic function or to provide a functional requirement (see the article "Materials Selection" in Ref 22) Therefore, in many instances the material choice is dictated not by corrosion, but by characteristics such as strength, reflectivity, wear resistance, and dimensional stability In some situations a corrosion-resistant alloy may not be satisfactory

High-Temperature Service. The subject of designing for high-temperature service is outside the scope of this article

(see the article "Design for High-Temperature Applications" in this Volume and Ref 58, 73, and 74); however, high temperatures always accelerate corrosion processes, and certain gases or liquids, which are considered innocuous under ambient conditions, become aggressive to materials when hot A tenfold change in corrosion rate is not uncommon for a temperature change of 30 °C under aqueous conditions The same tenfold change (or considerably worse) can occur with

a 20 °C change under high-temperature conditions If temperatures are too high, the material might oxidize (i.e., scale) Thick scales and metal loss result from overheating (loss of water cooling, absence of insulation, etc.) Heat-transfer contributions will increase as the scale/deposit thicknesses increase

Candidate high-temperature materials need to be strong and resistant to oxidation or to other corrosion processes that might involve complex multioxidant environments having highly volatile phases and molten salts It is important for designers and others to recognize that several corrosion elements might simultaneously be involved in an application (e.g., oxygen, halogen, and sulfur) (Table 6) Material selection for high-temperature service needs to be reviewed for each individual part and application Alloy steels and more sophisticated alloys based on nickel and cobalt are most commonly used, in which key elements for high-temperature corrosion resistance include chromium, aluminum, silicon, and rare-earth additions for scale retention

Table 6 Types of corrosion and corrodents encountered in high-temperature processes or components

Process components Temperature, °C Types of

corrosion or corrodent Chemical/petrochemical

Ethylene steam cracking furnace tubes to 1000 Carburization, oxidation

Steam reforming tubes to 1000 Oxidation, carburization

Vinyl chloride crackers to 650 Halide gas

Hydrocracking heaters, reactors to 550 H 2 S and H 2

Petroleum coke calcining recuperators 816 Oxidation, sulfidation

Cat cracking regenerators to 800 Oxidation

Flare stack tips 950-1090 Oxidation, thermal fatigue, sulfidation, chlorination, dewpoint

Carbon disulfide furnace tubes 850 Sulfidation, carburization, deposits

Melamine production (urea)-reactors 450-500 Nitriding

Other processes

Titanium production reactor vessels 900 Oxidation, chlorination

Nitric acid catalyst grid 930 Oxidation, nitriding, sulfidation

Nuclear reprocessing reactors 750-800 Oxidation (steam), fluorination (HF)

Oil-fired boiler superheaters 850-900 Fuel ash corrosion

Gas turbine blades corrosion to 950 Sulfates, chlorides, oxidation, ash

Waste incinerators superheaters 480 Chlorination, sulfidation, oxidation, molten salts

Fiberglass manufacturing recuperators 1090 Oxidation, sulfidation, molten salts

Source: Ref 58

Different behavior can arise at similar temperatures, depending on the source of heat, such as electrical heating elements; fuel combustion, flue gases and deposits; flame impingement; friction and wear When hot gases cool, condensation can cause acid dewpoint conditions, thereby changing the material choice to a corrosion-resistant alloy

Design Factors That Influence Corrosion. The basic factors that most influence design for corrosion resistance are summarized in Table 7 (Ref 58) Each factor plays a unique yet not always unrelated role with other factors For example, localized corrosion damage adjacent to the spindle support of a stainless steel paddle stirrer (Fig 28) resulted not only from crevice corrosion (oxygen differential cell) but also because of galvanic corrosion, caused by a small steel retaining screw (anodic) that had been used inadvertently for assembly Ultimately, the stirrer support loosened, which allowed further deterioration by fretting

Table 7 Corrosion factors that can influence design considerations

Environment Natural

Chemical Storage/transit Stress Residual stress from fabrication

Operating stress static, variable, alternating

Shape Joints, flanges

Crevices, deposits Liquid containment and entrapment

Compatibility Metals with metals

Metals with other materials Quality control

Movement Flowing fluids

Parts moving in fluids Two-phase fluids

Temperature Oxidation, scales

Heat-transfer effects Molten deposits Condensation and dewpoint

Control Surface cleaning and preparation

Coatings Cathodic protection Inhibitors

Inspection Planned maintenance

Location. Exposure to winds and airborne particulates can lead to deterioration of structures.Designs that leave structures exposed to the elements should be carefully reviewed, because atmospheric corrosion is significantly affected

by temperature, relative humidity, rainfall, and pollutants Also important are the season and location of on-site fabrication, assembly, and painting.Codes of practice must be adapted to the location and the season

Shape. Geometrical form is basic to design.The objective is to minimize or avoid situations that worsen corrosion These situations can range from stagnation (e.g., retained fluids and/or solids; contaminated water used for hydrotesting) to sustained fluid flow (e.g., erosion/cavitation in components moving in or contacted by fluids, as well as splashing or droplet impingement)

Common examples of stagnation include nondraining structures, dead ends, badly located components, and poor assembly or maintenance practices (Fig 29) General problems include localized corrosion associated with differential aeration (oxygen concentration cells), crevice corrosion, and deposit corrosion

Fig 29 Examples of how design and assembly can affect localized corrosion by creating crevices and traps

where corrosive liquids can accumulate (a) Storage containers or vessels should allow complete drainage; otherwise, corrosive species can concentrate in vessel bottom, and debris may accumulate if the vessel is open

to the atmosphere (b) Structural members should be designed to avoid retention of liquids; L-shaped sections should be used with open side down, and exposed seams should be avoided (c) Incorrect trimming or poor design of seals and gaskets can create crevice sites (d) Drain valves should be designed with sloping bottoms

to avoid pitting of the base of the valve (e) Nonhorizontal tubing can leave pools of liquid at shutdown (f) to (j) Examples of poor assembly that can lead to premature corrosion problems (f) Nonvertical heat exchanger assembly permits dead space that is prone to overheating if very hot gases are involved (g) Nonaligned assembly distorts the fastener, creating a crevice that can result in a loose fit and contribute to vibration, fretting and wear (h) Structural supports should allow good drainage; use of a slope at the bottom of the member allows liquid to run off, rather than impinge directly on the concrete support (i) Continuous weld for

horizontal stiffeners prevents traps and crevices from forming (j) Square sections formed from two L-shape members need to be continuously welded to seal out the external environment

Movement. Fluid movement need not be excessive to damage a material Much depends on the nature of the fluid and the hardness of the material.A geometric shape may create a sustained delivery of fluid or may locally disturb a laminar stream and lead to turbulence Replaceable baffle plates or deflectors are beneficial where circumstance permit their use; they eliminate the problem of impingement damage to the structurally significant component

Careful fabrication and inspection should eliminate or reduce poor profiles (e.g., welds, rivets, bolts), rubbing surfaces (e.g., wear, fretting), and galvanic effects due to the assembly of incompatible components Figure 30 shows typical situations in which geometric details influence flow

Fig 30 Effect of design features on flow (a) Disturbances to flow can create turbulence and cause

impingement damage (b) Direct impingement should be avoided; deflectors or baffle plates can be beneficial (c) Impingement from fluid overflowing from a collection tray can be avoided by relocating the structure, increasing the depth of the tray, or using a deflector (d) Splashing of concentrated fluid on container walls should be avoided (e) Weld backing plates or rings can create local turbulence and crevices (f) Slope or modified profiles should be provided to permit flow and minimize fluid retention

Compatibility. In plant environments, it is often necessary to use different materials in close proximity.Sometimes, components that were designed in isolation can end up in direct contact in the plant (Fig 31) In such instances, the ideals

of a total design concept become especially apparent, but usually in hindsight Direct contact of dissimilar metals introduces the possibility of galvanic corrosion, and small anodic (corroding) areas should be avoided wherever this contact is apparent

Fig 31 Design details that can affect galvanic corrosion (a) Fasteners should be more noble than the

components being fastened; undercuts should be avoided, and insulating washers should be used (b) Weld filler metals should be more noble than base metals Transition joints can be used when a galvanic couple is anticipated at the design stage, and weld beads should be properly oriented to minimize galvanic effects (c) Local damage can result from cuts across heavily worked areas End grains should not be left exposed (d) Galvanic corrosion is possible if a coated component is cut When necessary, the cathodic component of a couple should be coated (e) Ion transfer through a fluid can result in galvanic attack of less noble metals In the example shown at left, copper ions from the copper heater coil could deposit on the aluminum stirrer A nonmetallic stirrer would be better At right, the distance from a metal container to a heater coil should be increased to minimize ion transfer (f) Wood treated with copper preservatives can be corrosive to certain nails, especially those with nobility different from that of copper Aluminum cladding can also be at risk (g) Contact

of two metals through a fluid trap can be avoided by using a collection tray or a deflector

Galvanic corrosion resulting from metallurgical sources is well documented Problems such as weld decay and sensitization can generally be avoided by material selection or suitable fabrication techniques Less obvious instances of localized attack occur because of end-grain attack and stray-current effects, which can render designs ineffective

End-grain attack, or preferential attack of grains exposed by cross-cutting through a metal plate or rod (Fig 31c), occurs

in many corrosive fluids An example is the cut edges and punched-out "holes" in a stainless steel reactor tray (Fig 32)

Fig 32 End-grain corrosion along cut edges and punched holes in a reactor tray made from type 316 stainless

steel

Stray-current effects are common on underground iron or steel pipelines that are located close to electrical supply lines, or where stray currents can cause active corrosion at preferred sites For example, stray-current attack caused a titanium flange spacer to become anodic "during service" due to inappropriate grounding of welding equipment on an adjacent part

of the plant structure

Designers, when aware of compatibility effects, need to exercise their ingenuity to minimize the conditions that most favor increased corrosion currents Table 8 provides some suggestions

Table 8 Sources of increased corrosion currents and related design considerations

Metallurgical sources (both within the metal

and for relative contact between dissimilar

metals)

Difference in potential of dissimilar materials; distance apart; relative areas of anode and cathode; geometry (fluid retention); mechanical factors (e.g., cold work, plastic deformation, sensitization)

Environmental sources Conductivity and resistivity of fluid; changes in temperature; velocity and direction

of fluid flow; aeration; ambient environment (seasonal changes, etc.)

Miscellaneous sources Stray currents; conductive paths; composites (e.g., concrete rebars)

The most common design details relating to galvanic corrosion include jointed assemblies (Fig 29, 31) Where dissimilar metals are to be used, some consideration should be given to compatible materials known to have similar potentials (Fig 13) Care should be exercised because the galvanic series is limited and refers to specific environments, usually seawater

Where noncompatible materials are to be joined, it is advisable to use a more noble metal in a joint (Fig 31b) Effective insulation can be useful if it does not lead to crevice corrosion Some difficulties arise in the use of adhesives, which may not be sealants

The relative surface areas of anodic and cathodic surfaces should not be underestimated, because corrosion at a small anodic zone can be several hundred times greater than that for the same bimetallic components of similar area As noted already, synergistic effects must also be recognized, such as failure resulting from a combination of galvanic and crevice attack This example shows the unfortunate choice of a carbon steel bolt to tighten the spindle in a stainless steel paddle stirrer (Fig 33) The result was crevice corrosion at the stirrer support (Fig 28), which was exacerbated by the galvanic action caused by the carbon steel/stainless steel metal-to-metal contact

Fig 33 The incorrect choice of a carbon-steel retaining bolt for a stainless-steel spindle resulted in localized

galvanic corrosion of the paddle-stirrer assembly (see Fig 28)

Less obvious examples of galvanic corrosion occur when ion transfer results in the deposition of active and noncompatible deposits on a metal surface For example, an aluminum stirrer plate used in water was extensively pitted because the water bath was heated by a copper heater coil (Fig 31e) The pits resulted from deposition of copper ions from the heater element More rapid, but similar, damage occurred when a dental aspirator (Teflon-coated aluminum) was attacked by mercury from a tooth filling These two metals rank as a "high risk" combination for galvanic corrosion (see Fig 13) The aluminum section was rapidly pitted once the Teflon had first been worn away by sharp fragments of tooth enamel

Anodic components can on occasion be overdesigned (made thicker) to allow for the anticipated corrosion loss In other instances, easy replacement is a cost-effective option (Table 5)

Where metallic coatings are used, there may be a risk of galvanic corrosion, especially along the cut edges Rounded profiles and effective sealants or coatings are beneficial Transition joints can be introduced when different metals are to

be in close proximity These and other situations are illustrated in Fig 31 Another possibility is coating of the cathodic material for corrosion control Ineffective painting of an anode in an assembly can significantly reduce the desired service lifetime, because local defects (anodes) effectively multiply the risk of localized corrosion

Insulation represents another area for potential corrosion attack, although most problems arise because of poor installation Insulation types and properties vary considerably, and expert advice from suppliers is recommended The most common corrosion problems include crevice corrosion (where insulation and/or adhesives are tightly held against a metal surface, for example when straps or ties are too tight) and pitting corrosion (where moisture condenses on the metal, usually because the insulation barrier was too thin or was improperly installed) Moisture-absorbing tendencies vary from one insulation to another (Ref 75)

Wet-dry cycling has been known to lead to concentration effects (e.g., chloride ions from calcium-silicate insulation) There have been reported instances of chloride stress-corrosion cracking (SCC) in certain stainless steel pipes and vessels,

or pitting of these and other materials, such as aluminum, when contacted by insulation The early instances of SCC failure were mainly attributable to high chloride levels (500-1500 ppm) associated with asbestos-type materials The chloride levels have been significantly reduced in recent years to a level that is not expected to cause SCC Standards are now available, as are tests to evaluate insulation materials (Ref 76) A parallel German standard calls for zero nitrite content and <0.2% ammonia levels in elastomeric insulation that is used for copper and copper-alloy piping to reduce the risk of SCC (Ref 77)

Figure 34 shows some typical examples in which design and installation procedures could have been improved Other problems occur when insulation is torn or joints are misaligned or incorrectly sealed with duct tape or similar bandaging, none of which is recommended by insulation suppliers

Fig 34 Corrosion problems associated with improper use of insulation and lagging (a) Incorrect overlap in

lobster-back cladding does not allow fluid runoff (b) Poor installation left a gap in the insulation that allows easy access to the elements (c) Outer metal cladding was cut too short, leaving a gap with the inner insulation exposed (d) Insufficient insulation can allow water to enter; chloride in some insulation can result in pitting or stress-corrosion cracking of susceptible materials (e) Overtightened strapping can damage the insulation layer and cause fluid "dams" on vertical runs

Stress. From a general design philosophy, environments that promote metal dissolution can be considered more damaging if stresses are also involved In such circumstances, materials can fail catastrophically and unexpectedly Safety and health may also be significantly affected

A classic example of chloride SCC occurred in a type 304 (UNS S30400) stainless steel vessel (Fig 35) The corrosion cracks extended radially over the area where a new flanged outlet was welded into the vessel Residual stresses (from flame cutting) and the fluids inside the vessel (acidic with chlorides) were sufficient to cause this failure in a matter

stress-of weeks

Fig 35 Chloride SCC in a type 304 stainless-steel vessel after a new flange connection was welded into place

Figure 36 shows examples of using design detail to minimize stress Perfection is rarely attained in general practice, and some compromise on materials limitation, both chemical and mechanical, is necessary Mechanical loads can contribute

to corrosion, and corrosion (as a corrosive environment) can initiate or trigger mechanical failure Designs that introduce local stress concentrations directly or as a consequence of fabrication should be carefully considered

Fig 36 Design details that can minimize local stress concentrations (a) Corners should be given a generous

radius (b) Welds should be continuous to minimize sharp contours (c) Sharp profiles can be avoided by using alternative fastening systems (d) Too long an overhang without a support can lead to fatigue at the junction Flexible hose may help alleviate this situation (e) Side-supply pipework may be too rigid to sustain thermal shock from a recurring sequence that involves (1) air under pressure, (2) steam, and (3) cold water

Of particular importance in design are stress levels for the selected material: the influence of tensile, compressive, or shear stressing; alternating stresses; vibration or shock loading; service temperatures (thermal stressing); fatigue; and wear (fretting, friction) Profiles and shapes contribute to stress-related corrosion, especially if material selection dictates the use of materials that are susceptible to failure by SCC or corrosion fatigue (Ref 10, 22, 30, 78)

Materials selection is especially important wherever critical components are used Also important is the need for correct procedures at all stages of operation, including fabrication, transport, storage, startup, shutdown, and normal operation

Surfaces. Corrosion is a surface phenomenon, and the effects of poorly prepared surfaces, rough textures, and complex shapes and profiles can be expected to be deleterious (Ref 79) Figure 37 shows some examples in which design details could have considerably reduced the onset of corrosive damage resulting from ineffective cleaning or painting

Fig 37 Effects of design on effectiveness of cleaning or painting (a) Poor access in some structures makes

surface preparation and painting difficult; access to the types of areas shown should be maintained at a minimum of 45 mm (1 in.), or one-third of the height of the structure (b) Sharp corners and profiles should

be avoided if the structure is to be painted or coated

Designs should provide for surfaces that are free from deposits; access to remove retained soluble salts before painting; free-draining assemblies; proper handling of components to minimize distortion, scratches, and dents; and properly located components relative to adjacent equipment (to avoid carryover and spillages).Other recommended procedures for coating constructional materials are shown in Fig 38 (Ref 18)

Fig 38 Suggestions for steel construction to be coated (a) Avoid pockets or crevices that do not drain or

cannot be cleaned or coated properly (b) Joints should be continuous and solidly welded (c) Remove weld spatter (d) Use butt welds rather than lap welds or rivet joints (e) Keep stiffeners to outside of tank or vessel

(f) Eliminate crevice (void) at roof-to-shell interface in nonpressure vessel (g) Outlets should be flanged or pad type, not threaded Where pressure limits allow, slip-on flanges are preferred, because the inside surface of the attaching weld is readily available for radiusing and grinding Source: Ref 80

Neglect and poor (or no) maintenance caused localized pitting on the underside of a type 304 stainless steel vessel lid that was exposed to high humidity, steam, and chloride vapors Access in this example was possible, but not used Common engineering structural steelwork requires regular preventive maintenance, and restricted access makes this impossible Figure 37 shows situations in which surface cleaning and/or painting is difficult or impossible Condensation in critical areas can also contribute to corrosion.Typical structures susceptible to this phenomenon include automobile exhaust systems and chimneys or exhausts from high-temperature plants, such as boilers, kilns, furnaces, or incinerators

Painting and surface coating techniques have advanced in recent years and have provided sophisticated products that require careful mixing and application Maintenance procedures frequently require field application where some control (use of trained inspectors) is essential, as in offshore oil and gas rigs Inspection codes and procedures are available and total design should incorporate these wherever possible In critical areas, design for online monitoring and inspection will also be important The human factor in maintenance procedures is often questionable Adequate training and motivation are of primary importance in ensuring that design details are appreciated and implemented

References cited in this section

10 M.G Fontana, Corrosion Engineering, 3rd ed., McGraw-Hill Book Co., 1986

18 R.S Treseder, R Baboian, and C.G Munger, Ed., NACE Corrosion Engineer's Reference Book, 2nd ed.,

NACE, 1991

22 Corrosion, Vol 13, ASM Handbook (formerly Metals Handbook, 9th ed.), ASM International, 1987

30 F.P Ford, Stress Corrosion Cracking, Corrosion Processes, R.N Parkins, Ed., Applied Science, 1982

51 V.R Pludek, Design and Corrosion Control, MacMillan, 1977

52 R.J Landrum, Fundamentals of Designing for Corrosion Control, NACE International, 1989

53 R.N Parkins and K.A Chandler, Corrosion Control in Engineering Design, Department of Industry, Her

Majesty's Stationery Office, London, 1978

54 L.D Perrigo and G.A Jensen, Fundamentals of Corrosion Control Design, The Northern Engineer, Vol 13

(No 4), 1982, p 16

55 Designer Handbooks, Specialty Steel Industry of North America, Washington, D.C.; also publications

relative to design, Nickel Development Institute, Toronto, Canada

56 Guides to Practice in Corrosion Control, Department of Industry, Her Majesty's Stationery Office, London,

1979-1986

57 Engineering Design Guides, Design Council, British Standards Institute, Council of Engineering

Institutions, Oxford University Press, 1975-1979

58 P Elliott and J.S Llewyn-Leach, Corrosion Control Checklist for Design Offices, Department of Industry,

Her Majesty's Stationery Office, London, 1981

59 P Elliott, Corrosion Control in Engineering Design, audiovisual for Department of Industry, United

Kingdom, 1981

60 O.W Siebert, Classic Blunders in Corrosion Protection, Mater Perform., Vol 17 (No 4), 1978, p 33 and

Vol 22 (No 10), 1983

61 T.F Degnan, Mater Perform Vol 26 (No 1), 1987, p 11

62 P Elliott, Why Must History Repeat Itself?, Ind Corros., Feb/March 1991, p 8

63 P Elliott, Process Plant Corrosion Recognizing the Threat, Process Eng., Vol 65 (No 11), 1984, p 43

64 P Elliott, Understanding Corrosion Attack, Plant Eng., Oct 1993, p 68

65 P Elliott, Corrosion Survey, Supplement to Chem Eng., Sept 1973

66 P Elliott, Catch 22 and the UCS Factor Why Must History Repeat Itself?, Mater Perform., Vol 28 (No

7), 1989, p 70 and Vol 28 (No 8), 1989, p 75

67 Standards for Corrosion Testing of Metals, ASTM, 1990

68 R Baboian, Ed., Corrosion Tests and Standards: Applications and Interpretation, ASTM Manual Series,

MNL-20, 1995

69 H.J.H Wassell, Reliability of Engineered Products, Engineering Design Guide, Design Council, Oxford

University, 1980

70 P Elliott, We Never get Corrosion Problems, Super News, 1974, p 70

71 A Sparks, Steel Carriage by Sea, 2nd ed., Lloyd's of London Press, 1995

72 G Kobrin, Ed., Microbiologically Influenced Corrosion, NACE International, 1993

73 P Elliott, Practical Guide to High Temperature Alloys, Mater Perform., Vol 28, 1989, p 57

74 G.Y Lai, High Temperature Corrosion of Engineering Alloys, ASM International, 1990

75 W Pollock, Corrosion under Wet Insulation, NACE International, 1988

76 "Specification for Wicking-Type Thermal Insulation for Use Over Austenitic Stainless Steel," C 795,

Annual Book of ASTM Standards, ASTM

77 "Codes of Practice for Drinking Water Installations (TRWI)," 628.1.033:696.11:620.193, DIN, Teil 7, 1988

78 H.H Uhlig, Corrosion and Corrosion Control, 2nd ed., John Wiley & Sons, 1971, p 314

79 C.G Munger, Corrosion Prevention by Protective Coatings, NACE International, 1984

80 P.E Weaver, "Industrial Maintenance Painting," RP0178, NACE International, 1973, p 2

Design for Corrosion Resistance

F Peter Ford and Peter L Andresen, General Electric Corporate Research and Development Center; Peter Elliott, Corrosion and Materials Consultancy, Inc

Life Prediction and Management

As pointed out in previous sections, corrosion degradation is largely understood mechanistically, and logical mitigation actions and design decisions can be formulated with a reasonable scientific basis However, in recent years there has been

a further requirement in some industries that the lifetime of various components be predicted and that the technical (and economic) benefits of life extension actions be quantitatively defined Such further design requirements for life management are especially seen, for instance, in the aerospace and power industries

To meet these particular design criteria, it is necessary to derive accurate and verifiable life prediction methods An example is given below of such a derivation for the example of environmental degradation in light-water nuclear reactors (LWRs) The specific corrosion phenomena in these structures have been widely publicized and discussed, especially in a series of conferences (Ref 81, 82, 83, 84, 85, 86, 87, 88, 89, 90) organized by the National Association of Corrosion Engineers (NACE, now known as NACE International), the American Institute of Mining, Metallurgical, and Petroleum Engineers (AIME), and the American Nuclear Society (ANS), and in books (Ref 91, 92) Moreover, this particular topic

of corrosion in nuclear reactors has been addressed in the ASM Handbook (Ref 93) These incidences have covered all

forms of corrosion, including erosion corrosion, galvanic attack, general corrosion, intergranular corrosion, SCC, and corrosion-fatigue cracking in, primarily, carbon/low-alloy steels, austenitic stainless steels, nickel-base alloys, and zirconium-base alloys To one extent or another, these phenomena have affected the integrity of most of the components

in reactors, including irradiated core structures, piping, tubing, pressure vessel, service water systems, and heat exchangers Moreover, the publications have addressed the whole spectrum, from the mechanistic aspects of degradation

in a particular alloy/environment system to the adequacy of the existing design and life prediction methods

The objective of this section is to focus on one form of degradation, environmentally assisted cracking, not only because

it is one of the severest forms of life-limiting phenomena, but also because it represents an area where "fundamental" approaches to resolving the design/prediction problems are most advanced

There have been well-documented instances of environmentally assisted cracking (e.g., stress-corrosion and assisted cracking and corrosion fatigue) in high-temperature water of austenitic stainless steels, nickel-base alloys, low-alloy and carbon steels, and their weld metals in various subcomponents of LWRs These events have had a considerable economic impact (Ref 81) and are occurring at a time when competitive and deregulation pressures are forcing nuclear utilities (at least in the United States) to reduce costs (Ref 82) Thus, there is an incentive to develop validated life prediction methods to (a) evaluate the extent of a problem (i.e., is an incident an isolated occurrence or is it the precursor

strain-to a more widespread generic problem?) and (b) quantitatively define the benefit of a specific mitigation action and evaluate its cost-effectiveness

Unfortunately, the scatter in data from which such design or life prediction codes can be formulated is extreme This scatter applies to data obtained from both field experience (Ref 94) and "well-controlled" laboratory experiments This latter observation is illustrated in Fig 39 by crack-propagation rate/stress-intensity relationships for stainless steels (Ref 96) Similar data scatter has been noted for nickel-base alloys (Ref 97) and low-alloy steels (Ref 97) in 288 °C high-purity water In other instances, a large database does not exist, so it is impossible to define a usable life prediction algorithm

An example of this is irradiation-assisted stress-corrosion cracking of stainless steels in reactor cores

Fig 39 Crack-propagation rate vs stress-intensity data for stainless steel in 288 °C water Also shown is the

Nuclear Regulatory Commission disposition line Source: Ref 95

Consequently, the current life prediction codes for environmentally assisted cracking of structural materials in 274 to 345

°C water, representative of the coolant, used in boiling-water reactors (BWRs) and pressurized-water reactors (PWRs),

usually represent upper bounds of the database available at a given time Moreover, a single intensity relationship (e.g., ASME Section XI for life prediction/evaluation) or cyclic-stress/cycles-to-crack-initiation relationship (e.g., ASME Section III for fatigue design) is usually quoted with no account of the different cracking susceptibilities that are associated with the wide range of material/environment conditions in LWR systems This conservative "upper bound" scenario can put unreasonable constraints on the continued operation of specific plants that have been operating under system conditions far better than those on which the design or life evaluation code was originally based

crack-propagation-rate/stress-Various statistical (extreme-value, pattern-recognition) or neural-network approaches can be applied to the available cracking database to develop an algorithm for component life, or crack depth, as a function of all the reactor operating parameters This "classical" approach is limited, however, because there are comparatively few datasets for crack propagation, for example, that have been obtained under adequate control of system conditions (e.g., water purity, loading) and using sensitive enough techniques that continuously monitor propagation rates that are relevant to design lives of the order of 40 to 60 years for LWRs

This situation is offset by the increase in mechanistic understanding of environmentally assisted cracking, which has accelerated in the last 20 years due to the availability of experimental and analytical procedures that allow quantification and validation of various cracking hypotheses Because the precise shape of the crack depth/operational time relationship must be a function of the specific material, environmental, and stressing conditions, it follows that if there is a range in the actual system operating conditions, then there will be a predictable range in the observed cracking susceptibility, with the distribution of the range in cracking mirroring the distribution of system conditions that affect the propagation process This predictable range in observed cracking susceptibility provides a bridge between deterministic and statistical/probabilistic life prediction methods (Ref 96, 99)

Historically, environmentally assisted cracking has been divided into the "initiation" and "propagation" periods To a large extent, this division is arbitrary, because in most investigations, "initiation" is defined as the time at which a crack is detected, or when the load has relaxed a specific amount (in a strain-controlled situation) Such a definition of initiation, however, can correspond to a crack depth of significant metallurgical dimensions (e.g., >2 mm) Thus, for the purpose of lifetime modeling, it is proposed (Fig 40) that, phenomenologically, initiation is associated with microscopic crack formation at localized corrosion or mechanical-defect sites and is generally related to pitting, intergranular attack, scratches, weld defects, or design notches It is further proposed that the probability is high of such microscopic initiation sites existing or developing relatively early in the life of the component Then the problem of life prediction devolves to understanding the growth of small cracks from these geometrically separated initiation sites, and the coalescence of these small cracks to form a major crack, which may then accelerate or arrest, depending on the specific material, environment, and stress conditions

Fig 40 Proposed sequence of crack initiation, coalescence, and growth for steels undergoing subcritical

cracking in aqueous environments

Attention has been focused primarily on the propagation rates for "deep" cracks (e.g., approximately one grain diameter and greater), recognizing that the resultant life prediction can be conservative because the microscopic crack initiation and crack coalescence periods are not being accounted for

Numerous mechanisms have been advanced for crack propagation in ductile alloy/aqueous environment systems The usual procedure in developing a quantitative predictive model for crack propagation is to advance a hypothesis for the cracking mechanism and then independently quantify the relevant parameters in that mechanism The resultant prediction algorithms are then validated by comparison with observed crack-propagation data Once validated, these algorithms are used for design-life-extension decisions This process is described very briefly below for stainless steels in 288 °C water (the boiling-water reactor coolant), where the hypothesis is made that the cracking mechanism is slip dissolution (Ref 96,

100, 101)

Quantification of the Slip-Oxidation Mechanism for Stainless Steels in 288 °C High-Purity Water.

Various theories have been proposed to relate crack propagation in various ductile alloys, including stainless steels, to oxidation rates and the stress/strain conditions at the crack tip These theories were supported by an early correlation between the average oxidation current density on a straining surface and the crack-propagation rate for several systems (Ref 101, 102) Experimentally validated elements of these earlier proposals have been combined such that crack propagation in many systems can be correlated with the oxidation that occurs when the protective film at the crack tip is ruptured (Ref 103, 104)

Consider the change in oxidation charge density with time following the rupture of a protective film at the crack tip Initially, the oxidation rate and, hence, crack-advance rate are rapid and are typically controlled by activation or liquid diffusion kinetics as the bare metal dissolves In most LWR cracking systems, a protective oxide rapidly reforms at the bared surface and the total oxidation rate (and crack-tip advance) will slow with time Thus, crack advance can be maintained only if the film rupture process is repetitive Therefore, for a given crack-tip environment and material condition, the crack-propagation rate will be controlled by the change in oxidation-charge density with time and the frequency of film rupture at the strained crack tip This latter parameter is determined by the fracture strain of the film, f, and the strain rate at the crack tip, ct By invoking Faraday's law, the average environmentally controlled crack-propagation rate, t, can be related to the oxidation-charge density (metal oxidation current integrated over time) passed

between film-rupture events, Qf, and the strain rate at the crack tip, ct:

(Eq 23)

where M and are the atomic weight and density of the crack-tip metal, respectively

Because the oxidation-charge density on a bare surface varies with time at a rate that is dependent on the material and environment compositions, Eq 23 can be reformulated in terms of a power-law relationship:

Thus, the starting point in the development of a quantitative prediction method for environmentally assisted cracking is

Eq 24 However, to develop a useful methodology, it is necessary to redefine this fundamental equation in terms of

measurable engineering or operational parameters This redefinition involves: (a) defining the crack-tip alloy/environment composition (e.g., in terms of bulk alloy composition, anionic concentration or solution conductivity, dissolved-oxygen

content or corrosion potential); (b) measuring the reaction rates for the crack-tip alloy/environment system that corresponds to the "engineering" system; and (c) defining the crack-tip strain rate in terms of continuum parameters such

as stress, stress intensity, and loading frequency Extensive work has been conducted in these areas, which has been reviewed elsewhere (Ref 10)

As a result of these examinations of the tip metallurgical, chemical, and stressing conditions, practical propagation-rate algorithms of the following form have been developed for stainless steels in 288 °C BWR water:

n = f ( , EPR, c) (Eq 26)

where is the conductivity of coolant ( S · cm-1), c is the corrosion potential of the steel (mVSHE), EPR is the

measurement of grain-boundary chromium depletion due to heat treatment annealing or welding, K is the stress intensity

(ksi ), app is the applied strain rate (s-1), is the cyclic-loading frequency (s-1), K is the stress-intensity amplitude under cyclic loading, and AR is a parameter that is a function of the mean stress under cyclic loading

Validation of Life-Prediction Algorithms and Their Application. The overall comparison between the observed and theoretical crack-propagation rates in type 304/316 stainless steels in 288 °C water is shown in Fig 41 The laboratory database upon which this comparison was made was obtained under a wide range of stressing (static, monotonically increasing, and cyclic load), material (solution annealed vs various degrees of sensitization) and water composition (<10 ppb O2 to >8 ppm O2, <0.1 to 10 S · cm-1) It is seen that there is a reasonable agreement between observation and prediction

Fig 41 Comparisons between observed and theoretical crack-propagation rates for type 304/316 stainless

steels in 288 °C water This database represents a wide combination of stressing material and environmental

conditions Source: Ref 96

Changes in corrosion potential within the range expected in BWRs can have a significant effect on the cracking susceptibility of type 304/316 stainless steels, especially under constant-load conditions This predicted and observed effect is illustrated in Fig 42 for furnace-sensitized type 304 stainless steel under constant stress intensity (25 ksi )

in water with the conductivity in the range 0.1 to 0.3 S · cm-1 It is seen that over the corrosion potential range -550

mVSHE to +250 mVSHE (spanning "hydrogen-water" conditions to those under "normal" core conditions) the propagation rate can change three orders of magnitude From an operational design viewpoint, therefore, it is seen that considerable benefit may be predicted by developing actions that lower the corrosion potential of the stainless steel structures, thereby highlighting remedial actions that lower the effective concentration of oxidants (oxygen, hydrogen peroxide) in the coolant Solution conductivity is also predicted to have an effect on the cracking susceptibility, as indicated by the three theoretical relationships shown in Fig 42, thereby highlighting the quantitative value of maintaining water-purity control

crack-Fig 42 Observed and predicted sensitivity of stress-corrosion-cracking sensitivity to corrosion potential for

sensitized type 304 stainless steel in 288 °C water The data points are measurements made in the laboratory

or in reactors The curves are the predicted relationships for the indicated conductivities The numbered data points were obtained at the Harwell variable-energy cyclotron The circled numbers were with the proton irradiation turned on, and the uncircled numbers were with the irradiation off Similarly the data point * was obtained under fast neutron irradiation in a boiling-water-reactor core

So far, the comparisons between observation and theory have centered on material/environment systems variables that

affect n in Eq 25 and 26 The effect of stressing/straining conditions on the cracking susceptibility occur primarily

through their effect on the crack-tip strain rate in Eq 27, 28, and 29 It follows that because the crack tip does not

recognize how the strain rate is maintained, the cracking susceptibility for a given material/environment condition should

adhere to the same crack-propagation rate/crack-tip strain-rate relationship, regardless of the stressing/straining mode The truth to this statement is illustrated in Fig 43, which shows the theoretical and observed crack-propagation rate strain-rate relationship for a severely sensitized type 304 stainless steel in 8 ppm O2, 0.5 S · cm-1 water Movement along the strain-rate axis has been achieved by increasing stress intensity under constant-load conditions, increasing applied strain rate under monotonically increasing strain conditions, or cyclic loading under a variety of stress-intensity amplitude, mean stress, and loading frequency conditions The single theoretical relationship line in Fig 43 adequately predicts the cracking under this wide range of loading modes, indicating that the prediction method applies to stress-corrosion cracking (SCC), strain-induced cracking (SIC), and corrosion fatigue (CF)

Fig 43 Predicted and observed crack-propagation rate/crack-tip strain-rate relationships for sensitized type

304 stainless steel in 8 ppm oxygenated, 0.5 S · cm -1 purity water at 288 °C

The old lore that these types of cracking (SCC, SIC, CF) are separate phenomena with, by implication, different mitigation or design modification needs is probably incorrect For instance, it follows from Eq 25 that the sensitivity of the cracking susceptibility to the crack-tip strain rate will be a function of the material/environment conditions that affect

n (Eq 26) Thus, the slope of the crack-propagation-rate/strain-rate relationship will be relatively shallow for severe

environmental and material conditions (e.g., high dissolved oxygen, impure water, and high degrees of grain-boundary sensitization), and the relationship will be steep for less severe material/environmental conditions This predicted and observed (Fig 44) change in propagation-rate/strain-rate dependency with system conditions is significant when evaluating the validity of accelerated tests that are often used for development of design codes For instance, increasing the crack-tip strain rate, and hence cracking susceptibility, by using the "slow-strain-rate test" is a valid test acceleration procedure (because it is accelerating one of the rate-determining steps in the cracking mechanism), but the factor of improvement between a reference condition and a proposed mitigation condition will be less in this test than at the lower stressing or strain-rate conditions expected in the operating plant The relationship (i.e., Fig 44) also gives an explanation for the lore that the cracking susceptibility is more dependent on the specific environmental conditions under constant-load stress-corrosion conditions than under corrosion-fatigue conditions

Fig 44 Predicted and observed crack-propagation rate/crack-tip strain-rate relationships for stainless steels in

a variety of material/environment systems

In summary, therefore, it is apparent that the crack-prediction algorithms are able to quantitatively explain the changes in crack-propagation rates for type 304/316 stainless steel in water at 288 °C for a wide combination of water composition (corrosion, potential, conductivity), material sensitization, and stressing (constant load/displacement, cyclic load) conditions It follows, however, that because the cracking response is so sensitive to changes in combinations of system conditions, it is necessary to combine the predictive method with system-defining sensors and models (Fig 45) Provided this combining is done, it is then possible to make predictions of the extent of cracking in specific plant components (Fig 46) and the increase in life associated with specific system changes (Fig 47)

Fig 45 The integration of system monitors, sensors, and environmental/material models as inputs to a

crack-propagation-rate model

Fig 46 Theoretical and observed intergranular stress corrosion crackdepth vs operational-time relationships

for 28 in diameter schedule 80 type 304 stainless steel piping for two boiling-water reactors operating at different mean coolant conductivities Note the bracketing of the maximum crack depth in the lower-purity plant by the predicted curve, which is based on the maximum residual-stress profile and the predicted absence

of observable cracking in the higher-purity plant (in 240 operating months)

Fig 47 Predicted crack depth vs time response for defected 28 in diameter schedule 80 recirculation piping in

a given boiling-water reactor to defined changes in water purity Also shown is the crack-depth limit that can be resolved by nondestructive testing (NDT)

References cited in this section

10 M.G Fontana, Corrosion Engineering, 3rd ed., McGraw-Hill Book Co., 1986

81 R.L Jones, "Corrosion Experience in U.S Light Water Reactors NACE 50th Anniversary Perspective," Paper 168, presented at Corrosion 93, NACE, 1993

82 R.L Jones, "Critical Corrosion Issues and Mitigation Strategies Impacting the Operability of LWRs," Paper 103, presented at Corrosion 96, NACE, 1996

83 Conf Proc., Environmental Degradation of Materials in Nuclear Systems Light Water Reactors, J

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84 Conf Proc., Environmental Degradation of Materials in Nuclear Systems Light Water Reactors, J

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85 Conf Proc., Environmental Degradation of Materials in Nuclear Systems Light Water Reactors, J Weeks

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86 Conf Proc., Environmental Degradation of Materials in Nuclear Systems Light Water Reactors, G

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88 Conf Proc., Environmental Degradation of Materials in Nuclear Systems Light Water Reactors, R Gold

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89 Conf Proc., Environmental Degradation of Materials in Nuclear Systems Light Water Reactors, R Gold

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Design for Corrosion Resistance

F Peter Ford and Peter L Andresen, General Electric Corporate Research and Development Center; Peter Elliott, Corrosion and Materials Consultancy, Inc

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