Ref. 12.4-1 describes the difficulty of defining damage-causing temperatures as follows:
“In many cases, it is difficult to make qualitative or quantitative statements regarding the damage-causing effect of thermal loads, because there are a large number of effects that contributed to the damage (Ill. 12.1-3).” Even the term “thermal load” is problematic upon closer consideration. The first problem is determining above which temperature the damaging thermal loads begin. Only in rare cases is the part temperature the only load, as in the case of an unallowable structural change due to overtemperature. Normally, impeded thermal strain that is inseparable from the temperature effect occurs in a hot part and causes corresponding mechanical stress. Many other effects which are not causally dependant on temperature, such as mechanical loads, chemical attacks, and metallurgic processes, can in many cases be damaging at certain temperature levels (Ill. 12.1-3), but should be dealt with separately. It is not always necessary for the temperature to exceed normally allowable values for damage to occur. Therefore, temperature is often not the cause of damage.
In the following, “damage caused by thermal loads” is understood to be “damage that would not have occurred without the influence of temperatures that are considerably higher than room temperature (>600°C).”
The following examples focus on material behavior. The engine part-specific behavior is explained more thoroughly in the chapters on hot parts, i.e. the combustion chamber, turbine, afterburner, tailpipe, and thrust reverser.
Illustration 12.4-1: Material properties react to temperature changes and have a considerable influence on damage mechanisms. In addition to strength, almost all design-relevant material properties are affected by increasing temperatures (Ref. 12.4-7). These changes are usually not linearly related to the temperature, nor do they always change in the same direction. This greatly increases the complexity of engine part design and dimensioning. The following are descriptions of several effects, but with no claim to completeness:
Static strength is dependent on time and temperature. The loss of strength accelerates as temperature increases (Ill. 12.5.1-2). This behavior is additionally influenced by the fact that changes occur in the material when it is under stress. Examples include pore formation, changes of the precipitation phase, or the appearance of brittle phases. Factors such as the orientation of the grain boundaries and the grain size influence the location and frequency of load-specific weak points and demand special attention.
Breaking strain: Normally, during hot tensile tests (short-time), the breaking strain increases with temperature. However, some Ni-based alloys have a more or less pronounced minimum (Ref. 12.4-10) at temperatures around 600°C, above which the plastic deformability may increase exponentially (bottom right). At very high temperatures that already lead to the softening of the material (solidus temperature), a material that has been ductile up to that point may behave very brittly (Ill. 12.4-1). Naturally, this tendency is also noticeable in the notch impact strength, which is important for the part behavior in the case of FOD and containment situations.
Strain behavior during long-term stress: The creep strain is dependent on the temperature and stress levels. The creep strain generally increases along with the stress levels until fracture (Ills. 12.5.1-3 and 12.5.1-4). If deformation parameters are properly selected, then superplastic behavior can be attained. In this case, the recovery processes in the material are balanced by the damage due to the strain. This behavior is used when forming metal.
Elasticity: The E-modulus usually decreases as temperature increases in a material-specific curve. This is important for the vibration frequencies of a turbine blade, for example, and must be considered in order to prevent resonance. The E-modulus is usually not dependent on the time at high temperature. An interesting phenomenon is the orientation-dependency of the E-modulus in the thermal strain of the single-crystal version of Ni-based alloys (also see Ill. 18.104.22.168-7) for the prevention of resonance. The orientation of the grains is also noticeable in the thermal cycle behavior (Ill. 12.6.2-7).
Hardening and plastic deformation: Forging processes use the strength losses and increased plastic deformability at high temperatures, but also take advantage of the reduction in hardenings. Conversely, targeted thermo-mechanical treatment can be used to increase strength. This is due to faster diffusion processes.
Residual stresses and hardenings: Creep leads to changes and induction of residual stresses and reduction of hardenings. If desired compressive residual stresses and hardenings are reduced by machining or shot peening, the operating behavior will be worsened. For example, a long interval before repeated regenerative shot peening can lead to a dynamic fatigue fracture in a titanium alloy blade root that is subjected to fretting. Creep strain can also shift and/or induce residual stresses. Thermal fatigue is based on this effect (Chapter 12.6.2). If creep deformation occurs in a limited part area (e.g. compression in the hot leading edge of a turbine blade, Ill. 12.5.1-14), residual stresses build up in this area after cooling.
Notch behavior: if the notch effect decreases due to creep-dependent stress reduction in the notch, the dynamic strength increases relative to that at room temperature. This can be true for both LCF and HCF. The notch impact strength also normally increases with temperature. This behavior is especially important for coatings such as diffusion coatings and metallic overlays of type MCrAlY. Similar to parts made from intermetallic phases, these coatings behave brittly up to several hundred °C before showing satisfactory ductile behavior. This makes them prone to cracking at low operating temperatures (start-up and shut-down), which has a negative influence on their thermal fatigue behavior (Chapter 12.6-2). High operating temperatures, on the other hand, lead to creep strain in the coating and cause it to wrinkle (Ill. 12.6.2-15). However, at very high temperatures in the solidus range, the notch effect can decrease considerably.
Structure: Structures can change at high temperatures. At very high temperatures, recrystallization with the formation of new grains can occur in the area of plastic deformations. If the solution annealing temperature of a hardening phase is exceeded, then the material may suffer a permanent strength loss. This phase can withdraw again later during long-time operation. The hardening phase can also orient itself dependent on the direction of the main stress (e.g. coarsening). The dropping-out of brittle phases (e.g. sigma-phase, Laves-phase) over longer times can affect the toughness and therefore also the FOD behavior, for example.
High temperatures cause diffusion coatings to undergo changes such as pore formation or redistribution of alloy components (Ills. 12.4-4 and 12.4-6).
Fracture- and notch impact strength, fracture mechanical behavior: Both parameters usually experience a characteristic steep rise as temperature increases (Ill. 12.3-5). The steep rise in the fracture toughness occurs at a lower temperature than that of the notch impact toughness. This also influences properties such as the growth-capable flaw size or the crack growth under cyclical loads.
Thermal strain: The thermal strain coefficient normally increases linearly with temperature. The thermal strain is very important for maintaining clearances at seals (Volume 2, Chapter 7.1.2). Nickel alloys have greater thermal strain than titanium alloys. This makes the transition from titanium alloys to nickel alloys in housings and rotors in the rear compressor area difficult. For this reason, a disk made from steel with suitable thermal strain (13% Cr steel) is inserted between Ni and Ti disks. This steel disk can have corrosion problems during operation. Fittings and bolts that do not loosen at high temperatures require coordinated thermal strain across the entire operating temperature range. The thermal strain naturally also influences the thermal fatigue behavior. A high thermal strain coefficient will perform especially poorly. In ceramic thermal barriers, especially, the thermal strain relative to the metallic base material is decisive.
Thermal conductivity: The thermal conductivity is important for the heat-up and cool-down behavior of the engine parts and therefore vital for maintaining clearances. During rubbing, the thermal conductivity influences the heating-up and damaging of the parts. There are materials such as aluminum and magnesium that have a thermal conductivity that is virtually independent of temperature in the operating temperature range. Ni-based alloys and titanium alloys, on the other hand, both exhibit a considerable increase in thermal conductivity at higher temperatures. The thermal conductivity of low alloy steels decreases considerably up to temperatures of about 500°C. The thermal conductivity of high alloy steels such as 13% Cr steels and 18/8 CrNi steels increases as temperatures approach 500 °C (Ref. 12.5.1-2).
Illustration 12.4-2 (Ref. 12.4-8): The higher the temperature of Ni-based alloys, the greater their proportion by volume of g'-phase. This diagram shows that the solution annealing temperature increases as the high g' proportion increases. Because the g' phase is an important indicator for damaging overtemperature, these temperatures are important. Experienced professionals can draw conclusions regarding the exceeded minimum temperature from a metallographic investigation of the changes of the g' phase (Ref. 12.4-12). These conclusions become less precise if longer operating times following the overheating cause the g' phase to drop out again.
Illustration 12.4-3: Evidence that overheating occurred, along with possible clues regarding the time, duration, and temperature levels of the overheating, is extremely important for understanding damage and therefore also a requirement for successful specific remedies. Even the macro-damage symptoms of a blade can reveal a great deal (top diagrams). In extreme cases with part temperatures in the melting range (liquidus temperature) and/or extensive oxidation (burning, Ills. 22.214.171.124-7 and 126.96.36.199-9), blade sections will be missing without indentifiable fracture surfaces. Instead, the separation plane will be rounded off and rough.
If the temperatures were a bit lower, in the range of grain boundary softening (solidus temperature), then blade zones will split and remain open and/or break off. This usually occurs at the blade tips and at the especially thin rear edge. Despite oxidation, fractured structures are recognizable in all cases.
In the described cases, it can be assumed that the damage is consequential and that other parts were also unallowably damaged. For this reason, engine-specific overhaul manuals often also prescribe the scrapping of the turbine disks of the following stages.
Crack fields that are oriented in a parallel direction in the edge area, with no pronounced orange-peel effect, indicate that a thermal shock occurred. This is a type of overheating of very short duration (Ill. 12.6.2-3). This cracking can be especially pronounced in parts with a diffusion coating that is brittle at certain temperatures (Ill. 12.6.2-15).
If high temperatures (at the upper limit of the design) act constantly and/or cyclically over a long time on part areas (especially blade inlet edges), they can cause “washed-out” short cracks (due to erosion and spalling of the oxides, Ill. 12.6.2-10) and “wrinkled” surface material removal that is usually dark in color (orange peel effect, Ills. 188.8.131.52-1, 184.108.40.206-7, and 12.6.2-10). This situation indicates poor cooling (partial blocking of cooling air ducts) or poor temperature distribution in the gas flow, for example. In the case of boroscope inspections, it indicates that the blading is almost at the end of its life. However, it does not mean that a spontaneous failure is imminent.
The bottom diagrams show typical signs of overheating in the micro-range. Fractures that occur near the solidus temperature, especially in single crystal blades, may have a pronounced crystalline structure (cleavage cracks). This is similar to fatigue fractures that occur due to dynamic stress at normal operating temperatures. A characteristic sign of overheating is droplets on the crack surfaces (middle diagram) that can be seen with an electron microscope (Ref. 12.4-4). This allows one to differentiate between these fractures and fracture surfaces with stage 1 cracks (Ills. 12.2-4 and 12.2-6).
Short-time temperatures that are considerably above the solution annealing temperature can be detected by changes in the hardening phase (g' phase). For Ni-based alloys, material-specific solution annealing temperatures are usually in the range of 950 °C to 1250°C (Ill. 12.4-2). Depending on the part temperature and duration of overheating, this phase goes into either partial or total solution (bottom right diagram). One must be careful of misinterpretations when a long period (hours) of normal operating temperatures followed the overheating. In this case a new hardening effect can be expected.
If one is very “lucky”, the part will have a coating that demonstrably reacts to overtemperature. Signs of this include cracking, diffusion with changed distribution of alloy elements, structural changes, and fused material (bottom middle diagram, Ill. 12.4-6).
Illustration 12.4-4: Liquid metal embrittlement (LME) has some of the same necessary conditions as stress corrosion cracking (Volume 1, Chapter 220.127.116.11). The danger of LME is present when a metallic part with sufficiently high tensile stress is coated by a metal melt. The important factor for dangerous coating with metal melt is the metallic contact. This requires situations such as new part surfaces or oxide coatings that have been torn open to the substrate. These conditions can be caused by LCF that occurs as defined with noticeable plastic deformations. The metal melt shoots into the material on the grain boundaries, and has an embrittling and crack-inducing effect. This damage can occur both immediately after contact with the melt, or later, such as when contaminants on the surface are melted by operating temperatures. The top left diagram depicts a case in which silver dripped onto the surface of a new part (metallic, unoxidized surface) during heat treatment in a vacuum furnace (top right diagram). The furnace had earlier been used for soldering with silver, and remnants of the silver solder had evidently become trapped in the insulating graphite mats above the part. A tiny drop of solder (smaller than the head of a pin) penetrated several centimeters into the disk. The disk was under sufficiently high tensile stress (probably forging stress and/or heat stress, top middle diagram).
The bottom diagrams show SEM images (Ref. 12.4-3). On the left, one can see the crack in the area of the silver droplet surrounded by other smaller droplets. This appearance is typical for a coating droplet and differs from smeared contaminants, which are also dangerous. The right diagram shows the formation of “flowerlets” that is typical for liquid metal embrittlement. These consist of small collections of solder on the gaping intercrystalline fracture surfaces (grain boundaries) at the start of the crack.
Dangerous metallic contaminants can end up on part surfaces in many different ways. This is especially common during manufacture and assembly.
Illustration 12.4-5: The external inspection of damaged thermally-stressed parts can help determine the causes, mechanisms, types, and extent of damage. This information can then be used to design measures for further inspections or remedies.
Cracking: The appearance of the cracks can reveal a great deal of information regarding their causes and damaging influences. A primary sign is the position of the crack in the structure. If the crack is intercrystalline (along the grain boundaries), it indicates the presence of a creep effect and/or a damage due to corrosion and/or oxidation. If the crack is transcrystalline (through the grain boundaries), dynamic fatigue (dynamic fatigue fracture, thermal fatigue, LCF) is more probable. A gaping crack that has not grown yet (crack edges fit together), and with no obvious plastic deformations, indicates brittle behavior under residual stress. One should investigate whether or not embrittling materials such as metal melts (Ill. 12.4-2) may have been present. Plastic deformations are signs of high strain due to overstress. Compressed and gaping crack edges are signs of temporary high compressive stress above the flow limit (e.g. thermal fatigue, Ill. 12.6.2-10). Tightly closed cracks indicate the presence of high compressive stress, such as that following a relocation of residual stress due to plastic deformation. The crack base can be especially important (Ill. 12.6.2-10). Important clues regarding the crack growth rate and possible environmental factors are rounding, oxidation, branching, and crack fields. The location and direction of the crack on the engine part yields important information regarding the direction of the damaging loads. Influences that can be dated to before or after the crack formation (e.g. machining grooves, scratches, splashed material) hold important information regarding the time of the damage.
Fracture pattern: Despite oxidation, the fracture pattern can usually be macroscopically evaluated in the same way as cracks and fractures that occurred at “non-oxidizing” temperatures. In fact, tarnishing can actually be advantageous in that it contains information regarding the temporal progress of the crack.
Although hot fracture surfaces, especially fresh reactive surfaces, have oxide coatings, experience has shown that successful microscopic analysis (SEM) is usually still possible (Ill. 12.4-7).
Melting and cracking of solder: The melting point of high-temperature solder increases during the soldering process as the low-melting components are diffused (boron, etc.). Due to the lower strength and brittleness of solder relative to the base material, the solder may crack open at high temperatures despite its sufficiently high melting point (Ill. 12.4-7). If the liquidus temperature of the solder is exceeded, the resulting increase in volume can cause pearls of solder to be secreted from the surface. Microscopic inspection of this type of fracture surface can use melt structures and doughy behavior (the SEM will show chewing gum-like fracture structures) to draw conclusions regarding overtemperatures. Similar secretions can be found on infiltrated sintered materials (such as in bearings and pressure disks; Volume 1, Ill. 5.1.5-2), on extremely inhomogenous materials, or in the area of segregations in Ni-based cast alloys.
Splashed pearls of melted material: This is another type of local overheating and damage. Melt pearls from welding and machining processes can compromise the strength of the surface they strike, and can also induce tension residual stresses. This decreases the dynamic strength considerably. Cases involving electron beam-welded titanium parts are well documented (compressor stators, disks). However, hot particles can also be scattered during friction welding, cutting, or grinding. Titanium parts are sensitive to splashed welding melts in two ways:
Drops of melted titanium alloys burn during flight in the air and are extremely hot. The poor thermal conductivity of titanium also slows cooling.
Titanium materials oxidize heavily and embrittle in the area of the drops. In addition, titanium alloys are sensitive to notches. The embrittlement promotes both micro-cracks during deformation and also dynamic fatigue fractures that originate in these cracks.
Indications of surface embrittlement: Cracking, especially in a pattern of concentric rings around imprints or craters, indicates embrittlement of the surface. In titanium alloys, this is usually an oxide coating that forms during contact with oxygen at temperatures above 600°C. Tarnishing is an indicator, and grey matte oxidation surfaces are a sure sign of a dangerously embrittled titanium surface. Embrittlement can also be expected in nickel alloys if foreign metal has melted onto the surface (brittle phases, embrittlement of the grain boundaries through liquid metal embrittlement, Ill. 12.4-4).
Tarnishing: Tarnishing is usually a sign of oxygen coming into contact with a hot surface. The colors depend solely on the thickness of the coating (colors of thin platelets). The oxide coating thickness is influenced by many other factors in addition to temperature, including the reactivity of the surface (e.g. material, fracture surface), atmosphere, fouling, and the reaction time. For this reason, only very vague conclusions can be made with regard to the maximum temperature. A great deal of experience and additional information (e.g. time) are necessary for a reasonably reliable estimation of temperatures. Special care is required in the case of titanium alloys. Their tarnishing changes decisively due to the presence of even small amounts of fouling.
Fouling and deposits: Special attention must be given to spots, splashes, or streaks that evidently altered the oxidation behavior of the surface. It is possible that a damaging surface reaction such as a grain boundary attack or diffusion may occur. This may mean that no protective oxide coating can build up in this zone.
Distortion of a part can be pure creep deformation under external stress and/or the result of released or relocated residual stress. This can occur at sufficiently high temperatures during normal operation. Another possibility for deformation is unusually large restricted thermal strain due to temperature gradients. Deformations should always be considered in connection with high residual stresses, the size and orientation of which can be measured either in the laboratory or in the field, depending on accessibility and conditions. Procedures that use X-rays can only measure the residual stresses of a thin surface zone. The hole-bore method is preferable for determining the stresses beneath the surface. Elaborate (expensive) “tomographic” procedures that use neutrons can accurately reveal the distribution and size of stresses even in thick cross-sections. These procedures border on destructive inspection, since holes must be drilled in the part. Unrestricted creep deformations do not create any noteworthy residual stresses.
Hardness: Changes in hardness can be measured in suspect part zones. Considerable changes in hardness beyond the given values for the part indicate a structural change. High operating temperatures may offer an explanation for this. This is true for both increases and decreases in hardness.
Metallographic inspection (see Ills. 12.4-8 and 12.4-10): If the surface of a part is treated like a metallographic section (polished, etched), then it can be directly inspected microscopically, depending on accessibility. Otherwise, an indirect inspection is possible with the aid of a synthetic mold (solidifying mass or foil).
Illustration 12.4-6 (also see Ills. 12.4-5, 12.4-7, and 12.4-8): The metallographic findings of thermally stressed parts is an important aid for assessing the extent of damage and determining damage mechanisms, processes, and causes. Even scanning electron microscopy has not made this inspection method obsolete, and the two can complement each other excellently when used together. This is especially true for micro-analysis and the analysis of extremely small particles such as precipitation phases or grain boundary precipitation.
Structural changes: Procedures such as solution annealing, precipitation (e.g. Ni alloys), hardening, tempering (e.g. C steels, Cr steels) and sensitizing (e.g. CrNi steels) can be recognized/verified. Of special import is the analysis of the appearance of the g' phase in the Ni alloys of the hot parts (Ill. 12.4-2). The precipitation particles are generally too small for an optical analysis. This analysis requires sufficient experience and background information and is preferably done with a combination of metallographic cuts and preparation, and an SEM inspection. Size, shape, volume, alignment, and orientation (rafting, see Ref. 12.5.1-6) are important indicators of operation-specific changes, and also indicate temperatures, times, and the acting direction of the stress.
Creep pores: Creep pores present the possibility of detecting the presence of life span-determining creep stress. This can be used to estimate the used life span (and therefore also the remaining part life, Ills. 12.5.1-7 and 12.5.1-9). Even with a great deal of experience, sufficiently accurate estimation of part life is only possible if the pore formation is not too pronounced. This, in turn, allows decisions to be made regarding reinstallation and/or regeneration.
The orientation of the pore-covered grain boundaries allows conclusions regarding creep-stressed part zones and the direction of stress. The incited creep mechanism can indicate the type and size of the loads (Ill. 12.5.1-8).
Grain boundaries: The location and extent of damaged grain boundaries (e.g. grains that have broken out) near the surface or in the cross-section can also yield important information regarding damaging influences. Accumulation or depletion of alloy components and foreign elements indicates material weaknesses and/or damage mechanisms. These can be oxidation or reactions with surface fouling. A typical case is liquid metal embrittlement (Ill. 12.4-4).
The sensitization of grain boundaries, such as occurs in insufficiently stabilized CrNi steels in certain temperature ranges (Volume 1, Ills. 18.104.22.168-9, 22.214.171.124-10), next to weld seams for example, can cause unallowable sensitivity to corrosion. If the part is coated (e.g. diffusion coating or applied coating), changes can be important indicators of damage (Ill. 12.4-8). Grain boundaries are primarily subjected to oxidation and hot gas corrosion (e.g. sulfidation). These damages can easily be recognized and evaluated in a cross-section. This type of analysis is required for estimating necessary reworking during repairs. Grain boundary precipitations or depletions indicate the sensitivity of the material to damage mechanisms. Melting of the grain boundaries indicates temperatures in the solidus range. Hot cracks should be mentioned in this context (Volume 2, Ill. 7.2.2-9.2), and are observed during welding, rubbing (Ill. 12.4-10), and grinding-induced overtemperatures.
Laminar damage: Materials that are used at temperatures with life-limiting oxidation usually survive due to the creation of a dense protective oxide coating. Accidents and disruptions of the desired oxide coating, such as splashes of Al following compressor damage, must be considered relevant to damage. These incidents include zones that react with fouling. One example is foreign metals such as silver from labyrinth coatings or silver-plated parts. These deposits can considerably accelerate sulfidation (Ill. 12.4-14). Especially dangerous are reactions with metal melts that weaken and embrittle the base material. Typical examples include lead from forgotten X-ray markings, shot peening remnants from unsuitable covering bands, or low-melting integral cast alloys (e.g. tin/bismuth).
Oxide coatings form rapidly on titanium alloys at temperatures above 600°C. In combination with structural changes, this causes embrittlement and the danger of cracking and a loss of dynamic strength. Damaged or insufficiently oxidation-resistant oxide coatings can form due to the influence of external factors (atmosphere, fouling). Sulfidation is promoted especially by oxide coatings that are not sufficiently dense due to a lack of oxygen (e.g. in poorly aerated hollow chambers, Volume 1, Ill. 126.96.36.199-2).
Crack symptoms: Whether a crack is gaping or closed, and whether its edges are compressed or pulled at the point of origin, depends on plastic deformations and induced residual stresses. However, these characteristics can hardly be used as a basis for conclusions (qualitative, if any) regarding these processes. Under ideal conditions, trans- or inter-crystalline, single or reticulating cracks can be assigned to damage mechanisms and load levels (e.g. creep or vibrations). The location and direction of the crack relative to weak points reveals information about other damage-relevant influences. Weak points include a special grain orientation, cavities, or production flaws, such as machining grooves or spark erosion surfaces.
The shape of the crack tip can be especially revealing with regard to the crack growth rate (Ill. 12.6.2-10). Pronounced oxidation with rounding of the crack base indicates a very slow crack growth rate. A sharp crack tip without oxidation indicates rapid, and therefore dangerous, crack growth.
Grain formation: Recrystallization indicates very high operating temperatures (extremely close to solidus temperature). New grains form in single-crystal materials in the area of critical plastic deformations. This type of structural change is caused by FOD or shot peening. Flawed grains in single crystal materials are caused by problems with the casting process. Typical symptoms include plastic deformations due to shrinkage during cooling in an overly stiff mold.
In forged material, unusual grain size (too large or small) may be related to a heat-treatment (Ill. 12.5.1-16). Grain growth indicates overly long annealing times in production or unusually high operating temperatures over long periods. This is more likely to affect static parts such as combustion chamber components, which are subjected to relatively minor mechanical loads.
Hardness: Absolute values and gradients of micro-hardness determined by metallographic inspection can be used in combination with other (material-specific) findings (e.g. structural anomalies) to evaluate damaging influences. For example, dangerous overheating of the annulus area of turbine disks (hot gas encroachment, Ill. 188.8.131.52-11) can be detected during overhaul by a decrease in strength relative to specified minimum values.
Illustration 12.4-7: In many cases, thermally stressed engine parts can be successfully inspected with a scanning electron microscope. The SEM can be used to evaluate and document the fracture surface structures, and also analyze their composition.
It is not always necessary to prepare the part itself or a destructively removed sample for the specimen chamber of the SEM. In some cases, impressions done with foils or molding materials and metallized with gold for the inspection are sufficient.
The composition of spalled particles and coatings can be analyzed with an SEM. In addition, the structure (e.g. porosity) of deposits and particles can reveal information regarding the place (e.g. compressor wear products) and time (e.g. certain ingested dusts) at which they originated (Ill. 184.108.40.206-14).
Fracture surface analysis: In many cases, the oxidation of the fracture surface is not so advanced that analysis with an SEM is not possible. In any case, it is a mistake to try to remove the (formed!) oxide layer in an attempt to find an undamaged fracture surface for analysis underneath.
Another situation involves coatings that have splashed onto the fracture surface, such as after heavy rubbing. Depending on the medium and how it acts in proper conditions (e.g. ultrasound bath), it may or may not be possible to remove the coatings without damaging the fracture surface.
Striations that are created in the HCF range (Ill. 12.2-6) are usually so fine, that they can no longer be recognized after short oxidation times. The crack progress of thermal fatigue cracks (LCF), on the other hand, can often be easily recognized by the crack growth lines, their point of origin can be identified (possible initial weak point), and their speed and number of cycles can be determined. Even the area where the crack originated can often still be inspected for causal weak points (e.g. cavities). The thickness and distribution of the oxidation coating can yield information about its temporal progress. Melt beads or pores (creep pores) are characteristic symptoms. It should be mentioned that fractures created in a laboratory (preferably after cooling the sample with liquid nitrogen) should be in areas with suspected creep damage. These unoxidized fracture surfaces can be excellent for analyzing creep pores. One must be aware that creep pores that have grown together have “soft” fracture structures, which makes them easy to confuse with part areas that have been melted by extreme overheating.
Evaluation of metallographic cuts or metallographically prepared surfaces: The precipitation phase (g' phase ) can be made analyzable with the aid of a proper sequence of polishing and etching. Due to the small particle size, this cannot be done optically to a satisfactory degree, and must be rather be done in an SEM. This enables conclusions regarding exceeded threshold temperatures (solution annealing, Ill. 12.4-2), direction of stresses, and long-term changes. Similar methods have proven themselves in the analysis of creep porosity. For example, this can make the evaluation of damage (i.e. residual life span) and the direction of stress possible (Ill. 12.5.1-7).
Microanalysis: In addition to the inspection of surfaces and spalled particles, it is also possible to determine the distribution of element concentrations. This can be extremely helpful for evaluating changes in diffusion coatings (Ill. 12.4-8). Identifying the diffusion of damaging elements such as silver or sulfur (sulfidation, Volume 1, Ill. 220.127.116.11-4) is also very important for understanding damage.
Illustrations 12.4-8 (Ref. 12.4-1): Diffusion coatings serve to prevent unallowable oxidation and hot gas corrosion on hot part surfaces. They are frequently manufactured with the aid of aluminum diffusion at high temperatures in suitable atmospheres. At operating temperatures, the aluminum forms a protective aluminum oxide coating. These coatings are especially suited to minimizing oxidation. In other applications with considerable corrosion, such as sulfidation protection, other metals (chromium, etc.) are diffused in (chromizing). In the diffusion process during production, part of the coating with an especially high Al content grows on the base material (build-up zone). Another part diffuses into the base material (diffusion zone). Diffusion coatings are usually brittle at low operating temperatures up to a few hundred °C (Ill. 12.6.2-15). Only above these temperatures are they able to plastically absorb strain of the type that occurs due to thermal fatigue. Therefore, the startup and shutdown process is especially important for cracking in diffusion coatings. This cracking, in turn, influences the thermal fatigue behavior. The coatings may change during operation, resulting in typical effects (Ill. 12.6.2-15):
“A”: Coating cracks occur at low temperatures at which the coating is still relatively brittle.
“B”: Creation of “oxide nails” at coating cracks near the coating base. The relatively low Al content no longer provides sufficient oxidation protection in this area.
“C”: Rippling due to cyclical thermal strain and plastic coating deformation at high operating temperatures.
“D”: Delamination of the coating due to cracking in brittle phases.
“E”: Diffusion processes during operation cause brittle phases to form. This affects the base material and coating in dependency on temperature and time.
“F”: Pore formation at the transition to the substrate in connection with the Kirkendall Effect (diffusion processes to even out alloy components).
“G”: Spalling of the coating due to a combination of oxidation, erosion, and cyclical thermal strain. This is the main criterion for coating life.
“H”: Structural changes in the coating can give professionals important clues regarding the actual operating temperatures. For example, certain structural changes reveal the occurrence of damaging overtemperatures.
“I”: Interdiffusion between the substrate and coating changes the structure and composition of both. These diffusion processes can also occur at normal temperatures, albeit slowly. They are minimized with the aid of so-called diffusion barriers, such as intermediate layers of platinum, for example.
“K”: Melting of the transition zone with coating separation at temperatures around 1250 °C.
Illustration 12.4-9: Several individually cast turbine stator blades are often soldered together into segments. During repair, extensive soldering is used to close cracks and/or build up damaged blade areas (oxidation). The solder is made from powder that has a similar composition to the blade material. In order to attain a melting point that is sufficiently below the softening point of the base material, the powder is mixed with a lower-melting additive such as boron. A paste-like binding agent provides good workability. The melting point of this connection rises due to diffusion during the soldering process until it is near that of the base material. Despite this, overheating will cause the solder to soften before the substrate reaches its solidus temperature. This causes the solder to crack open and/or secrete beads of melt. These characteristics can therefore be seen as indicators of high overtemperatures.
Illustration 12.4-10: Rubbingcan result in dangerous overheating (Volume 2, Ills. 7.2.2-9.2 and 8.2-21.2). This softens or melts the grain boundaries and causes them to crack under the simultaneously occurring thermal strain (hot cracks, bottom left detail). This cracking is generally oriented perpendicular to the direction of rubbing. The top diagram shows the example of a small gas turbine. An air-directing metal plate rubbed on the back side of the centripetal turbine disk, resulting in a damaged ring zone with a radially oriented crack field (arrow). Similar damages occur during machining in the production process (grinding, high-speed milling, separating) if the selected procedural parameters are not optimal. Because the surface being worked is usually greasy, the cracks are often discovered only after a heat treatment cycle, considerable operating time (e.g. during overhaul), or etching.
Unusual tarnishing can indicate damaging high operating temperatures. Areas with unusually intense tarnishing can also occur for other reasons in the annulus (hot gas incursion) or hub area (oil fires, etc.). A sign of dangerous overheating is a drop in strength below specified values and an increase in diameter due to permanent expansion of the disk (bottom diagram).
Oxidation and Hot Gas Corrosion
Damage due to oxidation and hot gas corrosion has already been discussed in Volume 1, Chapter 5.4.5. The following text supplements the earlier information.
Terms (also see Volume 1, Chapter 5.4.5): Technical literature defines oxidation as various damage mechanisms that all ultimately lead to oxide formation. Therefore, oxidation can be seen as a superordinate concept for various types of damage, and also as a subordinate concept of high-temperature corrosion. Oxidation can occur through direct reactions of the substrate with oxygen from the atmosphere, or through chains of reactions and diffusion processes (e.g. sulfidation).
Undesirable oxide formation damages engine parts in various ways. Geometric changes, strength losses, notch effects (grain boundary corrosion), and accelerated crack growth can act separately or in combination. The oxidation of bond layers worsens the bond strength of the overlaying coating (e.g. ceramic thermal barrier coating, Ill. 12.4-12).
Oxidation is unavoidable at the temperatures typical for engine hot parts. These temperatures are actually a prerequsite for preventing damaging oxidation, which is done through the formation of protective dense oxide coatings (usually Al2O3). For this reason, new parts are often pre-oxidized. The goal is to optimize these protective properties, which is done through a suitable selection of alloy components for the base material (generally, high Al content) and/or with the aid of surface coatings containing Al. Protective oxide coatings are thermally stable, seal out oxidizing gases, inert to damaging environmental factors, slow-growing, and bond strongly.
At high temperatures, all alloy components create oxides in differing amounts. If oxides are not sufficiently stable, they will chemically combine with one another. After long periods, only especially stable oxides remain. The oxidation rate depends especially on the alloy composition and the topography (or form) of the surface (Ill. 12.4-13). The protective effect of Al diffusion coatings is not good on high-temperature solders of the type frequently used on hot parts.
Oxide coatings grow through the already-present oxide layer through the diffusion of metal and/or oxygen. Oxide coatings have a greater volume than the metal from which they form, creating internal stresses in the oxide between the oxide and base material. Residual stresses cause oxides to break off, resulting in separation of the surface. The degree of this separation is determined by the life span of the anti-oxidation coating. This is true for both diffusion coatings and applied coatings (e.g. MCrAlY coatings, Ill. 12.4-13). The oxidation life span of an aluminized coating increases with the thickness and Al-content of the coating.
Oxides on long-running hot parts are also more chemically stable than the base material against substances used in repair and manufacturing processes (acids, etc.). This creates considerable problems for the chemical removal of oxides (e.g. for invasive testing). There is a danger of unallowable damage (intercrystalline corrosion, etc.) being done to the material before the oxide is satisfactorily removed. For this reason, oxides are usually removed through a combination of abrasive (e.g. blasting) and chemical methods.
Illustration 12.4-11 (Ref. 12.4-4): These diagrams show the behavior of metallic and ceramic coatings on turbine blades in the hot gas flow, as determined by calculations of a life span model (compare with Volume 1, Ill. 18.104.22.168-4, sulfidation on metallic materials, and Ill. 22.214.171.124-4 for ZrO2 thermal barriers). The grey zone indicates the temperature range in which aggressive fused salts form. The schematic diagram at top left shows the zones of sulfidation for NiCrAlY-type metallic protective coatings, depending on the operating and material parameters. The dotted/solid border line indicates the life span under pure oxidation, with no sulfidation. The top right diagram shows the damage to ZrO2 thermal barriers under the influence of typical fused salts (Volume 1, Ills. 126.96.36.199-4 and 188.8.131.52-5). As one can see, the ceramic coating is also damaged by the fused salts. The life span without the influence of fused salt corresponds to the oxidation of an NiCrAlY-type bond coating (Ill. 12.4-13). Unlike the metallic coating, the ceramic coating has no pronounced minimums.
The bottom diagrams show the behavior of ZrO2 thermal barriers with a NiCrAlY bond coating on a Mar-M 247 Ni-based cast alloy. The diagrams show the influence of important operating parameters such as altitude, sulfur content, and pressure of the hot gas in the turbine. It is important to note that these diagrams are estimates based on computer calculations and therefore only reflect reality in their tendencies.
Influence of altitude: The bottom left diagram shows the considerable increase in sulfidation as altitude decreases. This is due to the increased amount of salt in the ingested air. Sea salt is necessary to break through the protective oxide coating so that the sulfur can have a corrosive effect. For this reason, aircraft that operate at low altitudes above water (e.g. submarine hunters), and/or operate from runways near the ocean, are understandably affected by especially intense hot gas corrosion. This can even increase through compressor washing, if salt deposits are carried into the turbine. On the other hand, engines of commercial aircraft, especially long-distance aircraft, are relatively free from sulfidation, since life span is determined by oxidation at altitudes above a few thousand meters. However, intense sulfidation in hollow LPT blades of civilian engines shows that this conclusion is not absolute (Volume 1, Ill. 184.108.40.206-2).
Influence of the sulfur content: Understandably, the coating life drops sharply as the amount of sulfur increases (bottom middle diagram). Oxidation only becomes dominant around 1000 °C, since the damaging fused salts evaporate at these temperatures.
Influence of hot gas pressure: The bottom right diagram shows that the pressure of the hot gas considerably increases sulfidation. The given pressures must be seen in connection with the “temperature window” for sulfidation (between 550°C and 950 °C). Low-temperature sulfidation (type II sulfidation) is more likely to occur in the low-pressure turbine, where the gas pressures are already quite low. However, pressures around 25 bar occur in the high-pressure turbine, where the part temperatures are around 900 °C. Modern engines in civilian applications have compressor exit pressures of up to 40 bar. From this perspective, type I sulfidation is the most likely in these engines.
The damage rate of a thermal barrier coating is determined by the interplay of the following factors:
Illustration 12.4-12 (Ref. 12.4-3): The chemical and physical properties (Ref. 12.4-11) of thermal barrier coatings (TBC) change under the influence of high temperatures, mechanical loads, and fouling (e.g. dust deposits). This affects the operating properties, resulting in phase changes, grain growth, and sintering. In order to minimize these effects, the oxide Y2 O3 is added to the ZrO2 , which stabilizes the TBC.
Sintering causes the E-modulus to increase after only a short period. Even a heat treatment of one hour at 1250°C results in a noticeable sintering effect (Ref. 12.4-4). This effect increases with higher temperatures (right diagram). It causes a decrease in fracture toughness and therefore earlier cracking under (thermal) strain.
Silicates formed out of dust ingested during afterburner operation and melted in the combustion chamber build up on the high-pressure blading. These deposits accelerate sintering in the thermal barrier coatings. Silicon and calcium, especially, can penetrate into the ceramic coating. This hinders thermal contraction if the dust melt fills the gaps between the columnal crystals in physical vapor deposition coatings, or fills the segmentation cracks in thermal spray coatings (Ill. 220.127.116.11-4). During cooling, the gaps in the TBC can no longer close, creating compressive stress. This induces high tensile stress in the bond coating (Volume 1, 18.104.22.168-4 and Ill. 22.214.171.124-4), promoting crack initiation and crack growth.
Even at a thickness of merely 0.005 mm, thermally grown oxides (TGO) underneath the TBC on the bond coating, cause local delamination of the ceramic and promote thermal fatigue in weak areas of the base material.
The left diagram shows the influence of oxidation of the substrate surface or bond coating on the life span, which is characterized by spalling. Oxygen can reach the bond coating in two ways: through cracks and gaps, and/or through diffusion by the TBC, which acts as an ion conductor. If an oxide coating has formed to the degree that the bond strength is no longer sufficient, the coating will break off. The sensitivity of systems to this damage mechanism varies greatly (grey field in the left diagram).
Illustration 12.4-13: If one disregards cracking due to thermal fatigue (Ill. 12.6.2-15), the life-determining criteria for anti-oxidation coatings is material removal. This is determined by a combination of several influences such as oxidation, coating changes (diffusion), spalling of the oxide coatings (cyclical thermal fatigue), and erosion through particles and/or hot gases (Volume 1, Chapter 5.3.1). Consistent with cyclical tests in hot gas, practical experience has shown that the material removal (or spalling) is clearly related to the curvature of the coated surface (top diagram). Convex surfaces are most likely to spall. This damage increases as the curvature flattens (larger radius), the coating thickens, and the compressive stress increases (Ref. 12.4-13). Flat surfaces perform comparatively well, and concave surfaces perform considerably better. One explanation for this phenomenon can probably be found in the residual stresses that occur during temperature cycles in the coating and substrate. Compressive stress has a lifting effect on convex coatings, but presses the coating down on concave surfaces (top diagram). The bottom diagram shows the behavior of diffusion coatings and applied coatings such as thermal spray coatings. Oxidation resistance and hot gas corrosion resistance are not necessarily mutually inclusive. For example, systems with a high Cr content should be highly resistant to sulfidation, but they are sensitive to oxidation. This is also true for CoCrAlY coatings. On the other hand, NiCrAlY coatings are highly resistant to oxidation, but offer little protection against hot gas corrosion. CoNiCrAlY coatings seem to be a good compromise and offer acceptable protection from both types of damage. The most commonly used Al diffusion coatings consist of aluminides and are especially oxidation resistant. Platinum improves the hot gas corrosion performance of the aluminides.
The primary requirement for satisfactory remedies for damage due to oxidation and hot gas corrosion is determining the type of damage, i.e. the main damaging influence. Only after this has been accomplished can a suitable protective coating be selected with any hope of success. Verification of the improvement always first requires sufficiently realistic reproduction of the damage.
Example 12.4-1 (Ill. 12.4-14, Ref. 12.4-11):
Excerpt: “…(The OEM) is planning to replace silver coating on F110-GE-100 engine high pressure turbine disk front and rear fasteners after an inspection revealed that small cracks had developed in the disk during intensive testing….after more than 4,000 ground operation cycles (equivalent to about 2,000 flight hours). Small cracks were found after a routine inspection with a fluorescent penetrant of a disk…
The silver coating used on nuts and bolts securing the disk's front and rear retainer had reacted with sulfur in the engine's jet fuel, creating a sulfate compound that exposed the disk surface to acid. The acid led to corrosion and about half a dozen cracks in the disk.
Appearance of the cracks on the rear edge of the disk rather than the forward edge indicated that the problem was not structural…Forward edge cracks are related to structural changes in the engine because of the heat generated during the engine operations. The silver material was used to prevent the fasteners from seizing and to allow bolt removal after engine run.
…(the OEM) is planning to replace the silver coating on all disk nuts and bolts to prevent the cracks from developing in other … engines.”
Comments: The described phenomenon has evidently been known for quite some time (Ref. 12.4-5) and has been observed in several cases (Ref.12.4-7). Continued use of silver coating on potentially threatened nuts and bolts would be a clear case of false economy.
Illustration 12.4-14: Silver is used on hot parts made from Ni-based materials and high-alloy steels in order to prevent galling and fretting and to obtain a controllable coefficient of friction. For this reason, nuts, bolts, and their fitting surfaces are often silver-coated. Silver can dangerously damage parts made from nickel alloys (Volume 1, page 126.96.36.199-7) and titanium alloys.
Diagram 1: Initiation and promotion of sulfidation through a type of catalytic effect of silver on hot part surfaces (also see Volume 1, Ill. 188.8.131.52-4).
Diagram 2: At higher part temperatures (probably >700 °C) silver can dangerously diffuse into Ni alloys and high-alloy steels. This is especially likely if these parts are under high tensile stress, as is common in nuts and bolts. This damage occurs primarily in the thread area. It causes embrittlement, strength losses, and fractures. The diffusion is promoted in silver-coated parts by the metallic contact between the silver and the substrate. Damage through diffusion upon contact with silver is less likely in parts with protective oxidized surfaces.
Diagram 3: Incitement of sulfidation on hot parts made from Ni alloys through contact with silver-coated surfaces. This can also unallowably reduce the fatigue resistance (Refs. 12.4-5 and 12.4-11, see Example 12.4-1)
Diagram 4: Ref. 12.4-5 describes two compressor disks made from the titanium alloy Ti-7Al-4Mo that burst after cracks occurred in the bolt bores of the rotor shroud. The cracking was attributed to the contact between the titanium alloy and high-alloy steel (A-286) bolts that were silver-coated against fretting. Condensation water containing Cl evidently led to the formation of silver chloride at the higher operating temperatures, which resulted in silver deposits occurring in the bolt bores. This type of fouling, especially chlorides, can always be expected in marine environments. Ref. 12.4-7 mentions that long-term action of Ag has been observed to cause damage to the disk material, Waspalloy. Ag has also been observed to separate from the bolts and damage neighboring parts.
Diagram 5: Pitting corrosion occurred near silver deposits in the Ni-alloy rotor of a low-pressure turbine. The deposits probably occurred due to evaporated condensation water which contained dissolved silver compounds. The aggressive water (marine atmosphere?) apparently de-silvered the threaded connectors while standing, and was then thrown outward into the flange sockets when the engine was started up, and later evaporated. Therefore, this is a combination of the damage mechanisms from diagrams 1 and 4.
Diagram 6: This HPT disk from a fighter engine is made of an Ni alloy. Cracks occurred in both flanges after longer test runs (see Example 12.4-1). They were caused by the silver of the threaded connection.
In order to prevent these types of silver-related damage, the contact surfaces and glide surfaces must not be silver-coated. This has the drawback, which must be accepted, of making detachable connections (bolts, etc.) impossible to open without damaging them to the extent that they cannot be reused.
Summary: Silver-coated fasteners such as nuts and bolts should not be used in the hot part area.\\
12.4-1 A.Rossmann, “Untersuchung von Schäden als Folge thermischer Beanspruchung”, contribution to J.Grosch “Schadenskunde im Maschinenbau”, Volume 308, from the series “Kontakt & Studium Maschinenbau”, Expert Verlag, ISBN 3-8169-1202-8, 2nd Edition 1995, pages 162-187.
12.4-2 L. Engel, H. Klingele, “Rasterelektronenmikroskopische Untersuchungen von Metallschäden”, Carl Hanser Verlag München Wien, ISBN 3-446-13416-6, page 58.
12.4-3 T.E. Strangman, J.F. Neumann, A. Tasooji, “Thermal Barrier Coating Life Prediction, Model Development”, NASA Host Program (NAS3-23945). Proceedings N88-11183.
12.4-4 P.A. Langjahr, R. Oberacker, M.J. Hoffmann, “Langzeitverhalten und Einsatzgrenzen von plasmagespritzten CeO2 - und Y2O3-stabilisierten Zr O2-Wärmedämmschichten”, periodical “Materialwissenschaft und Werkstofftechnik”, 32, (2001) , pages 665-668.
12.4-5 ASM Handbook, Formerly Ninth Edition, Metals Handbook, Volume 13, Corrosion, chapter on “Corrosion in Aircraft Industry”, page 1041.
12.4-6 Ch.W. Siry, H. Wanzek, C.-P. Dau, , “Aspects of TBC service experience in aero engines”, periodical “Materialwissenschaft und Werkstofftechnik”, 32, 2001 , pages 650-653.
12.4-7 Rabi.S.Bhattacharya, S.Krishnamurthy, A.K. Rai, J.A.Kramer, “Threaded Fastener Coatings for Aerospace Applications”, periodical “Lubrication Engineering”, Volume 52, 3 , pages 237-242.
12.4-8 A.K.Koul, “Hot Section Materials for Small Turbines”, proceedings of the AGARD meeting “Technology Requirements for Small Gas Turbines”, October 1993, pages 40-1 to 40-9.
12.4-9 C. Sommer, M. Bayerlein, W. Hartnagel, “Deformation and Failure Mechanisms of DS CM 247 LC Under TMF and LCF Loading”, proceedings CP-569 of the AGARD meeting “Thermal Mechanical Fatigue of Aircraft Engine Materials”, 2-4 October 1995, pages 11-1 to 11-11.
12.4-10 Fa. Inco, prospectus specifications from “High Temperature High Strength Nickel Base Alloys”, July 1977.
12.4-11 “F110 Disk Crack Traced to Coating”, periodical “Aviation Week & Space Technology”, August 6, 1984, page 18.
12.4-12 J.M. Aurrecoechea, W.D. Brentnall, J.R. Gast, “Service Temperature Estimation of Turbine Blades Based on Microstructural Observations”, periodical “Journal of Engineering for Gas Turbines and Power”, April 1991, Vol.113, pages 251-259.
12.4-13 D.M. Nissley, “Thermal Barrier Coating Life Modeling in Aircraft Gas Turbine Engines”, periodical “Journal of Thermal Spray Technology”, Volume 6 (1), March 1997, pages 91-98.