MATERIALS SELECTION IN ATMOSPHERIC FLUIDIZED BED COMBUSTION SYSTEMS John Stringer Electric Power Research Institute Palo Alto, California, U.S.A. Syn...

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MATERIALS SELECTION IN ATMOSPHERIC FLUIDIZED BED COMBUSTION SYSTEMS John Stringer Electric Power Research Institute Palo Alto, California, U.S.A.

Synopsis The materials problems in an atmospheric fluidized bed combustion (AFBC) system include the erosion and corrosion of in-bed heat exchangers and their support structure; erosion of the bed containment; and perhaps also mechanical fatigue of the heat exchanger and thermomechanical fatigue of the distributor plate and associated structure, although examples of these failures have not been reported· Erosion, and perhaps corrosion, of the first-stage cyclone must be guarded against, although the problems should not be major. The various solids feed systems—coal, acceptor, bed material, recirculated cyclone catch—will suffer from wear problems unless appropriate precautions are taken. Corrosion of gas pass heat exchangers should not be a problem, and erosion should also be no problem so long as the appropriate gas velocity limits are observed. In-bed corrosion is probably controllable for boilers; for air-heaters the situation is not so clear. In-bed erosion does not appear to be an intrinsic characteristic of an AFBC, but is often experienced: the reasons for the appearance of erosion are not well-understood at the moment, which makes this problem a matter of particular concern. Introduction Any energy conversion system presents a range of problems in materials selection. The principles to guide the selection are relatively straightforward: o

The material selected for a component should have properties appropriate for the principal function of that component. Thus, if it is load-bearing, it should have adequate strength; if it is an electrical component it should have appropriate resistivity.


The material selected should have other properties which will enable it to survive in normal opration. Thus, although the prime requirement may be high temperature strength, some room-temperature toughness may also be required. However, if this is incompatible with the first requirement, it may be possible to modify the design or the operation to take account of the material's limitation.


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The material selected should have appropriate durability. A material may degrade in service by (for example) precipitation of a phase, grain growth, oxidation, corrosion, erosion; and eventually become unsuitable. The lifetime attainable must be adequate for the purpose: this is clearly a matter to be decided by economics, since it may be better to select a cheaper, less durable material and replace the component at intervals.


The material selected must be fabricable into the desired component.


The problem of joining the component to the rest of the structure must be addressed: if it is to be welded, the properties of the weld metal and of the heat affected zones must be adequate.


The cost of the material, the fabrication into the component, the joining to the rest of the structure, and the periodic replacement, if that is required, must be acceptable.


A further requirement is that the various properties required should be reasonably consistent and predictable, so that one may design with confidence. An associated requirement is that the material should be "inspectable": by which is meant that acceptance testing and evaluation for possible defects during service is highly desirable.

In a combustion system in which heat is to be extracted by means of heat exchangers, the major critical requirements are usually those of the heat exchangers. It is common, although not necessarily essential, for the working fluid to be at elevated pressures, and the primary requirement is for adequate strength at the maximum expected temperature. In the U.S., these criteria are defined by the ASME Boiler and Pressure Vessel Codes. These define a maximum allowable stress for different materials, which is the maximum stress that can be used by the designer. This is the lowest of six stresses: (1)

1/4 of the specified minimum tensile strength at room temperature


1/4 of the tensile strength at the use temperature


2/3 of the specified minimum yield strength at room temperature


2/3 of the yield strength at the use temperature


100% of the stress required to produce a creep of 0.01% in 1000 h (a creep rate of approximately 2 x 1 0 " ^ s"1)


67% of the average stress to produce rupture at the end of 100,000 h, or 80% of the minimum stress for rupture as determined from extrapolated data, whichever is lower.

Figure 1(a) shows the application of these criteria to determine the maximum allowable stress for T22 (2-1/4 Cr - 1 Mo steel), and Figure 1(b) shows the maximum allowable stresses for a number of alloys. Both these figures are taken from "Combustion: Fossil Power Systems" prepared by Combustion Engineering^1-'. Only alloys for which Code allowable stresses have been defined can be used for these purposes in the United States. Other countries have similar specifications, although the form differs in detail.



Clearly, the heat exchanger alloys are subject to corrosion by the combustion environment on one side, and by the working fluid on the other. In the U.S., this is not a matter covered by the Codes, but the boiler manufacturers use their own criteria. Table 1 shows the maximum allowable use temperatures for several alloys according to Babcock and Wilcox^ ': the maximum use temperatures based on mechanical properties relate to specific assumptions about tube dimensions, and are to be taken only as a general indication. The strength criteria relate to the mean metal mid-section temperature; the oxidation limit relates to the mean metal surface temperature. The latter is typically 20-30°C greater than the former, depending on the wall thickness and the heat flux. Most of the steels referred to in the previous section are relatively stable at these temperatures. However, carbon steel and 1/2 Mo steel (Tl) will graphitize after long exposures at 400 and 470°C respectively, and these should be regarded as their maximum use temperatures· Construction of an AFBC Several previous papers have discussed the general form of an AFBC, and the mode of operation, but from a materials point of view a slightly different description is appropriate. Figure 2 illustrates the principal features. The bed itself is contained within a combustor enclosure. In the case of boilers, this will usually be waterwall construction, in which the walls consist of tubes linked by steel webs. In U.S. practice, it is usual for these tubes to be vertical, with outside diameters in the range 50-100 mm; the webs between them are typically 10 mm thick and 20-40 mm wide. Water, usually preheated, enters these tubes at the bottom from a header. The water may or may not boil in the walls: in a pulverized coal (pc) boiler, the water boils in the walls, and the steam/water mixture rises to the drum at the top of the boiler where the water and steam are separated. A similar steam drum is at the top of an AFBC boiler, but there may be some evaporator surface within the bed itself. In smaller boilers, the walls may be refractory lined, either at the bottom to contain the bed, or all the way up. Evaporation will take place in the in-bed heat exchanger; the combustion gas leaving the bed may pass through a firetube heat exchanger in which the water is on the outside of the tubes· In this paper, the problems of small boilers will not be specifically addressed. In larger AFBC boilers, the steam separated from the water is further heated: this is superheating. It would be usual for at least the final stages of superheating to be done in an in-bed heat exchanger, since the gas temperatures are considerably lower than in a pc boiler; a gas-pass superheater would thus have to have a significantly greater surface area than an exchanger having the same duty in a pc boiler. Since the finishing superheater is the most expensive heat exchanger in a boiler, it would appear to make most sense to take advantage of the excellent heat transfer in the bed. The air distributor is most commonly a plate with air nozzles in it. The nozzles usually are at the top of standpipes 5-10 cm long. The material below the nozzles is unfluidized, and this layer of material acts as a thermal insulator for the plate itself. Sometimes a refractory layer will be applied to the plate for the same reason. In some designs the distributor plate


J. Stringer

resembles a membrane wall, with water-cooled tubes joined by membranes: holes are drilled in the membrane to form the distributor· Other designs do not use a distributor plate at all, but use a group of nozzles manifolded into one or more air pipes; the material below the nozzles is again unfluidized. This is called a sparge distributor· From a materials durability point of view, the issues are the maximum temperature to be experienced by the distributor, the differential between this and the coldest part of the air plenum, and the duty cycle. A means must be provided to remove bed material and since it may on occasion be necessary to dump the bed, the total drain area has normally to be fairly large· Furthermore, to avoid blockage each individual drain point may need to be quite large, typically 10 cm diameter or more. This can present diffi­ culties in achieving uniform air distribution· In operation, when the bed drain is not being used, the drain will be filled with bed material up to a stop valve· Both the valve and the drain itself will experience considerable temperature variations, and must tolerate hot solid streams passing through them, There must also be means of feeding coal and sorbent. These may be mixed before injection through a common feed, or they may be fed separately. One or both may be injected at the bottom of the bed, or at the top. Currently, underbed feeding of coal is most common· In this case, the coal is crushed, typically to below 4 mm, and pneumatically injected· It is desirable for the air/coal ratio in this feed to be as low as possible, because the air that enters the bed this way must be subtracted from that blown through the distributor (for a given overall air/coal ratio)· A relatively high injection velocity is also desirable, since it promotes more uniform distribution of coal in the bed· It was believed by many designers that it was difficult to feed more than 1 m of bed with a single feed point; this is roughly equivalent to 0.3 MW(e) which is approximately 100 kg/h coal. If 15% of the air is fed through the solids feed system, this is equivalent to approximately 2 kg air/ 1 kg coal (for a 3% S Coal, Ca/S = 3) or 200 kg/h through a single nozzle. If the exit velocity is to be 30 m/s, this corresponds to a feed point diameter of approximately 50 mm. More recently, to simplify the coal feed problem, designers have been inclined to increase the area per feedpoint towards 2 m ; this would involve either increasing the velocity or increasing the feed port diameter. In fact, the tendency is to do the former, since the greater separation of the feed points implies a desire for greater jet penetration. Control and partial load operation may require that a number of coal feed ports are not used· Flow-back of bed material must be prevented, either by use of purge air blown continually through the feed ports, or by the use of some form of close-off. Overbed feeding is quite different. Relatively large coal (2-5 cm diameter) is distributed over the surface of the bed using a spreader stoker. The coal fines, the fine limestone, the exfoliated lime produced by calcination and the fine calcium sulfate produced by exfoliation or abrasion of the sulfated lime particles, will be elutriated from the bed. Good combustion efficiency and sorbent utilization require that this be captured and recirculated to the bed. The capturing is done in the primary cyclone. For a cyclone to operate efficiently, the gas velocity on entry is of the order of 30 m/s. The captured solids are then returned to the bed: since they are fines, they must be returned to the bottom of the bed. It is inconvenient to split the cyclone catch, so that if there is 1 primary cyclone per 5 MW(e) (which is probably reasonable) there will be one injection point per 15 m of bed. Depending on the recirculation ratio, the velocity at this point may also be high: to achieve reasonable mixing, velocities as high as 60 m/s are sometimes used.

FLUIDIZED BED BOILERS: DESIGN AND APPLICATION The cyclone catch i s hot, so that either i t must be cooled, or reinjected hot into the bed· There are advantages and disadvantages to both schemes. If i t i s cooled, one method i s to use a water-cooled screw to extract the dust from the bottom of the cyclone, and meter i t into the air transport system. Materials Problems in the AFBC The materials problems in the AFBC can be conveniently l i s t e d in terms of the system description given in the previous s e c t i o n . (1)

Corrosion of the in-bed heat exchanger and the in-bed support structure


Erosion of the in-bed heat exchanger, the water walls and other components


Erosion of feed l i n e s ,


Thermomechanical fatigue of the furnace enclosure and the a i r d i s t r i b u t o r plenum assembly


Mechanical fatigue of the in-bed heat exchanger and support structure


Erosion of the freeboard heat exchanger


Corrosion of the freeboard heat exchanger


Erosion of the primary cyclone

including the feed s p l i t t e r



Erosion and wear of the fines r e i n j e c t i o n system


Erosion /corrosion/thermal fatigue of the valves in the hot s o l i d s lines


Cold end acid corrosion

Some of these problems are reasonably similar to those experienced in pc systems. Because of the relatively low combustion temperature, the alkali release rate should be less in the AFBC than in the pc combustor. Furthermore, there should be no appreciable fusion of the ash. As a result, corrosion of the freeboard heat exchangers should not be an issue. Similarly, although the dust loading above the bed is high, it seems probable that the freeboard heat exchangers will follow the primary cyclones. In any event, it is unlikely that the dust will be more erosive than pc flyash, and it may well be less erosive. Consequently, if similar limits are used for the gas velocity through the freeboard and gas-pass heat exchangers as are used in the design of pc boilers (in the range 15-20 m/s) there should be no erosion. The cyclone presents more of a problem. It is probable that the primary cyclone will be refractory lined, in which case erosion shold be minimized, although the carry-over of spalled refractory must be a question. 30 m/s is a relatively high velocity for a particle-laden gas, and even with a refractory lining erosion should be monitored. In practice, erosion is not often reported even with all-metal cyclones: this may be because of the good flow patterns, and it may also be a consequence of the relatively short time highperformance cyclones have been run in AFBC systems. Other areas have so far received little systematic attention: thermo­ mechanical fatigue is a potential problem, but likely to be of more importance in large systems with appropriate heat recovery and a realistic operating


J. Stringer


cycle: there has been very little attention to this problem as yet. Feed line erosion, hot valve problems and so forth have received only cursory attention: some of them appear similar to problems in other well-known systems, and solutions may already be available. The three most important issues are numbers (10), (2), and (6) in the above list. In-bed corrosion has been studied in some detail, and there is a reasonably good understanding of the problem. In-bed erosion appears in some beds but not others: it appears therefore not to be an intrinsic problem for AFBC systems, but so far there is only very imperfect understanding of the factors involved, and little systematic research has been conducted. No reported instance of fatigue failure of in-bed structures has been reported, but it is recognized as significant potential problem in large units. These three issues will be addressed separately. In-bed Corrosion in AFBCs The possibility of in-bed sulfidation/oxidation corrosion was first suggested on theoretical grounds by Stringer and Ehrlich. The existence of the corrosion was demonstrated in a 2000-hour test conducted by the Coal Research Establishment (CRE) of the National Coal Board (NCB) under the sponsorship of EPRI. Since then, a detailed study of the mechanism has been conducted by CRE under joint EPRI/CRE sponsorshipJ ' ' and several papers have been published outlining the mechanism. Laboratory studies have been conducted by Stringer and Whittle, ' Akuezue, Stringer and Whittle, Mark, Stringer, Lin and Stevenson, ' Perkins, ' and Ficalora. ' On the basis of all this work, the following statements can be made. (1)

Above a metal temperature of approximately 600°C, an oxidation/ sulfidation corrosion process may occur.


This is induced by low local oxygen activities, which generate local high sulfur activities as a result of the dissociation of calcium sulfate.


The local low oxygen activities are detected by stabilized zirconia solid electrolyte probes. These indicate fluctuating oxgyen activity from close to 0.2 atm to as low as 10" atm. The frequency of the fluctuation is in the range 0.5-2.0 H z . ^ 1 4 " 1 8 '


The low oxygen activities appear to be associated with the emulsion phase, in which the coal particles are combusting: the high oxygen activités with bubbles.


The pattern and magnitude of the high and low oxygen activities vary from point to point in a given bed; the variations are different from bed to bed. However, the minimum oxygen activity is almost always close to 10" atm.


Calcium sulfate is necessary for the reaction to proceed. In the absence of a significant amount of calcium sulfate, either from the use of an acceptor, or from the coal ash itself, the sulfidation/ oxidation corrosion is absent.


The corrosion is not significantly affected by overall coal/oxygen ratio. However, some alloys which are sensitive to this form of attack may corrode more rapidly in substoichiometrie conditions.


The corrosion is not affected by the volatiles in the coal.


The corrosion is relatively insensitive to the sulfur content of the coal.


The corrosion is insensitive to changes in the coal ash chemistry (except for the presence of calcium, as indicated above) or the acceptor chemistry·


The corrosion is only reduced a little when the bed temperature is reduced from 900 to 850°C.


Above a metal temperature of 700°C or so, the corrosion rate of most alloys appears to be relatively insensitive to temperature.


Metal surfaces at temperatures above 400°C or so in fluidized beds are covered with a deposit composed of calcium sulfate, some calcium oxide, and components derived from the coal ash. The deposit is very compact: it is usually not possible to detect any porosity. It is well-bonded to the oxide on the metal. Its thickness may vary from 1 mm to less than 50 ym.


The corrosion does not appear to be related to the deposit in that the variation in corrosion behavior over the surface of a specimen is not related to the variations in the deposit thickness over that specimen, and variations which alter the deposit thickness have no effect on the corrosion. However, this is a matter of some debate.


In crevices, narrow enough that bed material cannot enter, conditions may become more severe than those in the bed itself, particularly under substoichiometric combustion conditions.


Nickel-based alloys are very sensitive to this attack, and may undergo catastrophic destruction, with formation of liquid nickel sulfides. High nickel alloys such as Incoloy 800 (45 Fe-33 Ni-21 Cr-0.4 Ti0.4 Λ1) are also sensitive, but there appears to be insufficient nickel for the reaction to become catastrophic. Cobalt-base alloys appear to be variable in their behavior: in general, they appear to be quite resistant (note that all tested have contained at least 22% Cr) but some severe localized corrosion has been reported, for example on Haynes 188 (Co-22 Ni-22 Cr-14 W-3.5 Fe-2 Mn-0.1 La). In general, iron-based alloys have exhibited the best resistance. For superheaters, with maximum metal temperatures of perhaps 650°C, Type 347 stainless steel appears to be the best of the materials tested, with Type 304 perhaps almost as good. Type 347 can be used at even higher temperatures, but close to bed temperature Type 304 is probably at the limit of its intrinsic oxidation resistance. Type 310 has excellent corrosion resistance, but it is prone to the precipi­ tation of brittle sigma phase. This can be avoided if the alloy chemistry is carefully controlled. A high-chromium Fecraly, GE2541 (Fe-25 Cr-4 Al-1 Y) has exhibited very good corrosion resistance, but its mechanical strength at elevated temperatures is very poor. It may be possible to use it as a cladding to protect a strong material. Lower chromium Fecralys appear not to be as resistant.


Welding presents problems. In a superheater, the colder end would normally be a low alloy ferritic steel, such as 2-1/4 Cr-1 Mo (T22), with the alloy changing to an austenitic steel such as Type 347 towards the hotter end. The weld metal for this would normally be a nickel-base alloy such as IN82, but this might be a little risky. The alternative is Type 309 austenitic steel (Fe-25 Cr-12 Ni) but this too has been less than wholly satisfactory.



J. Stringer


The early results suggested that the low alloy ferritic steels oxidized somewhat more rapidly than the same steels in air or a pc atmosphere· No sulfidation was detected. However, these results have been questioned because of uncertainties in the temperature ' Minchener et al. at measurement of tubes in the fluidized beds. CRE have been conducting a careful study of temperature measurements in fluidized beds, and the results should be available shortly. This is an issue of considerable technical significance.


Uncooled in-bed support materials are less of a problem than might be supposed, since the loading should not be high and a significant corrosion allowance can be added. Type 347 might be adequate, and alloys such as HK40 might also be satisfactory.

Mechanism of In-Bed Corrosion The corrosion process for the more highly-alloyed materials above approximately 600°C is sulfidation/oxidation. Beneath the initially-formed protective oxide (usually Cr 2 0 3 for the alloys used in these systems, less commonly A ^ O ^ ) sulfide particles appear in the alloy. Initially, for many of these alloys the sulfide particles are manganese-rich, and are small, randomly-distributed particles: these have no effect on the oxidation behavior. Later, the sulfides becomer larger and more numerous; they tend to be present along grain boundaries and other structural features of the alloy. These sulfides are chromium-rich. The chromium-rich sulfides then oxidize in situ to form internal Cr 2 0 3 . The metal matrix beneath the external scale is thus depleted in chromium, and if the protective oxide is lost for any reason (thermal cycling, erosion, the deposit falling off and taking the oxide with it) it will not be possible for the protective oxide to reform, and rapid oxidation ensues. The sulfur released by the internal oxidation of the sulfides does not appear to escape to the atmosphere, but is driven further into the alloy, reacting again with chromium, so that the band of internal oxides is always preceded by a band containing internal sulfides, and the rapid corrosion, once initiated, will continue to propagate. The sulfur activity may rise high enough for sulfidation of the base metal to take place: in the case of nickel-based alloys, this can result in catastrophic breakdown of the alloy, since a liquid eutectic forms between nickel and nickel sulfide at 637°C. Sulfides of the base metals are frequently observed between the non-protective oxide and the deposit on rapidly-corroding alloys. Figure 3 shows the extent of the corrosion for a number of alloys in several different tests, and Figure 4 illustrates the effect of process variables on the maximum corrosive penetration for two alloys: Incoloy 800H (Fe-32Ni19Cr-0.4A1-0.4Ti), which is sensitive to corrosion in the fbc environment, and Type 347H stainless steel (Fe-18Cr-10Ni-0.8Nb) which is relatively resistant. The observed corrosion morphologies imply that the local sulfur activity is in excess of 10"* atm., and probably in excess of 10~ atm., since this is the minimum required for iron or nickel sulfides to be formed. In the presence of pyrite from the coal, or indeed S0 2 from the oxidation of the pyrite, the existence of very low oxygen activities can result in high sulfur pressures. However, the fact that the corrosion does not appear in the absence of calcium sulfate implies that the crucial equilibrium is CaS0 4 = CaO + λ/2 S 2 + 3/2 0 2 On a thermodynamic phase stability diagram, (Figure 5) it can be seen that the oxygen and sulfur activity (partial pressures) are required to lie along a line representing the equilibrium between CaSO. and CaO, since the bed (and

FLUIDIZED BED BOILERS: DESIGN AND APPLICATION the deposit) normally contains large amounts of both of these phases. At 850 C, a maximum sulfur partial pressure of 10" atm. corresponds to an oxygen pressure of 10" atm. This corresponds to the triple point, at which a third solid phase, CaS, appears. The uniformity with which minimum oxygen partial pressures are reported suggests that the role of the calcium-containing species may be to buffer the system within the low oxygen activity region to these conditions. If this is so, the sulfur content of the coal should have essentially no influence since it is only necessary that the solid phases CaO and CaSCK should be present; the relative amounts do not matter. Equally, it appears that quite large variations in excess oxygen will have little effect; even prolonged substoichiometrie operation cannot alter the local chemistry until all the CaSO, has been reduced to CaS. However, in this case local environ­ ments may be dominated by the gas phase chemistry, in which case lower oxygen activities (at approximately the same sulfur activity) may appear, similar to a gasifier atmosphere. This is particularly likely to happen in crevices, where there is limited access for the solid species. The oxygen and sulfur pressures vary relatively little with temperature, and in particular the relationship with the oxide and sulfide stabilities of the elements in the alloys varies little with temperature. Oxygen probe studies have shown that the overall oxygen distribution varies from bed to bed, and can vary significantly with location in a given bed. The afbc at Battelle is markedly non-uniform in oxygen distribution, and it has been possible to show that the corrosion is significantly more severe in regions of the bed where the oxygen potential is mostly low as opposed to regions where the potential is mostly high.* Accordingly, if a bed can be designed or operated so that all the components at risk are in a region where the oxygen activity is mostly high, there should be little risk of corrosion. As yet, too little is known about the factors which determine the character­ istics of the oxygen distribution to allow predictions to be made. It seems clearly important to develop such an understanding as quickly as possible. It is clear from the reaction mechanism outlined above that sulfidation/ oxidation corrosion is a "breakaway" process: the initial rate is little different from that corresponding to protective oxidation, but after a certain time - the incubation period - the rate accelerates. This means that shortterm tests can give a misleadingly optimistic result, if the time of exposure is less than the incubation period, unless considerable care is taken in the interpretation of the results, and particularly of the development of the microstructure of the internal penetration. As yet, there is no method for predicting the length of the incubation period from short-term results, and it is clear that long-term testing is required, both to give engineering information on candidate materials of construction and to help in developing an understanding of the breakdown process itself. In-Bed Erosion in Fluidized Beds Erosion has been observed in a number of fluidized beds. When it appears, the local metal removal rate can be very rapid indeed. If a tube is penetrated, the resulting jet of steam will entrain bed material and cause considerable erosive damage to adjacent structures, so the problem is potentially very severe. However, the 6 ft x 6 ft AFBC of Babcock and Wilcox at Alliance has now run for over 10,000 hours under a variety of conditions, much of the time with fluidizing velocity in the 2-3 m/s range, and has exhibited no erosion at all, although measurements have been taken periodically to detect any loss of metal. Other beds have also run long times with no reported problems, and from this it can be deduced that erosion is not intrinsic to AFBCs. The



J. Stringer

following points summarize the possible mechanisms for those incidents of in-bed erosion that have been reported. (1)



Large fluidized bed ore roasters contain horizontal water-cooled(23) tubes (22) to control the bed temperature· Both Dorr-Oliver and Lurgiv ' note that erosion is absent if the fluidizing velocity is below 2 m/s or so· Above this velocity, erosion is observed; however, there are interesting differences: Lurgi observe erosion on the underside of the tubes, whereas Dorr-Oliver observe it on the upper surfaces· This implies that the erosion is caused by the long-range flow patterns— "gulf-streaming"—in the beds. The design figure of 2 m/s is interesting: erosion is relatively uncommon in AFBCs with fluidizing velocities of 1—1.5 m/s; most of the reported instances are for beds running at higher velocities, although as noted above a higher fluidizing velocity alone does not result in erosion . The Coal Utilization Research Laboratories (CURL) have run a circular shell boiler with bed diameter of 1 m for many thousands of hours under a wide variety of conditions. In the main, little or no erosion has been observed. However, when oil was burnt in a bed of coarse silica sand, severe erosion of the underside of the tubes was observed forming flats at the "four o'clock and eight o'clock" positions (six o'clock pointing downwards, into the gas flow). This implies that the erosiveness of the bed may be a factor, but Hoy points out that during other periods sand beds were used with no erosion: the reasons for this severe problem have not been elucidated. (25) De Coursin reported a case of severe erosion in the Fluidyne AFBC. An attempt was made to inject coal by blowing it into the top of the bed. To get good penetration of the coal, a relatively high velocity jet was used, and severe erosion of the in-bed tubes beneath the jet resulted. Other erosion incidents involving local jets have been reported: this is the single most frequent cause of problems. Jets from coal-feed ports, ash recirculation ports, even the jets from the holes in the air distributor nozzles have all produced serious erosion; the incidents will not be listed separately. The solutions are: keep the jet velocities as low as is consistent with other requirements; use deflector plates whenever possible, avoid having the jet pointing directly at any surface, but particularly the tubes. It is surprising how far a jet can penetrate in a fluidized bed. Often, jets have exit velocities of the order of 60 m/s; it would be better to restrict it to 30 m/s and achieve better mixing in some other way.


There have been some indications that erosion may be more severe in the splash zone at the top of the bed: the collapsing bubbles eject particles with relatively high velocities.


Erosion has been observed on the bends at the top of vertical tubes, for example in the Exxon PFBC, and at similar positions on sloping tubes, for example in the Georgetown AFBC. It appears that bubbles can move rapidly along such surfaces, carrying particles in their wake. At a bend, these particles may strike the metal, causing erosion. In the case of the Exxon problem, welding discs to the vertical tubes to disrupt the flow along the surface was effective in reducing the erosion, and similar techniques have been used elsewhere.


It has been observed that if larger particles (rocks from coarser overbed coal feeding, agglomeration as a result of bed temperature excursions) are present, there are manifestations of increased damage. For example, thermocouple sheaths are bent, and it has

FLUIDIZED BED BOILERS: DESIGN AND APPLICATION been suggested that this effect may be a factor in the erosion at the Georgetown AFBC. (2 ' (7)

The erosion observed in the Grimethorpe pfbc is different in character to any of these situations described above.* ' The uniformity of the wastage on tubes and the fact that it extends throughout a deep tube bundle are incompatible with any of the processes listed above. Instead, the process appears to be a cooperative bed motion which produces cavities beneath the tubes; these cavities then close abruptly, hammering the bed material against the underside of the tubes. This is very similar to the cavitation erosion observed in marine propellers for example. The local lateral motion of the bed following the closing of the cavities may be an additional factor. The cavities appear to be very extensive, running along a large fraction of the tube length. Since the erosion is related to a characteristic of the bed motion, the most effective way of combatting it is to modify the motion. Attempts to protect or armor the tubes are aimed at a symptom rather than the cause. This cooperative bed motion may have other dele­ terious consequences: the periodic pulsations may well lead to evential fatigue failures of structural components.

These brief summaries are not intended to be a complete survey of erosion in fluidized bed combustore: many other instances could be quoted. They serve merely to illustrate the possible factors: (a) (b) (c) (d) (e) (f) (g) (h)

Fluidizing velocity Erosiveness of bed material Long-range flow—"gulf-streaming" Jets Splash zone Fast bubbles on vertical or sloping tubes Rocks in the bed "Cavitation"-type attack from bed cooperative motion

Erosion Fundamentals There are several factors involved in erosion, some of which are understood, and some are not. These are summarized below. (a)

Effect of substrate ductility. Finnie suggested that ductile materials eroded by a cutting or plowing mechanism, whereas brittle materials eroded by brittle fracture. As a result, he was able to demonstrate that for ductile materials, te maximum erosion rate should be when the eroding particle impacted the surface at an angle of 30° or so; for brittle materials the maximum damage was at an impact angle of 90°. At room temperature, many observations have indicated that this is more or less true.


Effect of substrate and erodent hardness. At first, it appeared that erosion should decrease with increasing hardness of the substrate. However, several studies have shown that if steels are heat treated to widely different hardnesses, the erosion rate scarcely changes. It has also abeen suggested that soft particles will not erode surfaces harder than themselves. It appears that the cutting mechanism for erosion of ductile materials is only one possibility. A second mechanism is one or other variant of the damage accumulation models, in which the surface accumulates plastic damage until finally cracks appear parallel to the surface and flakes separate. It appears that cutting requires the particle to be harder than the substrate, but


J. Stringer that, provided there is a significant difference, the erosion is independent of the magnitude of the difference· In the case of damage accumulation, erosion can occur even if the erodent particle is softer than the substrate. (32) Effect of erodent shape. Raaskv suggested that the erosion in pc boilers was attributable to larger particles of quartz which had passed through the combustion zone without melting and, thus, preserved most of their angular shape. In processes such as grit blasting, it is known that angular particles produce more metal compared removal than rounded ones. Recently, Salid and Buckley the erosion of 1045 steel specimens, heat treated to produce a wide variation of microstructure and hardness, using glass beads and crushed glass as the erodents. It appeared that the crushed glass produced a cutting-type of wear and the beads a surface damage type; the rate of wear was typically an order of magnitude higher with the angular particles. There was no correlation with substrate hardness over the range R*47 to R A 68. Effect of velocity. Theory suggests that for ductile wear, the wear rate should vary with velocity raised to. a power n, where n is close to 2.3. Theory also suggests that for brittle wear, n should be closer to 3. Experiments frequently produce results of this order. There has been much discussion as to the existence of a threshold velocity, As noted above, boiler designers use a velocity criterion to avoid erosion, and in many systems it appears that there is no erosion at all below a certain velocity, and rapid erosion at a slightly hierher velocity. However, the experiments of Wright and Nagarajan show that, if a true threshold velocity exists, it is well below the design velocity, and probably below 1 m/s. Effect of temperature. The effect of temperature is not very great, and is somewhat obscure. Sometimes erosion increases with temperature and sometimes it decreases. It appear that the dependence on the angle of impact diminishes with increasing temperature. The issue is complicated by the participation of corrosion at elevated temperatures. Effect of particle size. If erosion is defined in the dimensionless form weight of metal removed/weight of impacting erodent, then it is independent of particle size for larger particles. This means that 1000 1-mg particles would produce the same damage as 1 1-g particle, provided all other conditions were the same. For small particles, the damage diminishes rapidly with particle size. There are two possible factors. The first is aerodynamic. As pointed out by Barkalow, Goebel and Pettit,* ' if particles entrained in a gas are directed towards a solid body, the gas stream lines flow around the body. Large particles follow ballistic paths and do not turn as the gas flow turns, striking the surface at the nominal impact angle. Very small particles remain entrained with the gas flow, and do not strike the surface at all. Intermediate sizes do strike the body, but the angles of impact are less than the nominal value. The second factor is connected with the possible existence of a genuine threshold, that small enough particles will be incapable of damaging the surface even if they strike it. This has not been demonstrated, but aerodynamic arguments seem inadequate to account for the rate of change observed in erosion. The variation of erosion with particle size has been discussed Cor the case of gas turbine erosion by Stringer and Drenker;^36 ' the size effect appears for particles below approximately 20pm, and erosion appears essentially to be absent for particles

FLUIDIZED BED BOILERS: DESIGN AND APPLICATION smaller than 3μ m or so. In practical terms, it is probably only the particles larger than 10μπι which contribute significantly to erosion in boiler systems. Erosivity of Fly Ash It is well known that the erosivity of fly ashes varies considerably. Figure 6 is taken from a study by Sverdrup^ ' of the erosivity of a large number of fly ashes collected from utility boilers burning a wide range of U.S. coals. Included in the study were bituminous, semi-bituminous and lignitic coals. As can be seen, there was an approximate correlation with total Si0 2 content, but with more than 55% SiC^, erosivity could vary by more than an order of magnitude with the same nominal silica content. Following Raask and others, it was believed that the amorphous, largely spherulitic, particles would not contribute to erosion, so instead the contents of crystalline quartz and mullite were determined. Figure 7 indicates that this may perhaps improve the overall correlation, but variations of an order of magnitude, at the same crystalline phase contents, were still observed. This is probably related to the size distribution of the erosive components, and perhaps their shape, but although the overall size distribution was measured, the size distributions of the individual components were not: this would be a tedious but not impossible exercise. Several laboratory tests of ash erosivity exist, mostly based on the same principle: the sample ash is blown through a grit blaster nozzle against a standard (usually steel) target: the weight loss of the target for a given weight of ash is the measure of the erosivity. The velocity of impact is seldom determined directly: the nozzle velocity of the carrier gas is usually of the order of 60 m/s. In most cases, the tests are conducted at room temperature. Sverdrup's apparatus is similar, but the particle velocities are determined directly using a laser doppler anemometer; the nominal angle of impact of the particle stream with the target surface can also be varied. Wright and Nagarajan's apparatus is more sophisticated: the particles are preheated and injected into an acceleration tube which collimates them: the velocity is measured in terms of the gas pressure drop across a nozzle which is calibrated beforehand. Again, the angle of impact can be varied; erosion loss is measured both as a weight loss and as a surface recession using a profilometer. The velocity can be varied over a fairly wide range, of the order of 10-80 m/s, and the temperature can be varied from room temperature up to 1000°C. In no case has the erosivity of fly ash measured in these laboratory tests been quantitatively correlated with boiler erosion, largely because of the lack of quantitative data from actual boilers. However, the boiler manufac­ turers believe that there is reasonable qualitative correlation. Erosion Resistance of Metals As noted above, there have been many laboratory erosion studies of different metals. However, it is often believed that one way of combatting an erosion problem is to select a construction material which has a better erosion resis­ tance, or to use an erosion resistant coating. Figure 8 shows data from Hansen:^ ' the erodent was AI2O3 with a mean particle size of 27μπι in a dry nitrogen carrier gas; the velocity was 170 m/s. The particle flow rate was 5 g/m, and the tests were continued for 3m. Two temperatures were studied: 20°C and 700°C. The impact angle was 90°; 30° would perhaps have given a better representation, but the relative effects are probably not affected. Several conclusions can be drawn from this figure. The first is that the erosion resistance of common structural materials do not vary by more than


J. Stringer


±20% at room temperature. The wear-resistant stellites do not give benefits in fine-particle erosive situations that they do in coarse particle erosion or sliding wear. This is because they rely on a relatively coarse distribution of hard carbide particles in a relatively soft matrix: in fine-particle erosion, the matrix is cut away between the carbide particles which are undermined and fall away. Mild steel is as good as anything else, except molybdenum and tungsten. At the higer temperature, most materials erode rather less, but the difference is seldom more than 30%. It can be concluded from this that a materials substitution is unlikely to be the answer to an erosion problem. Coatings offer more hope. Electroplated chromium is used to protect induced draft fans from erosion: in Sverdrup's tests, the erosion of 750μπι chromiumplated targets was 4-5 times less than Type 304 stainless steel, at virtually all impingement angles. Sverdrup has tested a wide range of potential coating materials: the results are summarized in Figure 9. Erosion-Corrosion Interactions It is clear that there are possible interactions between erosion and corrosion. For example, if a metal can form a relatively thick adherent oxide coat, this may confer some erosion resistance. However, if the oxide is brittle or poorly adherent, the erodent may knock it off, allowing the oxidation to occur in a rapid, non-protective fashion. Internal corrosion may weaken (or, less probably, strengthen) the surface layers of a metal, modifying its erosion response. If the erodent is plowing the surface, the corrodent may rapidly remove the protrusions developed. However, if the corrosion depends on the accumulation of a corrodent (e.g., a molten salt) at the metal surface, the erodent, by removing it, may reduce the corrosive attack. This possible interaction has been studied under high velocity, high tempera­ ture conditions appropriate to the gas turbine by Barkalow, Goebel and Pettit,* ' and under lower velocity conditions over a range of temperatures by Wright and Nagarajan.^34' Barkalow et al, examined the interaction of erosion and alkali-sulfate induced hot corrosion, and found that if the erosion was much more rapid than the potential corrosion, there was little effect on the erosion of the presence of the corrodent. Equally, if the corrosion was much more severe than the potential erosion, the presence of the erodent had little effect. However, if the two processes alone would have produced comparable damage, there was a marked interaction, with the damage resulting from the simultaneous presence of the corrodent and the erodent being much greater than the sum of the two acting separately. Wright and Nagarajan compared erosion in two gaseous atmospheres at 760°C: one, simulating a combustion gas, contained 3% oxygen with CC^, H 2 0 ' an(^ N 2 ' fc^e other was relatively high purity argon. The erodent was AI2O3 with mean particle size of 15um, the velocity was 43 m/s, and the impact angles were 30° and 90°. In general, the erosion rates were more rapid in the oxidizing atmosphere: for some alloys (Fe-25Cr-4A1-1Y, for example) the effect was small; for others (IN100, for example) the erosion rate was between 5 and 10 times as rapid. The effects appeared to be more marked at 90°, reducing the angular dependence of the erosion. In-Bed Erosion:

Concluding Remarks

The conditions within a fluidized bed do not fit the normal concept of particle erosion. The mean particle velocity within a fluidized bed is lower than the fluidizing velocity, and although there is a distribution of velocities, very few particles indeed will have velocities higher than (say) three times the fluidizing velocity; the fast particles will have a very short mean free path unless they are within a bubble, and then the time of flight

FLUIDIZED BED BOILERS: DESIGN AND APPLICATION will be short· Generally, the erosion damage produced by a particle impacting a metal surface at 6 m/s is negligible, and even though the particle loading is very high, the small concentration of ash particles close enough to a surface to hit it would seem largely to counteract that· It has been suggested^ 39 ' 40 ' that the wear process in a fluidized bed is somewhat different to independent particle erosion: instead, the particle will be "loaded" onto the surface by a block of other particles. There has been limited systematic study of erosion in fluidized beds: the major study is that by Wood and Woodford, ' but unfortunately this was confined to a fluidization regime different from the bubbling regime used in most currently operating beds· The important conclusions were that the erosion mode appeared to be a surface fatigue process rather than a cutting process, and that the particle hardness had only a secondary effect on the wear· Clearly, more study is required· Generally, the appearance of erosion in a bed can be regarded as a design or operational problem, rather than a materials problem· The use of pins and fins on tubes to limit erosion has been proposed, and appears to be successful in some cases. Within the range of materials that can be chosen for in-bed tubes, the erosion resistance does not vary very much· Coatings might be effective if a localized area is suffering damage: an aluminide coating was tried at Georgetown, but was not effective· A floating disc device has been used to prevent return of bed material down fuel ports or ash reinjection ports when they are not in use; and this is subject to considerable erosion when the feed is on, and also thermal shocks on the turning on or off of the feed as well as mechanical shocks against its retaining caqe. A variety of materials has been tested for this arduous application.*42* To date, a highdensity silicon carbide appears best· Mechanical Fatigue of In-Bed Tubes The forces on tubes were examined by Kennedy^ , amongst others. The net conclusion was the oscillating forces due to the passage of bubbles were relatively small, largely because the bubbles were small in comparison to the tube length, so the overall forces on the tube tened to balance out. However, Turner and Irving^ ' recently suggested that a cooperative motion, akin to the passage of a planar acoustic wave front through the bed, was possible. This may be the same as the process described in item (7) in the previous section. Under these conditions, voids open beneath the tubes extending over a large fraction of the tube length, and then close abruptly: accelerometers indicate a significant mechanical pulse on the tubes. From time to time thinner components (thermocouple sheaths, sampling tubes and the like) are distorted within the bed, demonstrating the existence of considerable mechanical forces. At times, the whole bed vibrates, and rigid supports are required. No detailed research has so far been conducted on mechanical fatigue within beds, but such work, if only to define the possible limits, would seem to be highly desirable. References 1.

"Combustion: Fossil Power Systems." Engineering, Windsor, Conn., 1981).


"Steam: 1975).


J. Stringer and S. Ehrlich. ASME Paper 76-WA/CD-4 (1976); see also J. Stringer, "Ash Deposits and Corrosion Due to Impurities in Combustion Gases," R. W. Bryers, ed., (Hemisphere Pubi. Co., Washington, D.C.,1977).

Its Generation and Use."

J. G. Singer, ed., (Combustion

(Babcock and Wilcox, Alliance, Ohio,



J. Stringer


NCB Coal Research Establishment, "Materials Problems in Fluidized Bed Combustion Systems," Final Report on EPRI Research Project RP388, Report CS-1449 (May 1980)


A. J. Minchener, et al. "Materials Problems in Fluidized Bed Combustion Systems," Interim Report on EPRI Resarch Project RP979-1, Report CS-1475 (August, 1980),


A. J. Minchener, et al. "Materials Problems in Fluidized Bed Combustion Systems," Final Report on EPRI Research Project RP979-11, Report CS-1853 (May 1981 ).


A. J. Minchener, et al. "Materials Problems in Fluidized Bed Combustion Systems," Final Report on EPRI Research Project RP979-11 (in press).


See for example J. Stringer, R. D. LaNauze, and E. A. Rogers. Proc. 5th Intl. Conf. on Fluidized Bed Combustion (U.S. DOE, Washington, D.C., 1977) 682. A. J. Minchener and J. Stringer. Corrosion-Erosion-Wear in Emerging Fossil Energy Systems," ed. A. V. Levy (NACE, Houston, Texas, 1982); J. Stringer, A. J. Minchener, and D. M. Lloyd, "Corrosion in Coal Conversion Systems," ed. B. Meadowcroft and M. I. Manning (Butterworth1s, London, 1983).


J. Stringer and D. P. Whittle. Proc. VGB Intl. Conf. on Corrosion and Deposition in Power Plants, Essen (1977).


H. C. Akuezue, J. Stringer, and D. P. Whittle, in preparation; H. C. Akuezue, "Calcium Sulfate-Induced Accelerated Corrosion," M. S. Thesis, Lawrence Berkeley Laboratory, Materials and Molecular Research Division, University of California; LBL 10286 (Dec. 1979).


K. Mark, J. Stringer, J. S. Lin, and D. A. Stevenson. Paper presented to the Spring Meeting of the Electrochemical Society, San Francisco, May 1983; to be published.


R. A. Perkins, personal communication.


P. J. Ficalora and J. H. DeVan. "Corrosion-Erosion-Wear in Emerging Fossil Energy Systems," ed. A. V. Levy (NACE, Houston, Texas, 1982) 476.


M. J. Cooke, A. J. B. Cutler, and E. Raask. (1972) 153.


J. Stringer and A. J. Minchener. Proc. 7th Int. Conf. on Fluidized Bed Combustion (U.S. Department of Energy, Washington, D.C., 1982) 1010.


D. P. Saari and R. J. Davis, ibid, 995.


C. F. Holt, A. A. Boiarsky, and H. E. Carlton. (1982).


M. A. Rocazella and I.G. Wright, paper presented to Spring Meeting of the Electrochemical Society, San Francisco, May 1983 (to be published).


I. G. Wright and A. J. Minchener.


W. R. Apblet, Jr., personal communication (see also ref. 4 ) .


P. L. Daniel, personal communication.


A. Leon, personal communication.

J. Inst. Fuel, _45_

ASME Paper No. 82-GT-225

ASME Paper 82-GT-226 (April 1982).

FLUIDIZED BED BOILERS: DESIGN AND APPLICATION L. Reh, personal communication. H. R. Hoy, personal communication, D. G. DeCoursin. Materials and Components in Fossil Energy Applications (U.S. DOE Newsletter DOE/FE-0053/34; Oct. 1981) 7. L. A. Ruth and M. S. Nutkis, in discussion of a paper by H. H. Krause, et al., J. Eng. Power 101 (1979) 1. Georgetown University; Quarterly Reports prepared for U.S. Department of Energy Contract No. DE-AC21-76ET10381. S. Ehrlich, personal communication. W. R. Apblett, Jr., personal communication. E. Carls, comments in discussion, 7th Intl. Conf. on Fluidized Bed Combustion (U.S. DOE, Washington, D.C., 1982). I. Finnie, Wear, J_9 (1972) 81. E. Raask, Paper in Proc. 5th International Conference on Erosion by Liquid and Solid Impact, (Cavendish Labortory, University of Cambridge, U.K., 1979). J. Salik and D. H. Buckley, NASA Technical Paper TP-1755 (1981). I. G. Wright and V. Nagarajan, Final Report to EPRI on Research Project RP979-8 (in press). R. H. Barkalow, J. G. Goebel and F. S. Pettit, Final Report on EPRI Research Project RP979-4, Report No. CS-14448 (May 1980). J. Stringer and S. Drenker, Proc. Amer. Power Conf., _43_ (1981) 943. E. F. Sverdrup et al., Final Report to EPRI on Research Project RP1649-4; Report No. CS-1979 (August 1981). J. S. Hansen, "Erosion: Prevention and Useful Applications" ed. W. F. Adler; ASTM Special Technical Publication 664 (American Society for Testing Materials, Philadelphia, 1979) 148. S. A. Jansson. "Corrosion-Erosion-Wear in Emerging Fossil Energy Systems," ed. A. V. Levy (NACE, Houston, Texas, 1982) 548. J. Stringer, paper presented to NACE/DOE/LBL Conf. on Corrosion-Erosion-Wear in Emerging Fossil Energy Systems, Berkeley, California (Jan. 1982); not submitted for publication in the Proceedings (see previous reference). R. T. Wood and D. A. Woodford. "Tube Erosion in Fluidized Beds," Final Report to New York State Energy Research Development Authority, Report No. 81-12 (Dec. 1980). D. R. Petrak.

Unpublished reports to EPRI on Project RP979-13.


J. Stringer


T. C. Kennedy. "A Study of Forces on Immersed Tubes in Fluidized Beds,' Topical Report to EPRI on Project RP718-2. (EPRI Report No. CS-1542, September 1980).


M. J. Turner and D. Irving. Proc. 7th Intl. Conf. on Fluidized Bed Combustion (U.S. DOE, Washington, D.C., 1982) 831.

Table 1 Maximum Use Temperatures for Several Steels


Maximum Use Temperature, °C Oxidation/Graphitization Criteria (metal surface)

Carbon Steel, SA 106

Strength Criteria (metal mid-section)



550 565 580 650

510 560 595 650

Austenitic Stainless Steels: Type 304H 760


Ferritic Alloy Steels: 1 /2CrJ/2Mo 11/4Cr-J/2Mo 3/4Cr-1 Mo 9 Cr-1 Mo


Material: 2% C M Mo Steel Specification: SA-213 Grade T 22

1 Carbon Steel SA-192 2 Carbon Steel SA-210 A-1 3 C-V2M0 SA-209T-1

20,000i 18,000| 16,000

14,000 -Maximum

\ i\ Allowable S t r e s s ' ^ i \


cô 12,000 S 10,0001 0.666 Yield S t r e n g t h - V \ V \ W

8000 0.25 Tensile Strength-SA 6000 "Stress for Rupture in 100,000 Hours 4000 Stress of Creep Rate Ji\ 2000 "of 0.01% in 1000 Hours ^ * 0

i ■ » ■ i ■ i I i j-JL 400 800 1200 0 Metal Temperature, F


Fig. 1(a). Use of ASME Boiler Code criteria to establish allowable stress for a 2%Cr-lMo steel.

200 400 600 800 1000 12001400 Metal Temperature, °F

Fig. 1(b). Effect of temperature on ASME Boiler Code allowable stresses for grades of steel tubing.

From "Combustion: Fossil Power Systems" ed. J. G. Singer (1981)


J. Stringer



Freeboard heat exchanger-vj Dense phase — "emulsion


In-bed heat exchanger ~ " v Ç f ^ ^ " X X * ? \ ^ Small bubbles near distributor, which agglomerate as they rise, forming large bubbles, which burst as they reach the bed surface, ejecting particles into the freeboard



Fines reinjection port Bubble capsAir plenum Materials issues • Corrosion of in-bed heat exchanger • Erosion of in-bed heat exchanger - Caused by jets (fines reinjection) -Caused by long-range bed currents -In splash zone -General erosion • Corrosion of freeboard heat exchanger • Erosion of freeboard heat exchanger • Erosion of cyclones • Erosion of fines transport equipment • Hot solids valve erosion Figure 2. Schematic of an Atmospheric Pressure Fluidi zed Bed Combustor, Highlighting the Materials Issues



Alloy Type 304 SS

Program ORNL


Type 310 SS



Temperature CC)


Scale thickness

870 870 870 760 760 613 631 655


870 870 870

Type 309 SS


880 880

Type 347 SS


610 612 615 641 642 880 880 650 760 760

Nitronic 50


540 540 650 650



540 540 650 650

Manaurite 36X




900 880

Hk 40

Corrosive penetration

860 510 570 546 554 562 567 586 613

Type 316 SS


Time (h)

2000 2000

p L_l 0


I 100



150 200 (μτη)

Rgure 3. A comparison of the reported corrosion behavior of alloys in several tests ( 6 ) .



250 300

J. Stringer



Type 347 HSS



540650 760 840900 540650 760 840900 Nominal Metal Temperature (°C) Figure 4(a). Effect of Process Variables on Corrosive Penetration in AFBC (8) CORROSIVE PENETRATION IN FLUIDIZED BEDS Effect of Coal Incoloy 800 H

Type 347 HSS Illinois No. 6 ash sintering temperature 850°C

Eccles ash sintering temperature 1100°C

Illinois No. 6 char

540 650 760 840 900 540 650 760 840 900 Nominal Metal Temperature (°C) Figure 4(b). Effect of Process Variables on Corrosive Penetration in AFBC (8).




Type 347 HSS



s i C

2. .1 g δ

Penrith limestone (pure)

50 0

None Oxide penetration only, in all other tests sulfide/oxide penetration was found

50 0 100







> I i

M _




E 100




B ■

Wenlock limestone containing 27% Si0 2

50 0


l 540■ 6501 7601 840 1II —650 ■ 760 ■ 840 ■ 900 B1 900 540 Nominal Metal Temperature (°C)

Figure 4 ( c ) .

Effect of Process Variables on Corrosive Penetration in AFBC ( 8 ) . CORROSIVE PENETRATION IN FLUIDIZED BEDS Effect of Excess Air Incoloy 800 H

100 c o ^


Type 347 HSS * No specimen

50 |— 0 50 0

Excess Air

x " B B BI ' "




540 650 760 840 900 540 650 760 840 900 Nominal Metal Temperature (°C)

Figure 4(d). Effect of Process Variables on Corrosive Penetration in AFBC (8).


J. Stringer CORROSIVE PENETRATION IN FLUIDIZED BEDS Effect of Coal Chlorine Content Incoloy 800 H

Type 347 HSS

100 ^ c


50 0





il Φ


Illinois No. 6 0.1%


Welbeck 0.5% CI

Û. Φ




p fc








_ M * I *

150 E D E 100




Illinois No. 6 no acceptor Oxide penetration only, in all other tests sulfide/oxide penetration was found

Welbeck no acceptor


540 650 760 840 900 540 650 760 840 900 Nominal Metal Temperature (°C) Figure 4(e). Effect of Process Variables on Corrosive Penetration in AFBC (8). -2f


-23 -22 -2I -20 -I9 -I8 -I7 -I6 -I5 -I4 -I3 -I2 -Il -IO -9 -8 Figure 5. The thermodynamic phase stability diagram for the Ca-O-S system at three temperatures.


40 30 SK)2 Content of Fly Ash (wt%)

Figure 6. Erosion rate versus SiCL content by chemical analysis of bituminous fly ashes. From Sverdrup (37).




I L Pwr. & Lgt. «15 (Havana)

M ft /


J3Q Texas Utilities (Monticello)




'S E

S.Calf.Ed. · / ( Mohave) /


/Lansing (Erickson)

/ / /

· 143 Utah Pwr. & Lgt. (NaughtonNo.3)



i 2



A150 Penelec ( Keystone )

- Utah Pwr. & Lgt. (Huntington »133 0.01 Canyon) 132 Tx. Util.(Big B r t m n j / ^


/ / / y S / ^ r 1 •40 I n. d ,,? w r·

(Homer City)

·53 LDulsvilleG &E 1 Cane Run)

· 93 Ariz. Pub. Service (Four Corners)

Figure 7. Erosion rate versus crystalline phase content for a number of fly ashes. From Sverdrup (37).



J. Stringer T . - 6 A I - 4V Hoynes 9 3 25Cr iron Haynes Stelhte 6K Haynes Stellite 3 Haynes Stellite 6Θ 3 0 4 stainless s teel 316 sta inless steel Haynes 188 Haynes 25 4 3 0 sta inless steel HK-40 Inconel 6 0 0 RA-330 Incoloy 8 0 0 H Beta ΙΠ Ti Incoloy 8 0 0 RA 3 3 3 Inconel 671 Mild ste el Mo W 0.4 0.6 0.8 1.0 1.2 1.4 RELATIVE EROSION FACTOR


Figure 8. Comparative erosion resistance of several metals and alloys. Erodent A ^ O ^ mean particle size 27 ym, velocity 170m/s, dry nitrogen carrier. From Hansen (38). Gas Velocity 700 ft/sec Target Angle 40 e Target toNozzle Spacing 0.5 Inch Nozzle Diameter 0.0625 Inch

304 Stainless Steel Run 19

Inconel 625 Run 24 Multlmet Run S

40 \

1075 Annealed Run 25 Hastelloy C Run 28


Hastelloy C 276 Run 23

TIC Coated LW- IN30 D Gun Coating on 1075 Steel Run 31 LW-1N30D Gun Coaling on Alloy 25 Run 3β

300 Grams Ash Delivered

Figure 9. Ash abrader test results showing comparative erosion resistance of various protective systems to high velocity bombardment at 40°. From Sverdrup (37).