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THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS
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, MATERIALS ISSUES FOR HIGH-TEMPERATURE COMPONENTS
IN INDIRECTLY-FIRED CYCLES
Ian G. Wright
111111111111111111111
Oak Ridge National Laboratory
1 Bethel Valley Road
Oak Ridge, Tennessee 37831
John Stringer
Electric Power Research Institute
3412 Hillview Avenue
Palo Alto, California 94303
ABSTRACT
Indirectly-fired cycles provide one means of using a fuel other
than natural gas or distillates of various purities to generate power
using a gas turbine. In a closed cycle, the fuel typically is used to
heat a clean working fluid which is then expanded through a gas
turbine, after which it is cooled and recompressed before being
recirculated through the heating circuit. In an open cycle, the heated
working fluid (usually air) is exhausted to the atmosphere after
expansion in the turbine and passage through heat recovery devices.
In both cases, the temperature of the working fluid may be boosted
before entry to the turbine by supplementary firing of a premium fuel
such as natural gas in a topping combustor. A major advantage of
such indirectly-fired cycles is that the concerns arising from the use
of a dirty fuel in other advanced cycles are confined to the fireside
surfaces of the heat exchange equipment, whereas the gas turbine is
exposed to a relatively benign environment. One limitation of such
systems is that the emissions problems are the same as for a
conventional coal-fired boiler although, on an power outputnormalized basis, the emissions from an indirectly-fired cycle may be
lower. The requirements of the potential candidate materials for the
various components in the circuit are discussed, and the critical
issues for each are identified.
(2250°F) with a additional ceramic heat exchanger.
Current
programs involving indirectly-fired gas turbine cycles are aimed at
high cycle efficiencies, of the order of 47 percent based on the higher
heating value (HHV) of the fuel, and involve open cycle systems in
which air is heated to 760°C (I400°F) in a metallic heat exchanger,
followed by further heating to 982°C (1800°F) in a natural gas-fired
ceramic heat exchanger [Klan, 1993, 1994; Robson, et a]., 1996]. A
variant of this approach is where part of the coal is pyrolyzed to
produce the fuel gas used to fire the ceramic heat exchanger or the
turbine; in that case the air entering the turbine is heated to 1288°C
(2350° F). A further program envisions using a coal-fired ceramic
heat exchanger for the whole duty of heating air to I200°C (2I92 ° F)
[LaHaye, et al., 19951.
Successful
implementation
of
indirectly-fired
cycle
technologies will require the development of a durable coal-fired
heat exchanger capable of heating the working fluid to very high
temperatures, in addition to adapting a gas turbine for this particular
duty. This paper discusses the materials issues associated with the
heat exchanger section of indirectly-fired cycles.
GENERAL DESCRIPTION OF MAJOR COMPONENTS
INTRODUCTION
Interest in increasing the efficiency of coal-fired power plants
has led to the examination of alternatives to the steam boilerRankine cycle systems, for which increases in efficiency have been
limited by the slow progress in improving the ability to handle
steam at temperatures much in excess of 565°C (I050°F), as well as
by the unavailability of easily accessible sources of naturallyoccurring low-temperature cooling water. Indirect-firing of gas
turbines in open or closed cycles is one approach to linking the
highe efficiencies possible via the Brayton cycle with coal asthe
fuel. An experimental program in the 1980's [Campbell and Lee,
1982) demonstrated a coal-fired, low-emissions heat exchanger
(fluidized-bed combustor) capable of heating air to 843°C (1550°F)
in a metallic heat exchanger, and to 954°C (1750°F) or I232°C
Figure 1 shows a schematic diagram of a hypothetical coal-fired
open-cycle system constructed so as to illustrate the major
components of interest.
Considerable license has been taken in suggesting locations
and interconnections of components that are not necessarily those
that would maximize the system efficiency. Also, no provision has
been made for the interconnections needed to facilitate part-load
operation. In addition, no indication has been offered of any energy
extraction from the flue gas stream from the coal-fired heat exchanger
although, in practice, some fonn of heat recovery boiler and steam
turbine system probably would be used. The main heat exchanger
has been deliberately drawn to resemble the configuration of a
modem coal-fired steam boiler, so that the location of the
components of interest has some practical logic, and the
Presented at the international Gas Turbine & Aeroengine Congress & Exhibition
Orlando, Florida —June 2–June 5,1997
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environments to which specific components are exposed at a given
location are well characterized. Given the importance of the spatial
arrangement of the heat transfer surfaces with respect to the flame in
controlling heat flux, fouling, erosion and corrosion, starting from a
known configuration should allow a rational analysis of the
probable implications of changes in size and location of the heat
exchanger components of interest.
use of a modified combustor on the gas turbine, with the associated
increase in capital cost.
Figure 2 shows a version of the hypothetical plant for operation
in closed-cycle mode. The major differences are that the working
fluid and the system pressure can be chosen to optimize the heat
transfer properties (although in this paper the working fluid is
assumed to be air). Also, the turbine in a closed cycle is used simply
to expand the gas and does not involve any combustion. Following
expansion in the closed-cycle turbine, the working fluid is further
cooled in a recuperator (which may preheat the combustion air, or
may simply reject heat to a cooling tower) before re-entering the
compressor. A complicating factor is that careful control of the fluid
flow to the turbine compressor must be exercised in a closed cycle to
ensure its proper functioning; this requires an ability to adjust the
temperature of the inlet fluid, and emphasizes the importance of the
functioning of the recuperator. Another difference is that in a closed
cycle supplementary firing of a premium fuel to increase the Rrr
requires a separate topping combustor and heat exchanger to heat the
working fluid.
Figure 1. Schematic Elevation of a Coal-Fired OpenCycle Gas Turbine System
The working fluid is air which is compressed in the turbinedriven compressor before it enters the main body of the
'conventional' heat exchanger (conventional in the sense that it
uses metallic alloys). There the air is heated to a temperature
consistent with the maximum practical operating temperature of the
metallic materials used before passing to the advanced, hightemperature heat exchanger which, in this case, is taken to be
ceramic. The conventional heat exchanger is located in the furnace
zone for the same reasons as in a coal-fired boiler, to minimize
fireside corrosion problems by arranging for the coolest metal
surfaces to contact the highest-temperature gas, as well as to provide
an efficient means of containing the flame. The maximum working
fluid temperature is attained in the ceramic heat exchanger, which
may be located at the top of the furnace to permit radiant heating
while avoiding direct contact with molten coal ash or slag, or may
be sited deeper in the furnace zone to take advantage of increased
radiant heat flux from close proximity to the fireball, if contact with
molten ash or slag is acceptable. After expansion in the turbine, the
working fluid may be further used as a heat source to raise steam in a
heat recovery steam generator or, as suggested in Fig. 1, to preheat
the air entering the conventional heat exchanger before being used
as combustion air for firing the heat exchanger, and then exhausted to
the atmosphere.
Figure 2. Schematic Elevation of a Coal-Fired
Closed-Cycle Gas Turbine System
The differences in heat transfer properties can lead to a
significantly smaller heating surface requirement for a closed cycle
compared to an open cycle, for the same heat absorption/power
output. Campbell and Lee (1982) indicate that from a cycle
performance viewpoint it is desirable to limit the pressure drop
across the heat exchanger to 3 to 5 percent of the inlet pressure.
Since it is important that the wall temperature of the fired surfaces is
maintained as close as possible to the working fluid temperature,
especially in the high heat flux regions of the radiant surface where
the flame temperature exceeds the melting temperature of the alloys
used, matching the heat flux to the coolant mass flow at the lower
inlet pressures of open cycles is a non-trivial problem. Further,
close control of the temperatures of the radiant heat transfer surface
leads to the need for good flame stability, as well as avoidance of
spiking from slagging and deslagging events. Hence, the ability of a
closed cycle heat exchanger to operate in a high heat flux zones
might also result in a simpler design. Working fluid considerations
are further discussed in Appendix I.
Since the power and efficiency of the Brayton cycle is a direct
function of the turbine rotor inlet temperature (Eli), it is desirable to
maximize this temperature. For modern, high-efficiency gas turbines,
the FUT is typically of the order of I288°C (2350°F), and is
projected to rise to 1450°C (2642°F) in advanced land-based gas
turbines. Since these very high temperatures may not be readily
achievable using the heat exchanger alone, at least in the firstgeneration units, one option is to fire a small amount of a
supplemental premium fuel in the turbine to further raise the
temperature of the working fluid. This approach would entail the
The cost-benefit arguments concerning the use of some level of
supplementary firing are addressed in the U.S. Department of
2
Energy's High Performance Power Systems program [Klara, 1994],
in which air is heated to 8I6°C (I500°F) in a coal-fired heat
exchanger and then further heated to 1260°C (2300°F) by natural
gas. A major advantage of the topping combustor is that it allows
the ceramic heat exchanger to be operated at a maximum temperature
commensurate with the strength and corrosion capabilities of the
material used, and also provides the ability to change that
temperature as experience accumulates. Obviously, there is a tradeoff between the cost of the extra components required and the
premium fuel used with the increased reliability of the system. In the
limit, the topping combustor might be absent, or may totally replace
the advanced heat exchanger in Figs 1 and 2. Thus there is a strong
incentive for increasing the reliable temperature capability of the
coal-fired heat exchanger to decrease the requirement for
supplemental firing in the topping combustor.
Figure 3 from Campbell et al. (1980) shows the relationship of
Rif to the overall efficiency of a closed-cycle system using a coalfired heat exchanger. System efficiencies in the range of 40 to 47
percent (HHV) for a simple, power-only 'closed-cycle gas turbine
system are suggested for a RIT of the order of that of state-of-the-art
power generation gas turbines, that is 1288°C (2350°F). Note that
they set the maximum temperature of the air exiting the metallic heat
exchanger at around 870 °C (1600°F), and used a 'primary' ceramic
heater for heating the air from 870 to 982°C (1600 to I800°F).
The maximum duty for the heat exchanger materials is to contain
the working fluid at the maximum temperature and pressure
consistent with the turbine design, while providing sufficient
resistance to the working fluid and fireside environments to give an
. acceptable lifetime for the heat exchanger. For an open-cycle system
the maximum system pressure will correspond to the design pressure
ratio of the turbine which, for modem power generation gas turbines
ranges from approximately 10 to 30; hence, the maximum pressure
would be approximately 1.1 to 3.1 MPa (162 to 456 psia). In a
closed cycle, the system operating pressure is determined by the
design and operating characteristics of the compressor (Lee, et al,
1980); small gas turbines (<50 MW) typically require pressures in
the range 0.4 to 3.4 MPa (60 to 500 psia) for good compressor
efficiency, whereas larger machines (50 to 350 MW) operate best at
pressures from 3.4 to 6.9 MPa (500 to 1,000 psia). At the higher
pressures, the problem of ensuring leak tightness of joints is
increased, which requires extra care in design and initial assembly,
as well as in materials selection to ensure that the potential effects of
aging and corrosion are assessed. This is especially important for a
closed cycle using a high-value working fluid.
For a heat exchanger surface in the form of a tube, the hoop stress
corresponding to these system pressures can be calculated from:
SA = d.P12t, where d is the tube diameter, P is the internal pressure,
and t is the wall thickness. Assuming a tube 5.1 an (2 in.) in
diameter with a wall thickness of 6.4 mm (0.25 in.), the hoop stress
ranges from 4.5 to 12.4 MPa (0.65 to 1.8 ksi) for an open cycle
system, to 1.7 to 13.8 MPa (0.24 to 2 ksi) for a closed cycle system
based on a small turbine, and from 13.8 to 27.6 MPa ( 2 to 4 Icsi) for a
large turbine closed-cycle system. Applying a design safety factor of
1.5, the required creep strength in the hoop direction at temperature
would range from approximately 6.8 to 18.6 MPa (0.97 to 2.7 Icsi) for
most open cycle turbine conditions, 2.6 to 20.7 MPa (0.36 to 3) for
closed cycle turbines up to 50 MW, and from 20.7 to 41.4 NfPa (3 to
6 hi) for closed cycles from 50 to 350 MW. Hence, a hoop strength
of 21 MPa (3 ksi) is sufficient for operation of the heat exchanger
with any current gas turbine in an open cycle, and up to 50 MW in
closed cycle, whereas larger closed cycle machines (50 to 350 MW)
require tubing hoop strengths up to approximately 42 MPa (6 hi).
(Si,)
MATERIALS SELECTION CONSIDERATIONS
In practice, the form of the heat exchanger surfaces may be
considerably different from those in a steam boiler, and there will be
other differences such as the fact that the lower-temperature portion
of the conventional heat exchanger will be significantly smaller,
since there is no requirement to provide any latent heat of
vaporization. Nevertheless, the fireside corrosion problems will be
similar to those in current coal-fired boilers, depending on the local
gas and metal temperatures. The gas turbine materials problems are
essentially the same, or less severe, than in a conventionally-fired
turbine using a clean fuel.
Conventional
Metallic
S 732°C/1350°F1
Heat
Exchanger
(Tm
In current coal-fired steam boilers, the maximum operating
temperature experienced by pressure boundaries does not exceed
approximately 732°C (I350°F). The application of heat transfer
surface materials in steam boilers is governed by the ASME Boiler
and Pressure Vessel (BPV) Code. This requirement has resulted in
the use of relatively few alloys in this application; since Codequalification requires the generation of long-term creep strength
data, the qualification process for new alloys is relatively lengthy
(and expensive).
(538)
(816)
(1093)
It is expected that the alloy selection criteria for steam boilers
for metal temperatures (T.) up to approximately 732°C (1350°F)
would be used for any coal-fired heat exchanger, since the internal
steam pressures are of the same order (up to 25 MPa/3.6 ksia) as the
maximum pressures contemplated for the working fluid in open- and
closed-cycle systems, there is a large database of experience of alloy
performance, and the construction techniques and practices have
been well proven. Hence, the selection of materials for the heat
exchanger surfaces of the conventional metallic heat exchanger in the
radiant furnace zone would start with seamless carbon steel or T-I I
(Fe- I.25Cr-0.5Mo) tubes; the maximum recommended use
(1371)
Rotor inlet Temperature 'Fr)
Figure 3. Variation of Efficiency of a Closed-Cycle
Gas Turbine With Turbine Inlet Temperature (after
Lee et al., 1980)
3
High-Temperature Metallic Heat Exchanger
(Tin
a 732°C/1 350°F)
Swindeman and Marriott (1994) and Ruth and Birlcs (1995)
recently assessed alloy properties required for power generation
plants as a function of increasing plant efficiency. The number of
alloys identified as having potential for application at the higher
temperatures decreased rapidly above approximately 815°C
(I500°F). The combination of properties desired in alloys for use in
the high-temperature metallic heat exchanger is difficult to achieve,
since the alloys must exhibit:
• increased creep strength,
• resistance to corrosion from the fireside environment,
• corrosion resistance to the working fluid environment, and
• ability to be fabricated by reasonable techniques.
temperature of these alloys in coal-fired boilers [Dooley and
McNaughton, 1996] are 454 to 510°C (850 to 950°F) and 552 to
566°C (1025 to 1050 °F), respectively. The water/steam-cooled
furnace wall tubes in steam boilers operate at metal wall temperatures
sufficiently low that they will provide service lives of >2 x 10 5
hours under all conditions except when direct flame impingement
occurs. The use temperatures of these tubes are corrosion-limited,
not creep-limited. However, the forms of fireside corrosion
experienced by low-alloy steel tubes in this location in coal-fired
systems are well documented for the coal types used in the U.S.
[Stringer and Wright, 1995], and guidelines for remedial measures
are available [Dooley and McNaughton, 1996].
Given that air as a working fluid will not have the heat
absorption capacity of steam-water mixtures (no latent heat buffer)
that allows excursions in heat flux without significantly increasing
tube wall temperature, it may be necessary to specify an alloy with
increased resistance to high-temperature oxidation, especially for the
high heat flux areas. Alloys for use in this location will be the same
as those considered for advanced steam condition plants (steam at 33
MPa/4.8 ksi and 625°C/I157°F) [Blum, 1994] and include 1-22 (Fe2.25Cr-IMo; 580 to 602°C/I075 to 1115°F) or modified versions of
this alloy, T-5 (Fe-5Cr-0.5Mo), and possibly 1-91 (Fe-9Cr-IMo) if
the increased cost can be justified by, for instance, the use of thinner
Creep Strength
Alloys approved or which are being
Wrought Alloys:
considered for approval (BPV Code Section VIII, Division 1) for
service at temperatures at or above 732°C (1350°F) are included in
Table I. Following Ruth and Birks (1995), Fig. 4 plots the
allowable stress as a function of temperature for some of these alloys
for the temperature range of interest; the horizontal shaded boxes
indicate the requirements for the two regimes of interest: open cycle
(all turbine sizes) and closed cycle <50 MW, and closed cycle >50
MW. Note that the requirements for small closed cycle systems are
equivalent to those for open cycle systems. The alloy with the
highest allowable stress at 899°C (I650°F) is Alloy 230. with
Inconel 617 a close second. These data indicate that this alloy
would be useable in an open cycle or small turbine closed cycle
system up to at least 860 °C (1580°F), and up to at least 780°C
(1436°F) in a closed-cycle heat exchanger for a large turbine. These
alloys represent the maximum practical performance for wrought
alloys strengthened by the conventional routes of solid solution
alloying and carbide or other precipitation.
The maximum temperature duty in steam boilers typically is
handled using type-300 series stainless steels such as 347H
(704°C; 1300°F), 304H (704-760°C; 1300-1400°F); or 321H (760815°C; 1400-1500°F) in the outlet sections of the superheaters;
these metal temperatures also are lower than those specified by the
BPV Code (except for type 32IH) because of fireside corrosion
considerations. For extremely corrosive duty, type 310 clad on a
higher-strength steel or Inconel 671 clad on lncoloy 80011 is used.
The compositions of these alloys are given in Table 1.
Table 1. Compositions of Candidate Alloys for the High-Temperature Heat Exchanger
Alloy
304H
321
34711
80011
NF709
I 253MA
556
HR6W
31011
HR3C
I 230
Inconel 617
HRI60
Inconel 671
Haynes 188
MaxU
se
T°C•
732
843
899
899
732
899
899
816
Chemical Composition, weight percent
Fe
Bal.
Bal.
Bal.
Ni
8-12
8-12
9-13
Bal.
Bal.
Bal.
Bal.
30-35
25
10-12
20
Bal.
Bal.
Bal.
1.3
0.0
3.5
43
20
20
Bal.
Bal.
Bal.
Bal.
20-24
3
Cr
18-20
18-20
17-19
Al
0.5
18
19-23
20
20-22
22
0.28
12
27-33
23
25
25
22
22
26-30
Co
Ti
Nb
Mo
10xC
10x
C
0.4
1.5
3
0.1
0.2
0.5
1
0.1
0.15
0.6
0.05
1
0.2
0.02
0.3
0.2-0.8
1.4-2.0
0.3
Other
1.2
9
I
0.4
0.2
2.4-3.0
0.2-0.5
21-23
N
0.03-0.08 Ce
2.5W, 0.6Ta,
0.02Zr
6W
0.2N,
0.25N
0.4
48
Bal.
C
0.08
0.08
0.08
0.1
0.5
0.25
0.4
1.2
Si
0.5
0.75
0.75
0.15
14W, 0.3La
0.05
0.07
0.1
1W, 1 Mn
*BPV Code data
4
13-15W, 0.02-0.12La
0.02La,
Creep rupture data in the longitudinal direction for Inconel
MA956, calculated from 213 of the reported 10,000 hour rupture
strength [McColvin and Smith, 1987], are shown as a curve in Fig. 4.
Data from a different source [Hurst, 19961 for longitudinal as well as
circumferential creep strength at I100°C (2012°F) for the three ODS
FeCrAl alloys in Table 2 were used in Fig. 4 to indicate the range of
properties. Using only the circumferential data as representative of
the capabilities of these alloys in a high-temperature heat exchanger,
it is projected that the best of these alloys could be used up to at
least 935°C (17)5°F) for all open-cycle sizes considered and small
turbine (<50 MW) closed systems, and up to I100°C for lower
pressure-ratio turbines in an open cycle system. Obviously these
projections only suggest the potential for these alloys, and more
extensive data are needed. Nevertheless, the very conservative
approach of using a 50 percent safety factor plus the 33 percent
margin associated with the BPV criteria lends some credibility to
the data shown in Fig. 4.
Metal Tang:ream •C
Figure 4. Allowable Stresses for Qualified and
Pending ASME Boiler and Pressure Vessel Code
High-Temperature Alloys (data for MA956 and other
ODS alloys were based on 2/3 10,000 hr creep
rupture strength)
Oxide Dispersion-Strengthened Alloys: There have been
significant developments in recent years in alloys strengthened by
an oxide dispersion (ODS) which can provide creep strength up to
approximately 90 percent of the alloy melting temperature. Table 2
shows the chemical compositions of three commercial iron-based
ODS alloys; the melting point of these alloys is 1480°C (2696°F).
Table 2. Compositions of Commercial Ferritic Oxide
Dispersion-Strengthened Alloys
Fe
Bal.
Composition, weight percent
Cr
Al
Ti
Mo
4.5
20.0
0.5
Bal.
Bal.
16.5
20.0
Alloy
Inconel
MA956
0DM751
PM2000
4.5
5.5
1.5
0.6
0.5
Y203
0.5
0.5
0.5
The ferritic ODS alloys based on Fe-Cr-Al have the potential for
application at higher temperatures than the modified
conventionally-strengthened alloys in which the strengthening
mechanisms degrade as the precipitated phases become less stable
with increasing temperature. Although the properties of this class of
alloys appear very promising, data are available usually for only
relatively short times, and none these alloys has been BPV Codequalified. There is a strong anisotropy in the creep strength of these
alloys. Since the microstructuml features associated with the
strengthening mechanism are usually aligned in the longitudinal
(axial) direction, maximum creep strength is in this direction.
Progress in processing and microstructural control has resulted in
increases in the circumferential creep strength, although it is still
significantly lower than the strength in the axial direction.
Unfortunately, the BPV Code does not consider the effects of
corrosion or of fatigue on the alloy strength limitations, and the
design engineer must decide his own allowances. Given a
reasonable understanding of the corrosion behavior of these alloy
types, a corrosion allowance factor can be produced if corrosion
information is available for the specific environment of interest.
Similarly, although there is no unified methodology for assessing
the effects of fatigue on tube life, boiler makers have their own
criteria. It appears that specific testing in the temperature and
environmental regimes of interest will be required to provide some
guidance in these areas.
Fabrication. In general, the wrought, high-temperature
alloys such as Alloy 230 and Inconel 617 are fabricable by
conventional methods with some modifications due to their high
strength. Since the issues specific to fabrication of these alloys are
widely known, they are not discussed further here.
The ODS alloys are not in wide use in tubular form, but are
routinely available in lengths up to 3.3 m (II ft). For full-scale
applications, tube lengths of 8 m (26 ft) would be desirable.
Fabrication of ODS alloys requires a different approach than with
conventional wrought alloys because of the critical dependence of
their high-temperature strength on the microstructure. The ductilebrittle transition temperature of these alloys (critical temperature for
50% risk of brittle failure) is 50-75 ° C (122-167°F)[McColvin,
19951, so that they must be worked hot; they can be readily bent
after heating to, for instance, 300°C (572°F). Plasma cutting and
wire cutting should be avoided since these processes result in
cracks forming in the alloy surface [Sporer, 19951. These alloys are
readily machinable using conventional methods with appropriate
parameters, but sharp corners should be avoided in component
design. Electrode discharge machining results in cracking in the
recast layer; this can be overcome with care, but then the process is
very slow. Some component designs that work well incorporate a
combination of welding and mechanical fastening. Overall,
however, it is considered best to avoid welding by design, or by
riveting.
Although ODS alloys can be successfully fusion welded, such
joints have a significantly reduced load-bearing capability at
temperature. This is because fusion locally destroys the controlled
distribution of the dispersed phase and disrupts the continuity of
the microstructure, which are the essential features that provide
high-temperature creep strength. Further, fusion welding can result
in cracking at grain boundaries. Laser welding can be performed
such that the dispersoid is retained, but it is typically coarsened
and, as with normal fusion welding, the grain structure of the alloy
is destroyed and some grain boundary cracking usually occurs. No
matter how good the weld, the grain structure and dispersoid
distribution will be disrupted [Wallach, 1995]. Nevertheless, these
changes in structure can be quite acceptable if the weld is outside
the area of critical interest.
deposits of nominally dry flyash. The temperature-dependence of
molten alkali sulfate-accelerated corrosion is considered to follow a
'bell-shaped' curve, as is illustrated in Fig. 5 [after Tamura, et al.,
1985] for several of the alloys of interest exposed to a simulated flue
gas and covered with a simulated deposit for 100 hours. For
essentially all alloys tested, the rate of attack accelerates rapidly
with increasing temperature above approximately 600°C (I112 °F),
reaches a maximum, and then rapidly decreases.
The actual
temperatures may shift with the specific molten salts involved, the
temperature difference between the gas and the alloy surface, and the
specific alloy. The reduced attack at the higher temperatures is
thought to result from the instability of the alkali sulfate species
involved, possibly as a result in a decline in the S02/S03
concentration in the flue gas. As evident from Fig. 5, type 310
stainless steel has the best resistance of the 300-series steels, and
the higher-strength alloys typically exhibit increasing resistance
with increasing chromium content. The highest corrosion resistance
usually is found with Inconel 671 (Ni-48Cr), which sometimes is
used as a cladding. Strengthening additions such as Mo can have a
detrimental influence on the susceptibility to molten salt corrosion.
Solid state bonding, whether by friction welding or explosive
joining, is a feasible method for joining ODS alloys. This process
results in discrete interfaces with unique restructured geometries
and may cause distortion of the grain size and direction. A variant of
friction joining involves a third body, which provides an
essentially mechanical bonding of the components of interest and
does not necessarily involve any melting [Thomas, et al., 1994].
This technique has promise for joining ODS alloys to themselves
and to other difficult-to-join materials. Explosive bonding can
readily be used for tubes, but local melting must be avoided. British
Gas [Starr, et al., 19941 successfully constructed a harp-type heat
exchanger that used ferritic ODS tubes (of 0DM751) by explosively
bonding the ODS tubes to wrought alloy stubs that were then
welded to the headers. The British Gas heat exchanger was part of a
closed-cycle demonstration unit that was run successfully with a
working fluid outlet temperature in excess of 1100°C (2012°F)
[Mabbutt, 1995]. Although the header and stubs were in this case
made of a wrought alloy with lower creep strength than the ODS
alloy, the use of thicker sections for the stubs was practically
acceptable. This approach allows for straightforward field
replacement of individual tubes, but requires that a damaged ODS
tube be completely replaced rather than repaired.
At metal temperatures higher than the upper minimum in the
bell-shaped curve but below the melting points of any simple salts
that might deposit, the main corrosion mechanisms would be
expected to involve gas-phase oxidation, sulfidation or
carburization. Field corrosion probe tests in the superheater regions
of steam boilers fired by the types of U. S. coals expected to produce
molten salt-related fireside corrosion were carried out at metal
temperatures in the range 538 to 871°C (1000 to 1600°F) as part of a
program to evaluate the performance of alloys capable of use in
higher-temperature cycles [Plumley, et al., 19791 After exposures
made for up to 8,000 hr, the performance of alloys Inconel 617 and
Haynes 188 was ranked as very good, with metal loss less than 0.7
nm/hr (0.25 mil/yr.), whereas type 310 stainless steel and Incoloy
800H lost 6 and 5 xun/hr (1.9 and 1.6 mil/yr), respectively.
Essentially, these alloys experienced high-temperature oxidation.
The lack of accelerated corrosion at the temperatures representing
the maximum in the bell-shaped curve was attributed to coal
blending being practiced to avoid this form of corrosion.
Transient liquid phase qui joining is a further possible
approach. If UP joining is done before the heat treatment used to
grow the large grains, the grains can be grown through the joint.
Overall, with ODS alloys it is difficult to attain the properties
of the parent material in a joint. However, there are approaches to
design and materials combinations that can be optimized for specific
applications. If the joint could be kept out of the most severe
service, at least the corrosion rate should be reduced compared to the
rest of the alloy.
Krause et al., (1994) also reported field corrosion data obtained
in support of a closed-cycle gas turbine system with a maximum
working fluid temperature of 843°C (1550 ° F). Those results were
obtained from both superheater and furnace wall regions of coal-fired
boilers and indicated that, in pulverized coal-fired systems burning
coal with sulfur levels in the range 0.8 to 3.1 percent and alkali-plus
sulfur-related corrosion indices (Borio, et al., 1972) up to 3.8, type
304H stainless steel exhibited high corrosion rates which led to the
recommendation that it not be used at metal temperatures above
704°C (1300°F). In contrast, type 310H, and Alloys 800H and 617
showed evidence of forming protective oxide scales in 2,000 hour
exposures at temperatures up to 871°C (1600°F); the projected
corrosion rates were less than 6 run/hr (2 mil/yr).
Environmental Resistance—Fireside. In addition to
the limitations of the strengthening mechanisms employed in these
various classes of alloys, at the higher temperatures environmental
compatibility considerations assume increasing importance. The
sources of corrosion attack on the fireside circuit of coal-fired steam
boilers, and the behavior of most candidate alloys at temperatures
typically encountered in current boilers are reasonably well
understood. The details have been exhaustively analyzed in a
number of reviews [for instance, Stringer and Wright, 1995]. The
potential for fireside corrosion can be minimized by ensuring that
the heat transfer surfaces are located where any fuel particles
entrained in the flue gas are solid, so that direct contact with molten
ash or slag is avoided. The sources of corrosion then involve the
temperature-driven acceleration of oxidation, sulfidation, "and/or
carburization, as well as the possibility of localized attack by the
formation of low-melting species such as alkali sulfates under
The ferritic ODS alloys have excellent resistance to gas-phase
oxidation, sulfidation, and carburization attack, but their corrosion
resistance in the presence of molten alkali salts is not well
established. Experience with molten salt corrosion in marine gas
turbines suggests that alloys which form alumina scales would be
expected to exhibit lower resistance to this form of attack compared
to chromia scale-forming alloys. However, initial results from
4,000+ hr boiler probe tests suggest that some alumina scale-forming
6
A further point for consideration is that, above approximately
1025°C (1877°F) in static dry air environments and at temperatures
as low as 800°C (1472°F) [Lowell, 1972] in flowing, dry air, the
protective oxide scale of Cr203 formed on most of the solid solutionand carbide-strengthened high-temperature alloys can be further
oxidized to 003 which is volatile. The rate of Cr03 formation and
subsequent loss is controlled by diffusion across the aerodynamic
boundary layer [Graham and Davis, 1971], so that the higher the gas
velocity (the thinner the boundary layer), the more rapid the loss.
Recent results [Opila and Jacobson, 1996] indicate that water vapor
has a significant effect on scale volatility, such that the vapor
pressure of an oxyhydroxide species, Cr02(OH)2, exceeds that of
Cr03 by many orders of magnitude at temperatures below
approximately 1100°C (2012°F) in an environment containing 10
percent by volume water vapor. This form of degradation not only
removes chromium from the alloy, so affecting alloy strength as well
as oxidation resistance, but also leads to deposition via
condensation of Cr03 at some point downstream, which also may be
undesirable.
160 _11140 2-5%0,-15%CO 2 -N t
34%Na 2 604-41%
K 2 SO4 - 25%Fe,O,
140 -Ippa test
2
8
120
100
80
60
40
20
0
In high-temperature, high-gas flow applications (such as gas turbine
hot gas path components), alloys or coatings that form a protective
scale of the relatively non-volatile A1203 typically are used instead
of Cr203 -forming alloys. There are techniques that could be used to
provide protective alumina scale-forming coatings on the internal
surfaces of tubing, although there is limited capacity for handling
long tube lengths, and any effect of the heat treatment cycle involved
on alloy properties would require consideration. Since the ODS
alloys MA956, 0DM751, and PM2000 are alumina-formers, they
would not require a coating on the working fluid side.
600 650 700 750 800 850
Temperature ( °C)
Figure 5. Temperature-Dependence of Molten SaltAccelerated Fireside Corrosion of High-Temperature
Alloys (after Tamura, at al., 1985)
Ceramic Heat Exchanger (T,
alloys (Fe3A1-based) may have corrosion resistance at least as good
as the best chromia-forming alloys under conditions where the 300series stainless steels suffered accelerated attack [Slough, 19961.
Apylronmental
Resistance—Working
1100°C/2012°F1
Materials Selection. In 1989, a detailed assessment of the
state of development of ceramic heat exchangers for pressurized duty
concluded that a large gap existed between the perceived
technology level and the actual design database, but that it was
important to pursue the concept [Bevilacqua Knight Inc., 1989].
The areas of main concern were:
(a) Lack of reproducibility of the structure and properties of the
ceramic in the required sizes, including insufficient data on the
size of critical flaws, effect of surface finish, and insufficient
thermal shock capability.
(b) Shortcomings in the then-current tube-string design including
joint wear leading to leakage, effects of deposition on tube
mechanical movement, and durability of seals; insufficient
information on modes of degradation and failure; and concern
that a single tube failure could result in the downstream failure
of other tubes.
(c) Lack of knowledge of the effects of corrosion, erosion, and
deposition at the temperatures of interest with the range of
available coals.
Fluid.
Resistance of the alloys to the high-velocity working fluid and, in
particular, the ability to form a stable oxide scale that is adherent and
will not produce particles of spatted oxide that would be an erosion
threat to the gas turbine also requires attention at the higher
temperatures. The rate of growth of oxide scales in air is typically
faster than in steam at a given metal temperature, and the composition
of the stales is likely to be different as a result of the different
effective oxygen partial pressures of steam and air. It is likely that
the available databases for scale growth and spallation in steam will
need to be supplemented (or replaced); hence, there is a need for
information on scale spallation behavior in air oxidation for all of
the candidate alloys. The ferritic ODS alloys have outstanding
resistance to scale spallation at high temperatures [Quadalckers, et
al., 1994]. Long-term oxidation testing under temperature-cycling
conditions has shown that these alloys essentially do not suffer
scale spallation for several thousand hours; when spallation
initiates it occurs as very small particles, and the alumina scale is
reformed until the aluminum in the alloy is depleted to very low
levels (of the order of 1.5 weight percent) before the onset of
formation of voluminous iron-rich scales which could be a
spallation-erosion threat.
Substantial progress has been made in all of these areas. The
leading candidate materials for this application are based on SiC in
the form of high-density sintered ceramic, as siliconizcd SiC ( ■ 8
volume percent Si) or reinforced with SiC fibers. Oxide ceramics
such as A1203 reinforced with SiC particles also are being
considered [Natesan, et al., 1995]. High-purity, high-density (<3
percent porosity) SIC has high thermal conductivity and the
7
potential for very little change in mechanical properties up to
approximately I500°C (2732°F); it also has good resistance to slow
crack growth up to I200°C (2192°F).
metallic inlet line. The configuration of the ceramic surfaces is
dictated to some extent by the location of the heat exchanger with
respect to the heat source and the need to handle running slag, ash
deposition or clean gas. In one design, a gas-fired ceramic heat
exchanger is located inside, near the outer diameter, of the main coalfired furnace [Solomon, et al., 1996], which allows the heat
exchanger to be heated by radiation from the coal flame and by
convection from the gas burner, but is susceptible to coal ash
fouling. The heat exchanger is made from U-shaped tubes, and the
gas flame fires along the channel formed by the stacked U-tubes.
Aerodynamic cleaning of the tubes is accomplished by using a long
gas flame fired at a slight angle to the axis of the channel.
Mechanical Properties. The strength of a ceramic is
dependent on the size and distribution of flaws, which are potential
initiation points for fracture. Deformation of ceramics may Occur by
diffusion and by grain boundary sliding, but at some critical stress
intensity, crack growth will occur from existing flaws. As a result,
the strength is expressed in statistical terms, usually based on the
approach suggested by Weibull [see, for instance, Ashby and Jones,
19863 which provides a measure of the fraction of identical samples
expected to survive loading to a specified tensile stress. The
variability of the strength is indicated by the Weibull modulus,
which is found finm testing. Hence, there is no single value for
representing the strength of a ceramic. Implicit in this approach is
the importance of the size and geometry of the part, as well as its
surface finish. In practice, since a statistical approach to design
implies that a failure will occur, the designer either needs to employ
a very large safety margin until experience is gained with the specific
material, or to devise a failure-tolerant design that allows easy
maintenance. An understanding of the likely failure mode is
essential for assessing the likely location of failure, the actions to be
taken when a failure occurs, and the spare parts inventory needed.
Joining the ceramic to itself, and to the header from the hightemperature metallic heat exchanger, still represents a challenge and
probably will require continuing development A particular
problem will be the application of expansion joints under
conditions where a significant thermal and pressure differential
exists across the tube wall, and where there is a strong possibility of
high-temperature corrosion. A further complicating factor is the
need to accommodate the different rates of heat-up and cool-down of
the component parts, especially since these plants will likely
experience some level of cycling duty.
Environmental Resistance. The actual maximum
temperature that can be realized in practice probably depends on the
resistance of the ceramic to fireside corrosion in the location in the
coal-fired furnace required to provide the desired heat flint The
important measure of corrosion is its effect on the size and
distribution of surface flaws that control the strength of the material.
Hence, selective attack of features such as grain boundaries to create
flaws or of pre-existing flaws must be minimized.
As with some metallic alloys, measurements of long-time mechanical
property data are sparse for ceramic materials. Some results from an
on-going effort to generate such data indicate that, to achieve
lifetimes approaching 100,000 hours in air for a commercial gas
pressure-sintered Si3N4, a static tensile stress of 300 MPa (43.5 ksi)
at 1038 and 1150°C (1900 and 2IO2°F), and 125 MPa (18 ksi) at
1350°C (2462°F) cannot be exceeded. For SiC, a static tensile stress
of 300 MPa at I038°C, 250 MPa (36 ksi) at 1150°C, and 180 MPa
(26 Its° at 1350°C cannot be exceeded. Failures in the SiC were a
result of slow crack growth that initiated from the specimen surface,
whereas it appeared that the fatigue failure in the Si3N4 was related
to the creep mechanism [Wereszczak, et al., 1996]. Note that these
strengths are all well in excess of those needed for the cycles
indicated in Fig. 4.
Silicon carbide typically forms a surface film of a-Si02 in
oxidizing environments at high temperatures, which is protective
(passive oxidation). Under certain conditions of reduced oxygen
partial pressure, which vary with temperature, the oxide formed is
SiO, which is volatile; this behavior is known as "active
oxidation." Such conditions could occur in areas of a combustor
which were persistently substoichiometric, as may be the case under
some conditions of low-NO. operation, or under slag layers where
unbumt carbon could create a locally reducing environment.
Design Approaches. The features that require attention in
any ceramic heat exchanger design are:
• prediction of the stress distribution, especially around joints and
at points of contact,
• ability to maintain leak-proof joints,
• design to avoid fireside fouling or incorporate methods to prevent
or remove deposits,
• strategies for repair or replacement in the field, and
• resistance to thermal cycling.
The configuration of the ceramic heat exchanger should utilize
shapes for the heat transfer surfaces that minimize any tensile
components of static or dynamic loading, including possible loads
from the build-up of deposits that could be molten, hence strongly
adherent. Various configurations of tubes have been extensively
explored using detailed stress analyses [Dapkunas, 1988] as a
guide. Designs have been based on U-tubes [Carpenter, et al., 1980]
or on straight tubes [LaHaye, 1995; Shenker and Torpey, 19961,
with emphasis on modular design to simplify manufacture and
inspection, and on approaches to accommodate the mismatch in
coefficient of thermal expansion between the ceramic headers and the
The resistance of SiC to molten slags or molten alkali sulfate
deposits is good if the deposit is acidic, but rapid attack can occur
in basic melts [Ferber, et al., 1985]. The melting temperature and
acid-base nature of slag or ash deposits from coals can be predicted
from the coal composition and can be modified by coal blending or
by the use of additives. Work in progress [Breder et at, 1996;
flannel, 1996] is aimed at quantifying the effects of slags from a range
of coals on the corrosion behavior and strength degradation of
several candidate ceramics.
Auxiliary Fired Heater/Topping Combustor
In either open or closed cycles, some form of topping combustor
appears to be an essential component for achieving acceptable cycle
efficiencies. In an open cycle, the auxiliary fuel probably will be
burned in a modified version of the combustor typically used on a
fired turbine. The design of this combustor will be complicated by
the need to handle the high-temperature air from the main heat
8
exchanger. The maximum temperature of the inlet air to the combustor
in conventionally-fired turbines is significantly less than 871 °C
(I600°F), so that modification to the cooling system used for the
combustor can, or the use of materials such as ceramic matrix
composites for the combustor liner will be necessary.
conventional wrought alloys, for which the required data are being
developed. The main issues concerning the application of these
alloys are similar to those faced in coal-fired steam power plants,
namely fireside corrosion in high-flux regions, including the
performance of weldments, and scale spallation on the working fluid
side at the higher temperatures. For the advanced alloys, the
generation of long-term creep rupture data is needed.
In the case where the fuel used in the gas turbine is pyrolysis
gas derived from the coal [Shenker and Torpey, 1996], the materials
selection for the transfer line from the pyrolyzer to the turbine is
complicated by the need to handle a hot, reducing gas coupled with
the need to avoid refractory linings which can potentially spall,
leading to erosion of the turbine. Since the pyrolysis gas will
probably carry entrained ash particles, the transfer line will likely
contain a hot gas filter, in this case, refractory linings must be
avoided only between the filter and the turbine.
This ODS-ferritic alloys have demonstrated high-temperature
strength capabilities, and have been successfully fabricated into a
closed-cycle heat exchanger configuration. Because of the need to
preserve the alloy microstructure of ODS alloys in the region where
joints are made, there is a need for demonstration of the capabilities
of the several available joining techniques in simulated service
conditions. There is also need for the quantification of the fireside
corrosion behavior in the regions of interest in a coal-fired system.
Further developments of microstructural and processing control to
increase in the creep strength in the circumferential direction are
desirable to further increase the temperature capability of these
alloys or allow the use of thinner sections. The cost of the ODS
alloy tubes currently is very high, partly as a result of the low
volume of current production.
In a closed cycle the auxiliary fired heater or topping combustor
will probably be a natural gas-fired heat exchanger, hence will
require similar considerations to the main ceramic heat exchanger.
Major differences will be the ability to better control the uniformity
of the heat flux on the tubes and the absence of molten salt
corrodents. For this unit, advantage probably could be taken of
work on ceramic heat exchangers for gas turbine recuperators, which
is aimed at turbine inlet temperatures in the range of I350°C
(2462°F). In one example [Yoshimura, et al., 1995], SN-84 SiC was
used for the tubes and tube-sheet, and a simple design was employed
which involved straight ceramic tubes and a clamping arrangement
using a compliant layer between the ceramic tubesheet and the
metallic headers. One end of the ceramic tube bundle was free to
elongate and rotate to accommodate changes caused by variation in
heat distribution.
A continuing need for the coal-fired ceramic heat exchanger is
an adequate body of design data for the as-fabricated condition that
will exist in commercial components. Inherent in this need is the
ability to manufacture larger parts/longer tubes with a higher
assurance of quality control. Techniques capable of identifying
surface defects of the order of 10 gm under field conditions would
greatly facilitate inspection and improve confidence in the operation
of these components. Joining of ceramic-to-ceramic and to metals
remains a challenge. For coal-fired applications, a major uncertainty
for the realization of the potential of the ceramic heat exchanger is the
effect of the fireside environment on the critical flaw size. Designs
need to address the practical implications of tube failure and
replacement in service.
SUMMARY
The materials available for construction of the heat exchangers
required in indirectly-fired gas turbine cycles have been considered
as three distinct classes:
Conventional wrought ferritic and austenitic alloys which, as
tubes 5.1 cm (2 in.) diameter and 6 mm (0.25 in.) wall thickness,
are suitable for use to temperatures up to 9I5°C (I680°F) in an
open-cycle system, and up to approximately 860 °C (I580°F) in
a closed-cycle system with a small turbine.
Oxide dispersion-strengthened (critic alloys which are capable
of use to at least 935 °C (I715°F) for all open cycle sizes
considered and small (<50 MW) closed-cycle turbine systems,
and up to I100°C (20I2°F) for lower pressure ratio turbines in
an open-cycle system.
(iii) Ceramics which can be used in heating the working fluid to
temperatures above approximately 935°C (1715°F).
Overall, there are materials available that can enable the
construction of externally-fired heat exchangers capable of
delivering air at the desired pressures at temperatures up to 935 ° C
(I715°F, open cycle and small closed-cycle) for an all-metallic
system and to possibly I300°C (2372°F) for a metallic plus ceramic
system. There are some gaps in the data required to commit to a fullsize plant, but there appears to be no reason why these needs cannot
be met with properly focused research, The high temperatures
possible with ODS alloys would allow the size of the following
ceramic heat exchanger to be reduced, with the possibility that more
appropriate or innovative designs could be better tailored to
accommodate the needs of the ceramics, especially the difficult
ceramic-to-metal joints.
Some of the unique properties of the class (ii) and (iii) materials
may render them unsuitable for the direct substitution of class (i)
materials in conventional heat exchanger designs. This is not
necessarily a drawback, since the most effective design of an
indirectly-fired heat exchanger is likely to use different modules
located in the gas path so that they encounter the most appropriate
heat flux and corrosive environment. Such a modular arrangement
also permits the design of each module to be most appropriate for the
properties of the materials being used.
The required strength and fabrication data are well established
for the conventional alloys, and long-term environmental resistance
data are being generated. This group also includes some advanced
ACKNOWLEDGMENTS
The authors are grateful the support of the Electric Power Research
Institute, and of the Fossil Energy Advanced Research and
Technology Development (AR&TD) Materials Program, U. S.
Department of Energy, under contract DE-AC05-840R21400 with
Lockheed Martin Energy Research, Inc. In addition, discussions
with colleagues at the Oak Ridge National Laboratory: in
particular, Drs. K. Breder, M. K. Ferber, It R. Judkins, Mr. It W.
Swindeman, Dr. A. W. Wereszczak, and with Prof. N. Birlcs of the
University of Pittsburgh are gratefully acknowledged.
9
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10
APPENDIX 1: WORKING FLUID CONSIDERATIONS
Wallach, R., "Joining MA-ODS Materials," Presented at AEA
Technologies Workshop on ODS Alloys, Risley, England, March
1995.
Wereszczalc, A. A., and Kirkland, T. P., "Creep Performance of
Candidate SiC and Si3N4 Materials for Land-Based Gas Turbine
Components," ASME paper No. 96-01-385, Presented at the Intl.
Lee et al. (1980) compared the pertinent properties of potential
working fluids and chose helium, helium/carbon dioxide (0.6 He/0.4
CO2 by mole), and air as candidates for a closed-cycle application
similar to that considered here. Under the conditions considered,
the heat transfer coefficients for the these three candidate working
fluids were in the order: 1.0 : 0.87 : 0.52, respectively, indicating
that with air a larger temperature drop would be experienced across
the heat exchanger wall (and a significantly larger surface area would
be needed) for the same duty. Lee et al. (1980) also showed that the
heat transfer coefficient in a coal-fired, closed cycle system varied as
the 0.6 power of the working fluid pressure, so that increasing the
total operating pressure will reduce the tube wall temperature and
the amount of surface area required.
Gas Turbine and Aeroengine Congress and Exhibition,
Birmingham, England, June 1996.
Yoshimura, Y., ltoh, K., Ohhori, K., Hori, M., Hattori, M.,
Yoshida, T., and Watanabe, K., "Development of Shell and TubeType Ceramic Heat Exchanger for C01301," ASME paper No. 95GT-208, Presented at the Intl. Gas Turbine and Aeroengine
Congress and Exposition, Houston, Texas, June 1995.
11

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