HK1071774B - Ferritic stainless steel having high temperature creep resistance - Google Patents

Description

Ferritic stainless steel with high temperature creep resistance

Inventor(s):

John F.Grubb

cross reference to related applications

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Statement relating to federally sponsored research or development

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Technical field and industrial applicability of the invention

The invention relates to a ferritic stainless steel alloy. More particularly, the present invention relates to a ferritic stainless steel alloy having microstructural stability and mechanical properties that make it particularly suitable for high temperature applications. Such applications include, but are not limited to, current collection interconnects in solid oxide fuel cells, kiln hardware, equipment used in chemical process, petrochemical, power generation and pollution control industries, and equipment used in processing molten copper and other molten metals.

Description of the invention

Fuel cells are efficient, environmentally friendly power generation devices. The basic principle of fuel cell operation is the generation of electricity by combustion of a fuel. The fuel is separated from the oxidant by a permeable barrier (permeable barrier) known as an electrolyte. The hydrogen atoms on the fuel side of the electrolyte are ionized. The generated protons pass through the electrolyte while the released electrons travel through an external circuit. On the air side of the electrolyte, opposite to the fuel side, two protons combine with oxygen atoms and two electrons to form water molecules, releasing heat and forming an electrical circuit. Energy is extracted from the process by using electronic work in an external circuit. For fuel cells operating at higher temperatures, the heat released by the reaction on the air side can also be used for fuel reforming (fuel reforming) or heating applications, increasing the overall operating efficiency of the cell.

Currently, a more attractive class of fuel cells is the Solid Oxide Fuel Cells (SOFC). SOFC's operate at high temperatures (1450-. Internal reforming recycles thermal energy and eliminates the need for expensive platinum-based metal catalysts. Both hydrogen and carbon monoxide may be used as fuels in SOFC. In the modification of the conventional fuel cell reaction described in detail previously, hydrogen is combined with oxygen. The electrolyte is an electrolyte capable of reacting with oxygen ions (O)^2-) Rather than permeable to protons. Therefore, SOFCs operate in the opposite direction relative to certain other types of fuel cells. In addition to combusting hydrogen, carbon monoxide is oxidized at the anode to carbon dioxide, releasing heat. This is an advantage because carbon monoxide is present in unrefined fuels and can poison low temperature fuel cells, reducing operating efficiency. Small SOFC's operate at efficiencies up to about 50%. To achieve even higher efficiencies, medium sized and larger SOFC's may be combined with gas turbines. The efficiency of the obtained mixed SOFC-gas turbine unit can reach 70%.

There are many variations on the design of SOFCs. The electrolyte is typically in the form of zirconia, which is stabilized by the addition of lattice change inhibiting oxides and provides high ionic conductivity when heated to high temperatures. Such oxide-stabilized materials are generally known and are referred to herein as "stabilized zirconia". Typically, SOFCs include Yttria Stabilized Zirconia (YSZ) as the stabilized zirconia electrolyte. The reported Coefficient of Thermal Expansion (CTE) of YSZ at 20 deg.C (68 deg.F) to 10000 deg.C (1832 deg.C) is about 11X 10^-6/℃。

Research has been conducted on tubular SOFCs that operate at very high temperatures (1800 ° f (982 ℃)) and are of relatively simple construction of large size. Tubular SOFCs can be scaled up by increasing the size and number of individual SOFC tubes in the device. "planar" SOFCs (PSOFCs) have been recently investigated. PSOFC's are fairly compact and consist of many planar cells. Typically, the anode and cathode plates are usually ceramic materials. Permeable nickel-zirconia cermets are also used for the anode.

Interconnects are required to collect the electrons produced by the fuel cell. The interconnects also act as physical barriers to oxidative and reductive gas flow. Thus, the materials used to form the fuel cell interconnects should be electrically conductive, oxidation resistant, and mechanically stable, and should have thermal expansion properties that substantially match the ceramic elements of the cells that may be physically disposed adjacent to the interconnects. Until recently, SOFC interconnects have typically been fabricated from ceramic materials that are electrically conductive at high temperatures and are typically LaCrO doped with CaO or SrO₃. Although ceramics are generally stable, ceramics are also brittle and relatively expensive when subjected to high temperatures for extended periods of time, being poor conductors of electricity relative to metals. Interconnects of certain metals have been developed for this purpose from chromium-based alloys. The alloy provides adequate oxidation resistance, as well as good thermal expansion matching with stabilized zirconia. However, the powder metallurgy used to make the alloy makes it very expensive, and significant capital investment is made in SOFC's made from the alloy.

Making SOFC interconnects from stainless steel is preferred over ceramic because steel has higher electrical conductivity and is less brittle than ceramic. However, problems associated with the use of stainless steel in SOFC interconnects include oxidation, thermal expansion, and creep issues. Oxidation can reduce the ability of the stainless steel to conduct electricity, thereby reducing the output of the cell over time. Standard austenitic stainless steel stainless provides good thermal expansion matching with conventional SOFC electrolyte ceramics. Ferritic stainless steels, which can have good thermal expansion matched to the ceramic electrolyte, have low creep resistance. For example, the inventors treated several commercially available stainless steels, including E-BRITE^®(UNS S44627)、AL 29-4-2^®(UNS S44800) and ALFA-IV^®Experiments with (AlloyDigest SS-677, ASM International) alloys demonstrated that E-BRITE^®The alloy has acceptable thermal expansion for SOFC, good thermal stability and forms desirable Cr₂O₃An oxide. However, E-BRITE^®The creep resistance of the alloy is less than that required for use in SOFCs.

Accordingly, there is a need for improved stainless steels having high temperature creep resistance, good thermal stability and other properties that make them suitable for use as current collecting interconnects in SOFC's and in other high temperature applications, such as equipment used in the chemical process, petrochemical, power generation and pollution control industries, as well as kiln hardware components and equipment used for processing molten metals.

Summary of The Invention

The present invention addresses the above-described needs by providing a ferritic stainless steel comprising greater than 25 weight percent chromium, 0.75 up to 1.5 weight percent molybdenum, up to 0.05 weight percent carbon and at least one of niobium, titanium, and tantalum, wherein the sum of the weight percentages of niobium, titanium, and tantalum satisfies the relationship 0.4 ≦ (% Nb +% Ti +1/2 (% Ta)) ≦ 1. The inventive steel has a CTE that is within a range of about 25% higher to about 25% lower than the CTE of the stabilized zirconia at 20 ℃ (68 ° F) to 1000 ℃ (1832 ° F). The steel of the invention also has at least one creep property selected from the group consisting of: a creep rupture strength of at least 1000psi at 900 ℃ (1652 ° F); a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° F).

The present invention also provides a method for manufacturing a ferritic stainless steel, wherein the method comprises forming a ferritic stainless steel comprising greater than 25 weight percent chromium, 0.75 to less than 1.5 weight percent molybdenum, up to 0.05 weight percent carbon, and at least one of niobium, titanium, and tantalum, wherein the total weight percent of niobium, titanium, and tantalum satisfies the relationship 0.4 ≦ (% Nb +% Ti +1/2 (% Ta)) ≦ 1. The CTE of the steel is within the range of about 25% higher to about 25% lower than the CTE of the stabilized zirconia, with preferred CTEs being 25% higher or lower than the CTE of the stabilized zirconia at 20 ℃ (68 ° F) to 1000 ℃ (1832 ° F) and within this range. The steel also has at least one creep property selected from the group consisting of: a creep rupture strength of at least 1000psi at 900 ℃ (1652 ° F); a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° F). In the subsequent step, the steel is solution annealed (solution annealing) and then cooled from the annealing temperature. Solution annealing is preferably performed at a temperature at least above the intended service temperature of the alloy and 1600F (871℃.). The solution annealed stainless steel is subjected to precipitation heat treatment to harden the steel, if necessary.

The invention also provides articles machined from the stainless steel of the invention. The article may be manufactured using conventional methods.

The stainless steels of the present invention exhibit improved high temperature mechanical properties, including improved high temperature creep resistance, relative to other ferritic stainless steels. The steel will also show good thermal expansion matching to YSZ, a stable zirconia commonly used as an electrolyte in SOFC's. Thus, the steel is suitable for use as current carrying interconnects and flow separators in SOFC's and may be used instead of ceramics. The steel is also suitable for use in high stress, high temperature applications, including, for example, oxygen sensor devices, certain chemical processes, petrochemicals, power generation and pollution control equipment, high temperature furnace hardware components, and molten metal processing equipment.

The reader will appreciate the foregoing details and advantages of the invention, as well as others, when discussing the following detailed description of embodiments of the invention. The reader also may comprehend additional details and advantages of the present invention upon making and/or using the stainless steels of the present invention.

Brief Description of Drawings

FIG. 1 is a graph of annealing temperature versus ASTM grain size for various ferritic stainless steels;

FIGS. 2(a) - (c) are graphs illustrating several mechanical properties of several ferritic stainless steels tested at different temperatures;

FIG. 3 is a graph of stress versus 1% creep strain time applied to several ferritic stainless steels tested at test temperatures of (a)800 ℃ (1472F), (b)850 ℃ (1562F), and (c)900 ℃ (1652F);

FIG. 4 is a graph of stress versus 2% creep strain time applied to several ferritic stainless steels at test temperatures of (a)800 ℃ (1472F), (b)850 ℃ (1562F), and (c)900 ℃ (1652F);

FIG. 5 is a graph of stress and time to failure for several ferritic stainless steels at test temperatures of (a)800 deg.C (1572 deg.F), (b)850 deg.C (1562 deg.F), and (c)900 deg.C (1652 deg.F);

FIG. 6 is a graph of time versus weight change of several ferritic stainless steels exposed to the atmosphere at 800 ℃ (1472 ° F) and depicting isothermal oxidation data thereof;

FIG. 7 depicts isothermal oxidation data obtained by exposing several ferritic stainless steels to the atmosphere at 800 ℃ (1472 ° F);

FIG. 8 depicts isothermal oxidation data obtained by exposing several ferritic stainless steels to the atmosphere at 900 ℃ (1652 ° F); and

fig. 9 depicts a graph of the relationship between cycle temperature and average cycle fatigue failure (CTF) value for several ferritic stainless steel samples 0.002 inches thick.

Detailed description of embodiments of the invention

It is hypothesized that replacing the ceramic SOFC interconnect with a stainless steel interconnect would provide advantages. However, initial work in this area has shown disadvantages in the various existing stainless steels considered. For example, austenitic nickel-based materials have been found to exhibit poor coefficients of thermal expansion ratios. Ferritic alloys formed from alumina were found to be deficient because they were not conductive after oxidation.

The present inventors also evaluated certain products available from Allegheny Ludlum Corporation, Pittsburgh, Pennsylvania under the trademark AL29-4-2 at elevated temperatures^®、ALFA-IV^®And E-BRITE^®The suitability of commercially available ferritic stainless steels for interconnection in SOFC's. AL29-4-2^®The alloy is described by UNS number S44800 and is juxtaposed in several ASTM numbers, including a 240. AL29-4-2^®With typical composition ranges (weight percent) of 28.0-30.0 chromium, 3.5-4.2 molybdenum, 2.0-2.5 nickel, balance iron and residual impurities. ALFA-IV^®The alloy is a patent alloy generally described in U.S. patent No. US4,414,023, rated for a composition of 20 wt.% chromium, 5 wt.% aluminum, and 0.3 wt.% rare earth metal. E-BRITE^®The nominal composition of the alloy is 26 weight percent chromium, 1 weight percent molybdenum stainless steel, which is described schematically in U.S. patent US3,807,991.

It was found that AL-29-4-2^®The alloy is severely embrittled at high temperatures due to the large amount of sigma phase precipitation (embrittlement). ALFA-IV^®The alloy showed a higher than adequate level of thermal expansion and was found to form an undesirable non-conductive a1₂O₃And (3) a membrane. Discovery of E-BRITE^®Alloy ratio AL-29-4-2^®And ALFA-IV^®Alloys are generally more suitable for SOFC interconnects, but they are still unsuitable, primarily due to unacceptably low creep resistance at high temperatures.

Ferritic stainless steels having enhanced high temperature mechanical properties, including enhanced high temperature creep resistance, relative to commercial forms of E-BRITE^®The alloys would be beneficial in applications such as SOFC interconnects and other high temperature applications. Through experimentation, the inventors have determined that ferritic stainless steels include greater than 25 weight percent chromium, 0.75 up to 1.5 weight percent molybdenum, up to 0.05 weight percent carbon, and 0.4 up to 1 weight percent niobium. Preferably, the carbon content of the alloy is limited to 0.005 wt.%, but the presence of niobium or another carbide former such as titanium in the alloy will also provide carbide stability in a range of up to a broader 0.05 wt.%, as described below.

The ferritic stainless steel of the invention is also characterized in that it has at least one creep property selected from the group consisting of: a creep rupture strength of at least 1000psi at 900 ℃ (1652 ° F); a time to 1% creep strain of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a time to 2% creep strain at 1000psi load, 900 ℃ (1652 ° F) of at least 200 hours.

Because YSZ is a stable zirconia electrolyte common in SOFC's, preferably, the inventive steel has a CTE that is within about 25% higher to about 25% lower of the CTE of YSZ at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f). As disclosed above, the CTE of YSZ is about 11X 10 over this temperature range^-6V. C. Thus, the CTE value ranges from about 25% higher to about 25% lower by about 8.25-13.75X 10^-6/℃。

At operating temperatures, less creep and/or stress relaxation of the metal elements of the SOFC leaves the device substantially stress-free shortly after that temperature. When the SOFC is subsequently cooled, if the CTE of the metal is less than the CTE of the stabilized zirconia electrolyte, the metal is placed in compression while the ceramic is placed in tension. It is well known that frangible materials are preferably stressed and can be accidentally damaged when placed under tension. Thus, it is preferred that the metal have at least the same CTE as the oxide stabilized ceramic. Thus, the CTE of the ferritic stainless steel of the invention is preferably at least as great as the CTE of stabilized zirconia, such as YSZ, and can be as high as greater than 25% of its CTE, such as at 20 (68F.) to 1000C (1832F.). The inventors have also found that in order to optimize the properties of the ferritic stainless steel of the invention for SOFC interconnections, it is preferred to solution anneal the steel during treatment and then cool from this annealing temperature. Solution annealing is preferably performed at a temperature at least above the intended service temperature of the alloy and 1600F (871℃.). The inventors have discovered that annealing the alloy at too high a temperature (e.g., above 2200F. (1204℃.)) for an extended period of time can result in excessive grain growth, which can compromise the hardness and formability of the alloy. Rapid cooling from this annealing temperature, such as by water quenching, has not been found to be necessary, but is not detrimental. Very slow cooling, such as by furnace cooling, has not been found to be necessary. Generally, air cooling or cooling by alternative methods at equivalent rates is generally preferred. Alloys used in applications where increased hardness is desired to alter certain mechanical properties may be solution annealed stainless steels by precipitation heat treating (precipitation heat treated) by conventional methods.

Chromium helps to increase the oxidation resistance of stainless steel and helps to form Cr that is conductive at high temperatures₂O₃Scale (scale). It is also primarily responsible for reducing the thermal expansion of the steel, so it is usually matched to that of zirconia. It is believed that steels containing less than about 25 wt.% chromium do not exhibit these desirable properties. However, when the chromium content is increased above about 35 wt.%, the steel becomes more difficult to hot work and, therefore, more expensive to manufacture. Moreover, such high chromium content steels will be more likely to form (from) an undesirable intermetallic sigma (FeCr) phase. Thus, the chromium content is preferably no greater than about 35 wt.%, more preferably no greater than about 30 wt.%, and even more preferably no greater than about 27.5 wt.%.

Molybdenum reduces thermal expansion. It is composed of a base, a cover and a coverAlso provides the function of strengthening solid solution, and forms strengthened Laves phase Fe together with niobium₂(Nb, Mo) precipitates. However, molybdenum mainly increases the tendency of stainless steels to precipitate undesirable sigma phases, and also undesirable chi (Fe, Cr, Mo) phases. Molybdenum also impairs the oxidation resistance of the steel and, in some cases, can contribute to catastrophic oxidation formation. For this reason, the molybdenum content of the stainless steel is preferably carefully controlled. Molybdenum content is from about 0.75 up to 1.5 wt.%, more preferably up to about 1.2 wt.%, which provides a particularly suitable balance between desirable and undesirable elemental effects on the properties of the alloy. In particular, experimental alloys made by the present inventors containing 0.9-1.1 wt% molybdenum show a particularly desirable balance of properties.

The role of carbon in ferritic stainless steels is well known. A carbon content of less than about 0.010 wt.% is desirable to achieve ductility in unstable alloys. To optimize performance, the carbon content needs to be less than 0.005 wt.%. However, the niobium content in the stainless steel of the present invention will mitigate the effect of many carbons. For this reason, a carbon content of up to about 0.05 wt.% is acceptable if sufficient carbide-forming elements are present to stabilize the carbon content. One of ordinary skill in the art can readily determine the amount of carbide-forming elements that must be present in a given alloy of the present invention to stabilize a given carbon content. If a welded article is to be formed from the steel of the invention, it is preferred to refer to a preferred upper limit of 0.005 wt.% to prevent hot cracking of the weld.

It has been found that the addition of a small amount of niobium can improve the creep or "sag" resistance (sag resistance) of ferritic stainless steels. Under normal circumstances, these niobium additions result in a Laves phase (Fe)₂(Ta, Nb, Mo)) precipitates are finely dispersed. Suitable niobium content in the stainless steel of the present invention is determined experimentally, as described below. It is believed that titanium may replace a portion of the niobium in the alloy. In addition, tantalum is similar to niobium in its effect on alloy properties, but tantalum is heavier and substantially more expensive than niobium. It is believed that on the basis of 2 wt% tantalum corresponding to 1 wt% niobium and titanium,tantalum may replace niobium and titanium, either completely or partially. Thus, it is believed that the enhanced properties observed by the present inventors for the stainless steel of the present invention may be obtained by including in the steel at least one of niobium, titanium, and tantalum, wherein the sum of the weight percentages of niobium, titanium, and tantalum satisfies the following relationship:

0.4≤(％Nb+％Ti+1/2(％Ta))≤1

preferably, the steel of the present invention comprises not more than 0.50 wt.% titanium.

One advantage of adding titanium to the stainless steel of the present invention is that it will remove nitrogen from the solid solution (solution) in the form of TiN. This will better prevent NbN and CrNbN precipitate formation, thus remaining for the formation of desirable lafuties (Fe)₂Nb) phase strengthening precipitates niobium (a more expensive alloying addition than titanium). It is also believed that the addition of titanium similarly removes carbon from solid solution, thereby better preventing the formation of NbC and NbCN. It is also noted that amounts of titanium above about 0.07 wt.% reduce the niobium-induced weld cracking problem.

In order to better ensure a significant increase in the high temperature performance, while limiting the investment associated with alloying additions, the sum of the weight percentages of niobium, titanium and tantalum in the steel of the invention is more closely controlled to satisfy the following relationship:

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75

wherein the maximum and preferred contents of titanium are the same as in the above relation.

In addition, the ferritic stainless steel of the present invention may contain addition of one or more rare earth elements to the above elements. These optional rare earth additions include, but are not limited to, up to about 0.1 wt.% cerium and up to about 0.05 wt.% lanthanum. It has been shown that the addition of rare earth elements as alloying additions is very beneficial for improving the oxidation resistance of iron-based alloys. This effect has been demonstrated for yttrium, lanthanum and cerium. Other rare earth elements are increasingly expensive and less effective, but may also be used for the above purposes. When these elements are added to the stainless steel of the present invention, it is not necessary to add only a single Rare Earth Metal (REM). Commercially produced mixtures of REM elements, known as misch metal alloys, can be used to provide economical REM doping. As is known in the art, misch metal alloys are naturally derived mixtures of metallic rare earth elements containing about 50 weight percent cerium with the balance being primarily lanthanum and neodymium.

Different mechanisms have been proposed for the effect of rare earth elements on the oxidation resistance of metal alloys. Currently, the most widely accepted mechanism is based on modification of internal surfaces such as oxide/oxide grain boundaries and oxide/metal interfaces. An improvement to this mechanism is the "poisoned" interface "model, in which REM atoms bind (tie up) sulfur at the oxide/metal interface. This mechanism is accepted to be supported by the finding that reducing sulfur in REM-free alloys to ultra-low levels (less than 1ppm) has nearly the same effect as adding REM to alloys with typical sulfur contents (3-100 ppm). Other theories that have been proposed include increased scale plasticity, promotion of protective oxide formation, and mechanical fixation of the scale to the metal by forming rare earth oxide stops (pegs). Regardless of the actual mechanism, it is important to the present invention that corrosion resistance is improved by the addition of REM. It is important not to add too much REM because these elements have limited solubility in iron-based alloys and excessive solute forms undesirable intermetallic phases, deep (deep) eutectics, or both, with very severe impairment of hot workability. High levels of REM can also lead to "over-doping", characterized by the formation of islands of REM oxide and an elevated oxidation rate.

The addition of other non-REM elements may also provide improved oxidation resistance. In particular, hafnium provides advantages similar to those provided by the addition of REM. Hafnium, however, is very expensive. Zirconium is much less costly and can be substituted in amounts similar to hafnium, although zirconium is less effective. As with REM elements, the amount of zirconium and/or hafnium contained in the alloy should not be too great, otherwise excessive amounts of undesirable intermetallic phases may form. Thus, hafnium and/or zirconium may be included in the alloy in a combined amount up to about 0.05 wt.%.

Other alloying elements and additives known in the art may also be added to the alloy to enhance or provide additional properties. Such additives include, for example, silicon, aluminum, tungsten, and manganese. In steel making, silicon is used as a deoxidizer. Which promotes the precipitation of the lafaxine phase and the undesirable sigma phase. In solid solution, silicon hardens the ferrite and makes it brittle. Thus, if present, the silicon content of the present alloy is preferably limited to less than about 1 wt.%, more preferably less than about 0.5 wt.%.

Aluminum is both a deoxidizer and a hardener. Since aluminum is a more effective deoxidizer than silicon, a lower residual content of aluminum is required to produce complete oxidation. If present, the aluminum content is preferably less than about 0.25 weight percent, more preferably from about 0.002 to 0.05 weight percent.

Tungsten generally works similarly to molybdenum, but is heavier, more expensive, and more difficult to alloy. It may be incorporated with the molybdenum but, if present, is preferably less than about 0.25 wt%.

Manganese is intentionally added to carbon steel to mitigate sulfur-induced hot shortness. It is typically present in stainless steel, but in the present alloy it is preferred to limit it to less than about 1 wt.%, more preferably to less than about 0.5 wt.%.

Unavoidable impurities may be present in the stainless steel of the present invention. Of which nitrogen, copper, sulfur and phosphorus are important. The molten Fe-Cr alloy readily absorbs nitrogen when in contact with air. As the chromium content of such alloys increases above about 18 wt.%, nitrogen removal becomes more difficult. Nitrogen in ferritic stainless steels is often embrittled by one of chromium nitride or aluminum nitride precipitates. Preferably, the nitrogen content of the steel of the invention is limited to less than about 0.04 wt.%, more preferably to less than about 0.010 wt.%. Sulfur is an unavoidable impurity in steel making and is generally undesirable. Which is easily removed during Argon Oxygen Decarburization (AOD) refining rather than Vacuum Induction Melting (VIM) refining. As known to those of ordinary skill in the art, AOD is a secondary refining process for the controlled oxidation of carbon in a steel melt in which oxygen, argon and nitrogen are injected into the molten metal bath through submerged, side-inserted tuyeres. VIM is a refining and remelting process in which metal is melted inside a vacuum chamber by induction heating.

The sulfur is preferably reduced to the lowest level that can be most easily achieved and, in any event, should preferably be no greater than about 0.010 weight percent. Phosphorus is a solid solution strengthening agent for steel and can cause brittleness. Phosphorus is not easily removed from stainless steel and therefore is not easily reduced to very low levels, but is preferably limited to less than about 0.050% by weight. Copper is not easily removed during the steel making process, but copper is almost harmless. High levels of copper (greater than about 2 wt.%) compromise the hot ductility and hot workability of ferritic stainless steels. In E-BRITE^®Copper is limited to no more than about 0.025 wt% of the alloy to better provide resistance to Stress Corrosion Cracking (SCC) in boiling magnesium chloride solutions. High SCC resistance is not a specific object of the present invention, and copper is preferably limited to less than about 0.25 wt%.

Before conducting tests to determine the performance of different ferritic stainless steels, 6 50 pound steel melts (heats) having the compositions described in table 1 below were prepared by VIM, numbered WC70-WC 75. All values shown are expressed in weight percent of the total molten steel weight.

TABLE 1

Molten steel	WC70	WC71	WC72	WC73	WC74	WC75
							CMnPSSiCrNiMoAlNbCeLaZrN	0.00260.0540.0100.00290.1625.520.0961.050.0020.12＜0.001＜0.001＜0.0010.0010	0.00260.0550.0100.00270.1525.980.0941.050.0020.68＜0.001＜0.001＜0.0010.0010	0.00380.0600.0100.00140.1425.630.0951.030.0020.130.0010.001＜0.0010.0008	0.00220.0490.0100.00110.1525.770.0941.040.0020.680.0030.001＜0.0010.0009	0.00230.0520.0100.00030.1525.690.0941.040.0020.710.0420.016＜0.0010.0011	0.00330.0530.0100.00060.1525.790.0951.040.0020.710.0090.0030.0110.0011

Molten steels WC70 and WC72 are representative of standard ferritic stainless steels having 0.37 wt.% or less niobium and 0.001 wt.% or less of one of cerium, lanthanum, and zirconium. The compositions found in the molten steels WC70 and WC72 represent E-BRITE^®Ferritic stainless steel. The molten steels WC71, WC73WC74 and WC75 have the general composition of the standard alloy with the following modifications made by the inventors: WC71 molten steel contains an increased niobium content; the WC73 steel ladle contains niobium and cerium; the WC74 steel water contains niobium, cerium and lanthanum; the WC75 steel contains niobium, cerium, lanthanum and zirconium. In Table 1, useThe "< 0.001" associated with cerium, lanthanum and zirconium indicates that these elements were not intentionally added and that chemical analysis indicates that the alloy lacks any significant amount of these elements. As described below, the inventors performed a standard E-BRITE^®The improvement of the alloy composition significantly improves the microstructural stability, mechanical properties and high temperature creep resistance.

The molten steel of table 1 was cast into ingots and treated prior to testing. Each ingot was cross-rolled at 2200F (1204℃) to spread the ingot into 5 inch (127 mm) wide bars. As is known in the art, cross rolling rolls a metal product in a rolling direction (rollingdirections) of about 90 degrees from the previous rolling direction. The cross rolled bar was then hot rolled through a series of rolling stands at a temperature of at least 2100 ° f (1149 ℃) a sufficient number of times to provide a 0.125 inch (3.18 mm) thick strip. The hot rolled strip was then water quenched, grit blasted, acid dip cleaned, and then cold rolled into a 0.040 inch (1.02 mm) thick strip.

After cold rolling, a sample of strip formed from each of the molten steels in table 1 was retained for recrystallization studies. Each of the remaining strips was annealed in-line (line annealed) at 1980 (1082) F (WC71-WC75 alloy) or 1725 (WC70 alloy) for 30 seconds time-at-temperature (time-at-temperature). After annealing, the strips were descaled by brief immersion in molten sodium salt and then pickled in a mixture of sulfuric acid, nitric acid and hydrogen fluoride. The partially annealed 0.040(1.02 mm) thick material was further cold rolled into foil (0.002 in/0.051 mm thick) for strip life cycle oxidation testing.

The following various tests were performed on fully processed strips formed from each molten steel to determine the microstructural stability, mechanical properties, creep/rupture strength and oxidation resistance of these 6 alloy compositions at temperatures representative of SOFC operation.

I. Study of recrystallization

Microstructural stability of 0.040(1.02 mm) thick strip samples formed from each molten steel that had been previously annealed, pickled and cold rolled was evaluated. Samples formed from each of the melts were annealed in a muffle furnace at 1750-. Then, the (mounted) profile was determined and polished for metallographic analysis. Particle size was evaluated according to ASTM standard E112 both at the sample centerline and near the sample surface. Tables 2 (centerline measurements) and 3 (near sample surface measurements) provide ASTM grain size results. The difference in particle size measurements at two different points on the same sample is indicated, for example, by "7.0/7.5". The larger the number of particle sizes, the smaller the particle size.

TABLE 2 centerline measurements

Molten steel number
							Annealing temperature	WC70	WC71	WC72	WC73	WC74	WC75
1750℉(954℃)1800℉(982℃)1850℉(1010℃)1900℉(1038℃)1950℉(1066℃)2000℉(1093℃)2050℉(1121℃)2100℉(1148℃)	7.0/7.57.57.0/8.06.0/7.54.5/7.03.0/5.53.0/4.02.0/2.5	-----6.5/7.53.0/5.03.0	7.0/7.57.0/7.55.0/6.54.54.0/4.54.02.52.5	-----6.0/6.54.03.5	------4.0/5.03.5/4.0	-----6.5/7.05.0/6.02.0/3.5

As indicated by the results of Table 2, which include measurements taken at the centerline of the sample after annealing, alloys of molten steel WC70 and WC72 with only trace amounts of niobium and rare earth elements readily recrystallized at 1750F (954℃.) and suffered significant grain growth at temperatures of about 1950F (1066℃.) and above. Alloys with more than trace amounts of niobium (liquid steel WC71), niobium and cerium (liquid steel WC73) and niobium, cerium, lanthanum and zirconium (liquid steel WC75) also showed no signs of recrystallization up to about 2000 ° f (1093 ℃). Alloys containing more than trace amounts of niobium, cerium and lanthanum (molten steel WC74) also did not show recrystallization up to about 2050F (1121℃). These results show that the addition of niobium, either alone or in combination with the rare earth element and zirconium, at least retards recrystallization at 200 ° f (93 ℃) compared to the unmodified form of the ferritic alloy.

TABLE 3 measurement at sample surface

Molten steel number
							Annealing temperature	WC70	WC71	WC72	WC73	WC74	WC75
1750℉(954℃)1800℉(982℃)1850℉(1010℃)1900℉(1038℃)1950℉(1066℃)2000℉(1093℃)2050℉(1121℃)2100℉(1148℃)	8.5/9.58.5/9.06.0/7.57.0/7.54.5/7.05.0/5.53.0/4.02.0/2.5	-----8.07.53.0	9.08.58.07.54.0/4.54.02.52.5	-----7.56.57.0	------4.0/5.03.5/4.0	-----7.5/8.07.02.0/3.5

The results shown in table 3, which include the particle size measurements obtained near the surface of the sample after annealing, are quite similar to those in table 2. It should be noted that the samples of molten steel WC71 tested at 1750F (954℃) are not equiaxed microstructures. Samples having a standard ferritic stainless steel composition, molten steels WC70 and WC72, showed recrystallization starting at about 1750F (954℃.) and significant recrystallization was observed at 1950F (1066℃.) and higher. In addition, the ferritic alloy modified by the present inventors did not show recrystallization until more than 1950F (1066 ℃ C.), and the alloy with niobium, cerium and lanthanum (molten steel WC74) did not show signs of recrystallization until 2000F (1093 ℃ C.). Thus, the addition of niobium, either alone or with zirconium and rare earth elements including but not limited to cerium and lanthanum, delays at least 200 ° F (93 ℃) recrystallization.

FIG. 1 graphically illustrates the effect of niobium addition, either alone or in combination with rare earth elements, on the recrystallization of various alloys. As noted in the discussion of tables 2 and 3 above, recrystallization is delayed by at least 200F (93℃) in alloys with added niobium, either alone or in combination with one or more rare earth elements, including cerium, lanthanum, and zirconium.

While not intending to be bound by any particular theory, it appears that the resistance to recrystallization of the modified alloys (including WC73-WC75) is a result of the presence of Laves phase precipitates in the sample. The Laves phase is an intermetallic phase that contributes to wear resistance and severely limits the ductility and impact resistance of the alloy material. Metallographic analysis of the annealed 0.040 inch (1.02 mm) thick material showed that the standard alloy (WC70 liquid steel) contained a small amount of laves phase precipitates, while the modified alloy tested contained significant portions of laves phase dispersed within and at grain boundaries. These precipitates impede grain boundary movement, thus preventing grain growth. Thus, the modified alloy has greater grain size stability than the standard alloy.

Mechanical test (Mechanical Testing)

Tensile specimens were machined from 0.040 inch (1.02 mm) thick annealed steel strip and tested. The high temperature test was carried out according to ASTM E21. The in-line tensile properties, calculated as the average properties of a minimum of two samples of each alloy, are shown in table 4 and figure 2.

TABLE 4

Molten steel	Test temperature	Hardness (Rb)	Yield stress (psi)	Tensile stress (ksi)	Elongation (%)
						WC70	77℉(25℃)1472℉(800℃)1562℉(850℃)1652℉(900℃)	79.0	49,6004,3674,5333,100	76,5006,7675,6004,233	27679876
WC71	77℉(25℃)1472℉(800℃)1562..℉(850℃)1652℉(900℃)	84.0	52,9007,3004,4333,475	80,00010,1606,7005,450	27503056
						WC73	77℉(25℃)1472℉(800℃)1562℉(850℃)1652℉(900℃)	84.4	51,3005,8005,6003,567	79,7008,5207,5675,733	26465058

WC75

77℉(25℃)1472℉(800℃)1562℉(850℃)1652℉(900℃)

84.6

49,3006,5674,9503,433

80,9009,7337,2755,667

23566785

As shown in table 4 and fig. 2, the modified steels (steels WC71, WC73 and WC75) showed higher yield and ultimate tensile strength values at high temperatures, usually at the expense of a slight decrease in elongation (offset by 0.2%). Samples that broke on or off the scale mark were excluded from the average elongation calculation.

As shown in table 4, the yield strength of the modified alloys (molten steel WC71, WC73, and WC75) was greater than that of the standard alloy (molten steel 70) at each test temperature, with one exception. The only anomalous result was found in molten steel WC71 at 1562F (850 ℃).

The tensile strength of the modified alloy is without exception higher than that of the standard alloy at all high temperatures. Typically, the hardness of the alloy is similar to the alloy tensile strength. This is also true in the present situation. Looking at table 4, it can be noted that the modified alloy not only has a higher hardness value than the standard alloy, but also has a greater tensile strength. Thus, the modified alloy possesses mechanical properties superior to those of the standard alloy.

Creep and Stress Rupture test (Creep and Stress Rupture Testing)

Creep is the time-dependent strain that occurs under stress. Creep strain occurring at a decreasing rate is called initial creep; secondary creep, which occurs at a minimum and nearly constant rate; the occurrence at an accelerated rate is called third creep. Creep of SOFC interconnects at high temperatures can cause damage to the cell integrity, leading to gas leakage. Creep strength is the stress that results in a given creep strain in a creep test at a given time in a given constant environment. Standard E-BRITE^®The creep strength of alloys such as those contained in molten steels WC70 and WC72 has been determined to be insufficient at the temperatures and stresses encountered in SOFC applications. However, the modification of the standard alloy of the present invention has been shown to significantly improve creep resistance.

Creep rupture strength is the stress that causes rupture in a creep test at a given time in a given constant environment. Creep-rupture test is a test in which the time for a sample to progressively deform and rupture is both measured. Creep rupture tests were performed using 0.040 inch (1.02 mm) thick materials made from standard alloys (molten steel WC70) and modified alloys (molten steel WC71, WC73 and WC 75). The standard alloy samples were annealed at 1715-. Samples of the three modified alloys were annealed at 1970-. The purpose of this test was to evaluate the effect of alloying additions in the modified alloy on creep strength. Since particle size has proven to be extremely important for creep and creep-rupture resistance, the fact that the modified and unmodified alloys have similar particle sizes (within 1-2ASTM grain size) demonstrates that the observed changes in creep resistance are due to composition and precipitation behavior.

Creep-rupture blank specimens were machined longitudinally from 0.040 inch (1.02 mm) thick annealed strip. Creep rupture tests were performed according to ASTM E139 to determine the time to 1% creep strain (FIGS. 3(a) - (c)), 2% creep strain (FIGS. 4(a) - (c)), and fracture (FIGS. 5(a) - (c)) at 800 deg.C (1472 deg.F), 850 deg.C (1562 deg.F), and 900 deg.C (1652 deg.F) for up to 1000 hours and under applied stress up to 3500 psi. The results are shown in FIGS. 3-5. The data included in fig. 3-5 are provided in tables 5-16 below.

TABLE 5WC70 molten steel, time of 1% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,0002,5002,8003,0003,1003,2003,5001,5001,8002,0002,2002,5007007507508008009001,1001,100	125.0120.03.850.011.06.84.8110.04.023.08.06.0300.03.85.04.04.02.51.01.0

1,3001,500

2.31.0

TABLE 6WC70 molten steel, time of 2% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,0002,5002,8003,0003,1003,2003,5001,5001,8002,0002,2002,5007007507508008009001,1001,1001,3001,500	1000.0320.09.5160.031.015.59.8300.08.839.517.523.0400.015.015.08.08.05.02.02.04.51.5

TABLE 7WC70 molten steel, time to fracture

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,5002,8003,0003,1003,2003,5001,4001,5001,7501,8002,2002,5007007508009001,1001,3001,500	822.577.5537.4160.172.558.2229.5520.3143.7145.948.3106.52205.0326.5177.4156.161.225.137.8

TABLE 8WC71 molten steel, time of 1% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)	2,0002,2002,3002,4502,5003,0001,7001,8002,000	370.0350.087.5185.014.030.092.575.053.0

1652℉(900℃)

2,5001,5001,6001,7001,8002,000

11.366.028.022.07.55.0

TABLE 9WC71 molten steel, time of 2% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,0002,2002,3002,4502,5003,0001,7001,8002,0002,5001,5001,6001,7001,8002,000	650.0505.0156.3285.029.048.0192.5180.0101.021.086.060.033.012.510.0

TABLE 10WC71 molten steel, time to fracture

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)	2,2002,3002,4502,5003,0001,7001,8002,0002,5001,500	954.4379.8662.4239.8131.0372.0652.9287.045.5203.4

1652℉(900℃)

1,6001,6001,7001,8002,000

175.0188.983.037.856.2

TABLE 11WC73 molten steel, time of 1% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,5002,6002,8003,0003,2003,3501,5001,7501,8501,9002,0002,5001,0001,1501,2001,4001,6001,8002,000	210.0200.0120.075.0375.060.0390.0500.0410.0122.036.02.3435.075.035.062.557.06.82.3

TABLE 12WC73 molten steel, time of 2% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)	2,5002,6002,8003,0003,2003,3501,500	355.0365.0161.3127.5380.090.0870.0

1562℉(850℃)1652℉(900℃)

1,7501,8501,9002,0002,5001,0001,1501,2001,4001,6001,8002,000

745.0503.8185.077.05.1742.5137.588.0125.071.013.55.0

TABLE 13WC73 molten steel, time to fracture

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,5002,6002,8003,0003,2003,3501,7501,8501,9002,0002,5001,1501,2001,4001,6001,800	862.4807.2310.3292.4390.2200.0894.3557.5226.5266.139.3316.6270.0270.5132.052.5

2,000

24.5

TABLE 14WC75 molten steel, time of 1% creep strain

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,3502,5002,5502,6502,7502,8003,0001,4001,5002,0002,0502,1502,2002,5001,0001,100l,2001,4001,5001,8002,000	225.O825.0130.050.0145.062.547.08.0400.0360.0102.032.060.019.01125.0105.06.540.027.04.53.5

TABLE 15WC75 molten steel, time of 2% creep strain

Temperature of	Stress (psi)	Time (hours)
				2,3502,550	365.0240.0

1472℉(800℃)1562℉(850℃)1652℉(900℃)

2,6502,7502,8003,0001,4001,5002,0002,0502,1502,2002,5001,0001,1001,2001,4001,5001,8002,000

102.5188.0118.872.517.0665.0550.0140.056.074.048.0315.0152.515.078.042.510.06.5

TABLE 16WC75 molten steel, time to fracture

Temperature of	Stress (psi)	Time (hours)
			1472℉(800℃)1562℉(850℃)1652℉(900℃)	2,3502,5502,6502,7502,8003,0002,0502,1002,2002,5001,1001,2001,4001,5001,8002,000	858.5494.4245.7253.9293.5147.0269.8140.0171.475.6470.064.2180.3131.158.440.4

Studying FIGS. 3-5, the compositional changes did not produce significant differences in creep resistance at the lowest test temperature of 800 deg.C (1472 deg.F.). Raising the temperature to 850 ℃ (1562 ° f) creates some difference between the creep resistance of the standard and modified alloys. Testing at 900 deg.C (1652 deg.F) showed a clear separation of creep strength properties between the different alloys. The modified alloys (liquid steels WC71, WC73 and WC75) generally have an increased creep resistance at higher test temperatures compared to the standard alloy (liquid steel WC 70). The results were consistent when tested at high test temperatures for the time to determine 1% creep, 2% creep and rupture, indicating that the modified alloy has superior creep resistance to the standard alloy. For example, based on the test data, it can be seen that the modified alloy exhibits a creep rupture strength of at least 1000psi at 400 hours at 900 ℃ (1652 ° f); a time to 1% creep strain of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a time to 2% creep strain at 1000psi load, 900 ℃ (1652 ° F) of at least 200 hours. In contrast, based on the test data, the standard alloy (WC70) showed creep rupture life of only about 156 hours at 900 ℃ (1652 ° F), 900psi lower stress. A standard alloy of molten steel WC70 showed a creep strain of 1% in 2.5 hours at 900 ℃ (1652 ° F), 900psi load; at 900 deg.C (1652 deg.F), 900psi load, the time to 2% creep strain is only 5.0 hours. These differences illustrate the significant improvement in creep and fracture resistance caused by alloy modification.

The increased creep resistance of the modified alloy in high temperature environments makes the alloy suitable for use in SOFC's and other high temperature applications.

Oxidation test

The isothermal oxidation performance of different alloys (molten steel WC70-WC75) was investigated. Two alloy samples were exposed to 800 ℃ (1472 ° f) and 900 ℃ (1652 ° f) for 500 hours. The sample is first degreased to remove grease and oil from the surface of the metal. Then, the sample was weighed and driedPlaced in an alumina crucible and exposed to a set long time at high temperature in a laboratory ambient atmosphere in a box furnace built with a solid hearth. Periodically, the samples were removed, weighed and placed back in the test kiln. The measured weight change was divided by the area of the sample to give the time-dependent change in specific gravity (mg/cm)²) Curve (c) of (d).

As shown in FIG. 6, the isothermal oxidation test at 800 ℃ (1472 ° F) produced similar weight changes for all samples. After 336 hours, the standard alloy steel (steel WC70) showed a slightly higher weight gain. However, after 500 hours, the weight gain was similar for all samples. The offset data point for the 336 hour sample of molten steel WC70 in fig. 5 may be caused by erroneous measurements when there is no evidence of significant scale spallation (separation of particles from the surface in the form of flakes). All samples showed a uniform charcoal grey without signs of discoloration or local erosion.

As shown in FIG. 7, the isothermal oxidation test at 850 deg.C (1562 deg.F) was limited to 3 samples of molten steels WC70, WC71, and WC 74. The sample made from molten steel WC71 with only the niobium content changed relative to the standard alloy showed a higher weight gain than the standard alloy (molten steel WC70) or the alloy modified by adding niobium, cerium and lanthanum (molten steel WC 74). After 168 hours, this difference was discernible, and after 500 hours, this difference became more pronounced.

As depicted in fig. 8, exposure to 900 ℃ (1652 ° f) shows similar results to those seen at lower temperatures. Also, the alloy modified by adding niobium only (molten steel WC71) had a slightly higher weight gain than the standard alloy (molten steel WC70) or the modified alloy with increased niobium, cerium and lanthanum (molten steel WC 74). The sample formed a relatively uniform charcoal gray scale with a light green undertone. Some evidence of local discoloration is evident.

The parabolic rate constant (parabo1ic rate constant) is a measure of the rate of oxidation. The constants summarize the overall weight change curve at a given temperature. The parabolic rate equation is of the form Δ M/a-k_pV, where Δ M/a is specific gravity change in mg/cm²，t＝Time, k_pParabolic rate constant. The parabolic rate constants resulting from the 500 hour oxidation exposure test performed on each alloy are listed in table 17.

TABLE 17

Rate constant (g)/cmh)
							Exposure temperature	WC70	WC71	WC72	WC73	WC74	WC75
1472℉/800℃1652℉/900℃	-13.5-12.1	-13.7-11.9	-13.8-12.5	-13.9-12.2	-13.8-12.2	-13.7-12.2

For the exposure tests performed, the calculated values were substantially within the scatter points (within ± 0.25 on a logarithmic scale).

Oxidation under thermal cycling conditions is generally more severe than oxidation at constant temperature. In general, there is a significant difference in the thermal expansion coefficients of the oxide and the metal. This can lead to a high degree of stress during thermal cycling, causing premature detachment of the protective oxide layer, known as spalling. The oxide spalls off exposing the as-received metal and then reoxidizes rapidly. Samples of the modified alloy molten steel were rolled into 0.002 inch (0.051 mm) thick foils and stamped (stamped) into cyclic oxidation test samples. These samples were then tested. The sample was heated for 2 minutes using an electric current and then rapidly cooled to room temperature. After 2 minutes at ambient, the sample was cycled back to the test temperature. The total number of cycles before filament breakage (filament breakage) due to through thickness oxidation was used as a measure of oxidation resistance under cycling conditions. Two samples were tested at 2100 deg.F (1149 deg.C), 2200 deg.F (1204 deg.C) and 2300 deg.F (1260 deg.C). The results depicted in fig. 9 show that the molten steel modified with niobium addition (molten steel WC71) shows poor resistance to cyclic oxidation and shows a general trend in isothermal oxidation tests (general tend). (CTF in FIG. 9 is "cycles to failure" of fatigue fracture.)

Coefficient of thermal expansion V

As discussed above, CTE is a key property of fuel cell interconnect materials. If the CTE's of the interconnects are too mismatched with the ceramic components of the fuel cell, the mechanical integrity of the cell, particularly the seal between the cell layers, can be compromised. Thus, in the stainless steel of the present invention, the conventional electrolyte CTE in SOFCs is in the range of about 25% higher to about 25% lower than the stabilized zirconia CTE at 20(68 ° f) to 1000 ℃ (1832 ° f). For the reasons stated above, it is preferred that the CTE of the steel be at least as great as the CTE of stabilized zirconia at 20 ℃ (68 ° F) to 1000 ℃ (1832 ° F), and can be as high as about 25% greater than the CTE of stabilized zirconia.

Testing of conventional E-BRITE^®Samples of the alloys were tested to determine the average CTE. E-BRITE^®The alloy (UNSS44627) comprises, on a weight% basis, up to 0.010 carbon, up to 0.40 manganese, up to 0.020 phosphorus, up to 0.020 sulfur, up to 0.40 silicon, 25.0-27.5 chromium, up to 0.50 nickel, 0.75-1.50 molybdenum, up to 0.015 nitrogen, up to 0.20 copper, 0.05-0.20 niobium and up to 0.50 (nickel + copper). The CTE test results are given in table 18 below.

Watch 18

		Laboratory A (× 10)/℃)		Laboratory B (× 10)/℃)
						Test temperature		Longitudinal direction	Transverse direction	Longitudinal direction	Transverse direction
(℃)	(℉)
		401002003004005006007008009009981000	104212392572752932111212921472165218101832	-9.9310.3810.7310.9311.1611.3511.6812.1812.5813.02-	-9.289.8110.210.5310.8711.0611.3311.7612.24-12.74	8.219.379.9810.3410.610.8911.0911.4511.9312.53-13.05	10.2210.0410.2510.5410.7911.0611.311.6112.0612.58-13.12

E-BRITE^®The low carbon limitations of the alloy and the limitations on nickel and copper (both individually and in admixture) are relaxed in the alloys of the present invention as generally described herein. It is believed that this change has no significant effect on the thermal expansion properties of the alloy. It is also considered that at least one of niobium, titanium and tantalum is contained in the alloy of the present invention so as to satisfy the following relational expression

0.4≤(％Nb+％Ti+1/2(％Ta))≤1

The CTE of the alloy is not substantially affected. All CTE values in Table 18 are within the range of about 25% higher to about 25% lower of 11X 10-6/° C, which is the approximate CTE of YSZ at 20 ℃ (68 ° F) to 1000 ℃ (1832 ° F).

Thus, the above test results demonstrate that the ferritic stainless steels of the present invention have improved high temperature mechanical properties relative to standard ferritic stainless steels. For example, as opposed to E-BRITE^®Ferritic stainless steel, the stainless steel of the invention shows improved stability of the microstructure, improved mechanical properties and greater creep resistance at high temperatures.

Any suitable conventional melting and refining practices may be used to prepare the slab or ingot of the steel of the present invention. The slab or ingot may be further processed by conventional means to form a product such as a strip, sheet or plate, solution annealed and optionally precipitation heat treated. For contemplated fuel cell applications, the steel may be precipitation heat treated at the use temperature (about 1600-. When lower service temperatures are involved, it may be desirable to precipitation harden the steel by exposing the steel to temperatures of about 1600F (871℃) for a time sufficient to suitably strengthen the material.

The steel can be processed to include components for SOFC's that include a stabilized zirconia-containing electrolyte. Such components include separators and interconnects for SOFC's that include a stabilized zirconia-containing electrolyte. The steel may also be processed into components for oxygen sensor devices comprising stabilized zirconia or into articles for use in other high temperature applications such as high temperature furnace hardware and equipment for processing molten copper and other molten metals. By way of example, SOFC's comprising the ferritic stainless steels of the present invention may include a ceramic anode, a ceramic cathode, and a stabilized zirconia electrolyte intermediate the anode and cathode. SOFC's may also include at least one of an interconnect and a separator comprising the present ferritic stainless steel and disposed adjacent to the ceramic electrolyte.

It should be understood that the foregoing description illustrates those aspects of the invention relevant to a clear understanding of the invention. Certain aspects of the present invention that would be apparent to those of ordinary skill in the art to simplify the present description are not presented so as to facilitate a better understanding of the present invention. While the invention has been described in conjunction with certain embodiments, it is contemplated that many modifications and variations of the present invention will be apparent to those of ordinary skill in the art in view of the foregoing description. It is intended that all such variations and modifications of the invention be covered by the foregoing description and the following claims.

Claims (45)

1. A ferritic stainless steel, comprising:

greater than 25 weight percent chromium;

0.75 to less than 1.5 weight percent molybdenum;

up to 0.05 wt% carbon; and

at least one of niobium, titanium, and tantalum, wherein the sum of the weight percentages of niobium, titanium, and tantalum satisfies the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤1，

Wherein the steel comprises no more than 0.5 wt.% titanium, the steel has a coefficient of thermal expansion in the range of about 25% higher to about 25% lower than the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f), and wherein the steel exhibits at least one creep property selected from the group consisting of: a creep rupture strength at 900 ℃ (1652 ° F) of at least 1000 psi; a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° F).

2. The ferritic stainless steel of claim 1, wherein the coefficient of thermal expansion of the steel is at least as great as the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

3. The ferritic stainless steel of claim 1, wherein the coefficient of thermal expansion of the steel is within the range of about 25% higher to about 25% lower of the coefficient of thermal expansion of yttria-stabilized zirconia at 20 ℃ (68 ° f) -1000 ℃ (1832 ° f).

4. The ferritic stainless steel of claim 1, wherein the steel comprises not greater than 0.005 weight percent carbon.

5. The ferritic stainless steel of claim 1 further comprising at least one element selected from the group consisting of up to 0.1 weight percent cerium, up to 0.05 weight percent lanthanum, and up to 0.05 weight percent zirconium.

6. The ferritic stainless steel of claim 1, wherein the steel comprises not greater than 35 weight percent chromium.

7. The ferritic stainless steel of claim 1 wherein the sum of the weight percentages of niobium, titanium, and tantalum satisfies the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

8. A ferritic stainless steel, comprising:

25 up to 35% by weight of chromium;

0.75 to less than 1.5 weight percent molybdenum;

up to 0.005 wt.% carbon;

at least one of niobium, titanium, and tantalum, wherein the steel contains no more than 0.50 weight percent titanium, and the sum of the weight percentages of niobium, titanium, and tantalum satisfies the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75，

Wherein the steel has a coefficient of thermal expansion in the range of about 25% higher to about 25% lower than the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° F) to 1000 ℃ (1832 ° F), and wherein the steel exhibits at least one creep property selected from the group consisting of: a creep rupture strength at 900 ℃ (1652 ° F) of at least 1000 psi; a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° F).

9. The ferritic stainless steel of claim 8, wherein the coefficient of thermal expansion of the steel is at least as great as the coefficient of thermal expansion of zirconia stabilized at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

10. The ferritic stainless steel of claim 8, wherein the coefficient of thermal expansion of the steel is at least as great as the coefficient of thermal expansion of yttria-stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

11. A method for making a ferritic stainless steel having a coefficient of thermal expansion in the range of from about 25% higher to about 25% lower than the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f) and having at least one creep property selected from the group consisting of: a creep rupture strength at 900 ℃ (1652 ° F) of at least 1000 psi; a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° f), the method comprising:

there is provided a ferritic stainless steel comprising greater than 25 weight percent chromium, 0.75 to less than 1.5 weight percent molybdenum, up to 0.05 weight percent carbon, and at least one of niobium, titanium, and tantalum, wherein the steel comprises no greater than 0.50 weight percent titanium, and the sum of the weight percentages of niobium, titanium, and tantalum satisfies the following relationship

0.5 percent to less than or equal to (Nb +% Ti +1/2 percent (Ta)) < 1; and

solution annealing the steel.

12. The method of claim 11, further comprising hardening the steel by precipitation heat treating the steel.

13. The method of claim 11, wherein solution annealing the steel comprises heating the steel at a temperature at least above the predetermined service temperature of the steel and 1600 ° f (871 ℃).

14. The method of claim 11, wherein the coefficient of thermal expansion of the steel is at least as great as the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

15. The method of claim 11, wherein the coefficient of thermal expansion of the steel is within a range of about 25% higher to about 25% lower of the coefficient of thermal expansion of yttria-stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

16. The method of claim 11, wherein the steel comprises no more than 0.005 wt.% carbon.

17. The method of claim 11, wherein the steel further comprises at least one element selected from the group consisting of: up to 0.1 wt.% cerium, up to 0.05 wt.% lanthanum and up to 0.05 wt.% zirconium.

18. The method of claim 11, wherein the steel comprises no more than 35 wt% chromium.

19. The method of claim 11, wherein the sum of the weight percentages of niobium, titanium, and tantalum in the stainless steel satisfies the relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

20. The method of claim 11, wherein the steel comprises 25 up to 35 weight percent chromium, 0.75 up to 1.5 weight percent molybdenum, up to 0.005 weight percent carbon, and at least one of niobium, titanium, and tantalum, wherein the steel comprises no more than 0.50 weight percent titanium, and the sum of the weight percent niobium, titanium, and tantalum satisfies the relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

21. A method for preparing an article selected from a component for a solid oxide fuel cell comprising a stabilized zirconia electrolyte and a component for an oxygen sensor device comprising a stabilized zirconia, the method comprising: providing a ferritic stainless steel comprising greater than 25 weight percent chromium; 0.75 to less than 1.5 weight percent molybdenum; up to 0.05 wt% carbon; and at least one of niobium, titanium, and tantalum, wherein the weight percentages of niobium, titanium, and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤1，

And wherein the steel comprises no more than 0.5 wt.% titanium, has a coefficient of thermal expansion in the range of from about 25% higher to about 25% lower than the coefficient of thermal expansion of stabilized zirconia at 20 ℃ to 1000 ℃, and has at least one creep property selected from the group consisting of: a creep rupture strength of at least 1000psi at 900 ℃; a 1% creep strain time of at least 100 hours at 900 ℃ under a load of 1000 psi; and a 2% creep strain time of at least 200 hours at 900 ℃ under a load of 1000psi,

solution annealing the steel; and are

The steel is processed into an article.

22. The method of claim 21, wherein the article is selected from the group consisting of a separator for a solid oxide fuel cell comprising a stabilized zirconia electrolyte and an interconnect for a solid oxide fuel cell comprising a stabilized zirconia electrolyte.

23. An article comprising a stabilized zirconia bearing component proximate to a ferritic stainless steel bearing component, the steel comprising:

greater than 25 weight percent chromium;

0.75 to less than 1.5 weight percent molybdenum;

up to 0.05 wt% carbon; and

at least one of niobium, titanium and tantalum, wherein the weight percentages of niobium, titanium and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤1，

Wherein the steel comprises no more than 0.5 wt.% titanium, the steel has a coefficient of thermal expansion in the range of from about 25% higher to about 25% lower than the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f), and has at least one creep property selected from the group consisting of: a creep rupture strength at 900 ℃ (1652 ° F) of at least 1000 psi; a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° F).

24. The article of claim 23, wherein the coefficient of thermal expansion of the steel is at least as great as the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

25. The article of claim 23, wherein the stabilized zirconia is yttria stabilized zirconia.

26. The article of claim 23, wherein the steel comprises no more than 0.005 wt.% carbon.

27. The article of claim 23, wherein the steel further comprises at least one element selected from the group consisting of: up to 0.1 wt.% cerium, up to 0.05 wt.% lanthanum and up to 0.05 wt.% zirconium.

28. The article of claim 23, wherein the steel has niobium, titanium, and tantalum weight percentages that satisfy the relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

29. The article of claim 23, wherein the steel comprises not greater than 35 wt% chromium.

30. The article of manufacture of claim 23, wherein the steel comprises:

25 up to 35% by weight of chromium;

0.75 to less than 1.5 weight percent molybdenum;

up to 0.005 wt.% carbon; and

at least one of niobium, titanium, and tantalum, wherein the steel contains not more than 0.50 wt% titanium, and the weight percentages of niobium, titanium, and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

31. The article of claim 23, wherein the article is an element selected from the group consisting of: elements for solid oxide fuel cells containing stabilized zirconia electrolytes and elements for oxygen sensor devices containing stabilized zirconia.

32. The article of claim 23, wherein the article is selected from the group consisting of a separator for a solid oxide fuel cell comprising a stabilized zirconia electrolyte and an interconnect for a solid oxide fuel cell comprising a stabilized zirconia electrolyte.

33. A solid oxide fuel cell comprising:

an anode;

a cathode;

an electrolyte comprising stabilized zirconia, wherein the electrolyte is between an anode and a cathode; and

an interconnect providing a current path from the anode, the interconnect comprising ferritic stainless steel, comprising

More than 25% by weight of chromium,

0.75 to a maximum of 1.5% by weight of molybdenum,

up to 0.05 wt.% carbon, and

at least one of niobium, titanium and tantalum, wherein the weight percentages of niobium, titanium and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤1，

Wherein the steel comprises no more than 0.5 wt.% titanium, the steel has a coefficient of thermal expansion in the range of about 25% higher to about 25% lower than the coefficient of thermal expansion of stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f), and exhibits at least one creep property selected from the group consisting of: a creep rupture strength at 900 ℃ (1652 ° F) of at least 1000 psi; a 1% creep strain time of at least 100 hours at 1000psi load, 900 ℃ (1652 ° F); and a 2% creep strain time of at least 200 hours at 1000psi load, 900 ℃ (1652 ° F).

34. The solid oxide fuel cell of claim 33, wherein the steel has a coefficient of thermal expansion at least as great as the coefficient of thermal expansion of the stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

35. The solid oxide fuel cell of claim 33, wherein the steel has a coefficient of thermal expansion at least as great as the coefficient of thermal expansion of yttria-stabilized zirconia at 20 ℃ (68 ° f) to 1000 ℃ (1832 ° f).

36. The solid oxide fuel cell of claim 33, wherein the steel comprises:

25 up to 35% by weight of chromium;

0.75 to less than 1.5 weight percent molybdenum;

up to 0.005 wt.% carbon; and

at least one of niobium, titanium, and tantalum, wherein the steel contains not more than 0.50 wt% titanium, and the weight percentages of niobium, titanium, and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

37. The solid oxide fuel cell of claim 33, wherein the weight percentages of niobium, titanium, and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

38. The method of claim 21, wherein the weight percentages of niobium, titanium, and tantalum satisfy the following relationship

0.5≤(％Nb+％Ti+1/2(％Ta))≤0.75。

39. The method of claim 21, wherein solution annealing the steel comprises heating the steel at a temperature at least above an intended use temperature of the steel and 1600 ° f.

40. The method of claim 21, wherein the coefficient of thermal expansion of the steel is at least as great as the coefficient of thermal expansion of stabilized zirconia at 20 ℃ -1000 ℃.

41. The method of claim 21, wherein the coefficient of thermal expansion of the steel is within a range of about 25% higher to about 25% lower than the coefficient of thermal expansion of yttria-stabilized zirconia at 20 ℃ -1000 ℃.

42. The method of claim 21, wherein the steel comprises no more than 0.005 wt.% carbon.

43. The method of claim 21, wherein the steel further comprises at least one element selected from the group consisting of: up to 0.1 wt.% cerium, up to 0.05 wt.% lanthanum and up to 0.05 wt.% zirconium.

44. The method of claim 21, wherein the steel comprises not greater than 35 wt% chromium.

45. The article of claim 30, wherein the component is selected from the group consisting of a separator for a solid oxide fuel cell comprising a stabilized zirconia electrolyte and an interconnect for a solid oxide fuel cell comprising a stabilized zirconia electrolyte.