METHOD FOR PRODUCING A COMPONENT MADE OF A NICKEL-CHROMIUM-ALUMINUM ALLOY AND PROVIDED WITH WELD SEAMS

Abstract

In a method for producing, and/or installing into a system, a component with one or more weld seams containing a nickel-chromium-aluminum alloy, with (in wt. %) >18 to 33% chromium, 1.8-4.0% aluminum, 0.01-7.0% iron, 0.001-0.50% silicon, 0.001-2.0% manganese, 0.00-0.60% titanium, respectively 0.0-0.05% magnesium and/or calcium, 0.005-0.12% carbon, 0.0005-0.050% nitrogen, 0.0001-0.020% oxygen, 0.001-0.030% phosphorus, max. 0.010% sulfur, max. 2.0% molybdenum, max. 2.0% tungsten, remainder ≥50% nickel and impurities, the component containing semi-finished products of wrought alloy, after welding, only the weld seams and surrounding heat-affected zones undergo annealing between greater than 980 and 1250° C. for 0.05 minutes-24 hours, then cooling in inert protective atmosphere, moving protective gas or air, where: Cr+Al≥28 and Fp≤39.9 with Fp=Cr+0.272*Fe+2.36*Al+2.22*Si+2.48*Ti+0.374*Mo+0.538*W−11.8*C, Cr, Fe, Al, Si, Ti, Mo, W and C being element wt. % concentrations.

Description

The invention relates to a method for the manufacture of a component, provided with welded seams, from a nickel-chromium-aluminum wrought alloy having excellent high-temperature corrosion resistance, good creep resistance and good processability.

Austenitic nickel-chromium-aluminum wrought alloys having different nickel, chromium and aluminum contents have long been used in furnace construction and in the chemical and petrochemical industry. For this service, a good high-temperature corrosion resistance is required, even in carburizing atmospheres generating “metal dusting”, as are a good hot strength/creep resistance.

In general, it must be pointed out that the high-temperature corrosion resistance of the alloys listed in Table 1 increases with increasing chromium content. All of these alloys form a chromium oxide layer (Cr₂O₃) with an underlying aluminum oxide layer (Al₂O₃), which is more or less closed and has the corresponding aluminum contents. Small additions of strongly oxygen-affine elements such as, for example, yttrium or cerium, improve the oxidation resistance. In the course of service in the area of application for establishment of the protective layer, the chromium content is slowly consumed. The lifetime of the material is therefore prolonged by a higher chromium content, since a higher content of the element chromium, which forms the protective layer, delays the point in time at which the chromium content goes below the critical limit and oxides other than pure chromium oxide (Cr₂O₃) are formed which, for example, may be iron-containing and nickel-containing oxides. A further increase of the high-temperature corrosion resistance can be achieved by additions of aluminum and silicon. Starting from a certain minimum content, these elements form a closed layer underneath the chromium oxide layer and in this way reduce the consumption of chromium.

In carburizing atmospheres (CO, H₂, CH₄, CO₂, H₂O mixtures containing further process-related non-oxidizing constituents), carbon may penetrate into the material and thus lead to formation of internal carbides. These cause a loss of notch impact strength. The melting point may also be lowered to very low values (as low as 350° C.), and transformation processes due to depletion of chromium in the matrix may occur.

A high resistance to carburization is attained by materials having low solubility for carbon and low carbon diffusion rate. Nickel alloys are therefore generally more resistant to carburization than iron alloys, since both the carbon diffusion and the carbon solubility in nickel are lower than in iron. An increase of the chromium content brings about a higher carburization resistance by formation of a protective chromium oxide layer, unless the oxygen partial pressure in the gas is too low for formation of this protective chromium oxide layer. At very low oxygen partial pressures, materials containing silicon or aluminum may be used, since they form a layer of silicon oxide or the even more stable aluminum oxide, which are formed at even much lower oxygen partial pressures in comparison with chromium oxide.

If the carbon activity is higher than 1 (i.e. the gas is not in equilibrium), “metal dusting” may occur in nickel, iron or cobalt alloys. In contact with the gas, the alloys can consume large quantities of carbon. The segregation processes that take place in carbon-supersaturated alloy lead to material damage. In the process, the alloy decomposes into a mixture of metal particles, graphite, carbides and/or oxides. This type of material destruction takes place in the temperature range of approximately 500 to 750° C.

Typical conditions for the occurrence of “metal dusting” are strongly carburizing gas mixtures of CO, H₂ and/or CH₄, such as occur in ammonia synthesis, in methanol plants, in metallurgical processes and even in heat-treatment furnaces among other possibilities.

The resistance to “metal dusting” tends to increase with increasing nickel content of the alloy, but even nickel alloys are not resistant to “metal dusting”.

The chromium and aluminum contents have a significant influence on the corrosion resistance under “metal dusting” conditions (see FIG. 2). Nickel alloys with low chromium content, such as Alloy 600 (N06600), for example (see Table 1), have relatively high corrosion rates under “metal dusting” conditions. Significantly more resistant alloys are the nickel alloys named Alloy 602 CA (N06025) with a chromium content of 25% and an aluminum content of 2.3% as well as Alloy 690 (N06690) with a chromium content of 30% (Hermse, C. G. M. and van Wortel, J. C.: Metal dusting: relationship between alloy composition and degradation rate. Corrosion Engineering, Science and Technology 44 (2009), pp. 182-185). The resistance to “metal dusting” increases with the total of chromium and aluminum (Cr+Al).

The hot strength and creep strength at the indicated temperatures (approximately 500 to 750° C.) are improved by a high carbon content among other possibilities. However, even high contents of solid-solution-strengthening elements such as chromium, aluminum, silicon, molybdenum and tungsten improve the hot strength or the creep strength. In the range of 500° C. to 900° C., additions of aluminum, titanium and/or niobium may improve the strength by precipitation of the γ′ and/or γ″ phase.

Examples of these alloys according to the prior art are listed in Table 1.

Alloys such as Alloy 602 CA (N06025), Alloy 693 (N06693) or Alloy 603 (N06603) have long been known for their excellent corrosion resistance in comparison with Alloy 600 (N06600) or Alloy 601 (N06601), on the basis of the high aluminum content of more than 1.8%. Alloy 602 CA (N06025), Alloy 693 (N06693), Alloy 603 (N06603) and Alloy 690 (N06690) exhibit excellent carburization resistance and “metal dusting” resistance due to their high chromium and/or aluminum contents. At the same time, alloys such as Alloy 602 CA (N06025), Alloy 693 (N06693), or Alloy 603 (N06603) exhibit, on the basis of the high carbon content or aluminum content, an excellent hot strength or creep strength in the temperature range in which “metal dusting” occurs. Alloy 602CA (N06025) and Alloy 603 (N06603) still have an excellent hot strength or creep strength even at temperatures above 1000° C. However, due to the high aluminum contents among other factors, the processability is impaired, wherein the impairment becomes greater with increasing aluminum content (Alloy 693-N06693). The same is true to a greater degree for silicon, which forms low-melting phases with nickel. In Alloy 602 CA (N06025) or Alloy 603 (N06603), the cold formability in particular is limited by a high proportion of primary carbides.

WO 2013/182177 A1 discloses (in weight-%) 24 to 33% chromium, 1.8 to 4.0% aluminum, 0.10 to 7.0% iron, 0.001 to 0.50% silicon, 0.005 to 2.0% manganese, 0.00 to 0.60% titanium, respectively 0.0002 to 0.05% magnesium and/or calcium, 0.005 to 0.12% carbon, 0.001 to 0.050% nitrogen, 0.0001-0.020% oxygen, 0.001 to 0.030% phosphorus, max. 0.010% sulfur, max. 2.0% molybdenum, max. 2.0% tungsten, the rest nickel and the common process-related impurities, wherein the following relationships must be satisfied:

$\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 28\mathrm{and}& \left(2a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}\le 39.9\mathrm{with}& \left(3a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+0.374*\mathrm{Mo}+0.538*W11.8*C& \left(4a\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti, Mo, W and C are the concentrations of the elements in question in mass-%. This nickel-chromium-aluminum alloy has an excellent high-temperature corrosion resistance under highly corrosive conditions, such as, for example, an excellent “metal dusting” resistance, a good corrosion resistance in air, a good phase stability, a good hot strength and a good creep strength with uniformly good processability.

U.S. Pat. No. 6,623,869 B1 discloses a metal material that consists of not more than 0.2% C, 0.01-4% Si, 0.05-2.0% Mn, not more than 0.04% P, not more than 0.015% S, 10-35% Cr, 30-78% Ni, 0.005-4.5% Al, 0.005-0.2% N, and one or both elements comprising 0.015-3% Cu and 0.015-3% Co, with iron representing the rest up to 100%. Therein the total of 40 Si+Ni+5 Al+40 N+10 (Cu+Co) is not less than 50, wherein the symbols of the elements denote the content of the corresponding elements in mass-%. The material has an outstanding corrosion resistance in an environment in which “metal dusting” may occur and it may therefore be used for furnace tubes, tube systems, heat-exchanger tubes and the like in petroleum refineries or petrochemical plants and may markedly improve the lifetime and the safety of the plant.

EP 0 508 058 A1 discloses an austenitic nickel-chromium-iron alloy consisting of (in weight-%) C: 0.12-0.3%, Cr: 23-30%, Fe: 8-11%, Al 1.8-2.4%, Y: 0.01-0.15%, Ti: 0.01-1.0%, Nb: 0.01-1.0%, Zr: 0.01-0.2%, Mg: 0.001-0.015%, Ca: 0.001-0.01%, N: max. 0.03%, Si: max. 0.5%, Mn: max. 0.25%, P: max. 0.02%, S: max. 0.01%, Ni: the rest, including unavoidable smelting-related impurities.

Alloys for furnace construction and the petrochemical industry must be capable of being welded for the manufacture of the individual components of a plant. The individual components may also have to be capable of being welded during installation in the plant. Because of their microstructure, welded seams often have a lower creep strength and/or tend to crack formation. The creep strength can be improved and the tendency to crack formation can be reduced by a heat treatment of the welded seam. Some examples of the prior art are listed in the following.

EP 0 234 200 A1 discloses a method as well as an apparatus for heat treatment of longitudinally seam-welded tubes of austenitic, ferritic or austenitic-ferritic non-rusting steels, wherein the tubes are annealed after the welding of the longitudinal seam. The method is characterized in that the tubes are solution-annealed only partly in the region of the welded seam and of the heat-affected zone, whereas the other regions are heat-treated at lower temperature. For high-alloy molybdenum-containing steels, annealing is carried out at a temperature above 1100° C., preferably above 1250° C., wherein the annealing temperature is maintained for longer than 5 s, preferably for approximately 25 s.

U.S. Pat. No. 3,865,639 A discloses a method for the manufacture of a welded subassembly, the welded parts of which withstand operating temperatures in the range of approximately 900 to 1050° C., wherein the parts consist of at least one high-alloy austenitic steel, which is alloyed with chromium, nickel and/or cobalt and contains less than 60 weight-% iron, up to 0.5 wt % carbon and low concentrations of additional elements such as manganese and silicon. In this case the indicated parts are subjected to a welding, wherein the welding leads to the formation of at least one welded seam that has a solidification front corresponding to a physical and chemical discontinuity in the seam, wherein the welded seam also consists of a high-alloy austenitic steel and then is treated such that the solidification front becomes chemically homogeneous, whereby the carbides of the welded seam are present to a maximum extent and according to a fine uniform distribution in the precipitated state, which has the consequence that the creep strength and the tensile strength of the welded seam are increased to values that are all at least approximately equal to those of the non-welded parts, in that at least the seam and the adjoining regions of the welded parts undergo a homogenizing heat treatment at a temperature between approximately 1100 and 1200° C. for a time period of several minutes to several hours, after which at least the seam and the adjacent regions are cooled to approximately 800° C. at a temperature-lowering rate of approximately 100° C./h and then with an air cooling to room temperature.

U.S. Pat. No. 3,046,167 A describes a method for the heat treatment of welded, hardenable chromium-nickel stainless steels in order then to impart a high level of ductility and toughness to them in the hardened state, wherein the method comprises the successive steps of welding as well as transformation of the welded products by annealing to a structure that is substantially martensitic, but contains some ferrite, after which the structure of the steel is restored, so that it is substantially austenitic, although it contains some ferrite. This is achieved by annealing at a temperature and for a time sufficient for re-austinitization of the martensite and ferrite and for breaking up the original cast structure of the welding. Thereafter back-transformation of the metal is carried out to a structure that is substantially martensitic, together with very little ferrite, followed by hardening of the metal by a suitable heat treatment to a state that is substantially completely martensitic. During processing of hardenable chromium-nickel stainless-steel products selected from the group of steels having an analysis of approximately 17% chromium, 7% nickel, 1% aluminum and the rest iron; approximately 17% chromium, 4% nickel, 3% copper and the rest iron; approximately 15% chromium, 7% nickel, 2% molybdenum, 1% aluminum and the rest iron; and approximately 12% chromium, 8% nickel, 6% molybdenum, 1% aluminum and the rest iron, wherein the method comprises welding of the products under shield gas as the first step, followed by transformation of the welded product by heating to approximately 760 to 954.4° C. and cooling to approximately room temperature to −73.3° C., then tempering of the transformed products at approximately 1065.6° C., then transformation of the products by heating and cooling as described above and finally hardening by heating to approximately 482.2 to 593.3° C.

U.S. Pat. No. 4,168,190 A describes a method and an apparatus for local solution annealing of austenitic stainless steel, which has been partly sensitized, e.g. by a local temperature elevation, without producing a sensitized structure at the thermal boundaries between the locally treated portion and the rest of the material, comprising the steps of (a) rapid heating of the sensitized parts of the material to a temperature at which the carbides pass into solution and (b) rapid quenching of the heated material.

The task underlying the invention consists in conceiving of a method for the manufacture of a component with welded seams from a nickel-chromium-aluminum wrought alloy, wherein the component consists partly of a nickel-chromium-aluminum wrought alloy and/or takes place for the installation of this component in a plant with one or more welded seams.

This task is accomplished by a method for the manufacture of a component with one or more welded seams and/or for installation of a component in a plant with one or more welded seams, which consist of a nickel-chromium-aluminum alloy, containing (in mass-%) more than 18 to 33% chromium, 1.8 to 4.0% aluminum, 0.01 to 7.0% iron, 0.001 to 0.50% silicon, 0.001 to 2.0% manganese, 0.00 to 0.60% titanium, respectively 0.0 to 0.05% magnesium and/or calcium, 0.005 to 0.12% carbon, 0.0005 to 0.050% nitrogen, 0.0001-0.020% oxygen, 0.001 to 0.030% phosphorus, max. 0.010% sulfur, max. 2.0% molybdenum, max. 2.0% tungsten, the rest nickel, greater than or equal to 50%, and the common process-related impurities, wherein the component consists partly or completely of semifinished products of this nickel-chromium-aluminum wrought alloy and, after the welding, only the welded seams from this nickel-chromium-aluminum wrought alloy and the heat-affected zones surrounding the welded seams are subjected, for homogenization of the welded seams and/or for reduction of stresses, to an annealing between higher than 980 and 1250° C. for times of 0.05 minutes to 24 hours, followed by a cooling in stationary shield gas or air, moving (blown) shield gas or air, with the consequence that the creep strength and the creep ductility of the welded seams are improved by this annealing.

wherein the following relationships must be satisfied:

$\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 28\mathrm{and}& \left(1a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}\le 39.9\mathrm{with}& \left(2a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3a\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti, Mo, W and C are the concentrations of the elements in question in mass-%.

Advantageous further developments of the subject matter of the invention can be inferred from the associated dependent claims.

Semifinished products of metal are partly completed products, such as sheet, strip, bar, forgings, seamless or longitudinally seam-welded tube or wire, for example.

A component may be joined together from two or more machined semifinished products and/or other components. The joints may be made, for example, with a fusion welding method, with extra weld filler if applicable. A weld filler may be supplied during welding, e.g. via welding rods or wires. A weld filler may also consist, in the case of corresponding configuration of the geometry of the welded seam, of one part at least of one of the adjoining semifinished products to be welded.

The nickel-chromium-aluminum wrought alloy, abbreviated as Alloy NiCrAl—H, cited in this invention is preferably smelted openly in an electric furnace or an arc furnace, followed by a treatment in a VOD (Vacuum, Oxidizing, Deoxidizing) system or VLF (Vacuum Ladle Furnace) system. However, a smelting and casting in vacuum is also possible. Thereafter the alloy is cast in ingots or electrodes or as a continuous casting for formation of a precursor product. If applicable, the precursor product is then annealed at temperatures between 900 and 1270° C. for 0.1 hours to 70 hours. Furthermore, it is possible to resmelt the alloy additionally one or more times with ESR (Electroslag Remelting) and/or VAR (Vacuum Arc Remelting). Then the alloy is introduced into the desired semifinished product mold. For this purpose, annealing is carried out if necessary at temperatures between 900° C. and 1270° C. for 0.1 hours to 70 hours, followed by hot-forming, if necessary with intermediate annealings between 900° C. and 1270° C. for 0.05 hours to 70 hours. The surface of the material may if necessary be chemically or mechanically stripped for cleaning intermediately (even several times) and/or at the end of the hot forming. Thereafter a cold forming with reduction ratios up to 98% may be carried out if necessary in the desired semifinished product mold, if necessary with intermediate annealings between 800° C. and 1250° C. for 0.05 minutes to 70 hours, if necessary under shield gas, such as, for example, argon or hydrogen, followed by a cooling in air, in the agitated annealing atmosphere or in the water bath. Thereafter a solution annealing is carried out in the temperature range from 800° C. to 1250° C. for 0.05 minutes to 70 hours, if necessary under shield gas, such as, for example, argon or hydrogen, followed by a cooling in air, in the agitated annealing atmosphere or in the water bath. If necessary, chemical and/or mechanical cleanings of the material surface may be carried out intermediately and/or after the last annealing.

Preferably, the solution annealing takes place between the following temperatures:

• • 1000 or >1000 to 1200° C. or <1200° C.
  • 1000 or >1000 to 1175° C. or <1175° C.
  • 1025 or >1025 to 1150° C. or <1150° C.
  • 1050 or >1050 to 1130° C. or <1130° C.
  • 1080 or >1080 to 1130° C. or <1130° C.

Preferably, the solution annealing takes place within the following time intervals:

• • 0.05 minutes to 24 hours
  • 0.05 minutes to 8 hours
  • 0.05 minutes to 4 hours
  • 0.1 minutes to 1 hour
  • 1 minute to 1 hour
  • 5 minutes to 1 hour

The nickel-chromium-aluminum wrought alloy used in this invention can be manufactured and used well in the semifinished product forms of strip, sheet, bar, forgings, wire, longitudinally seam-welded tube and seamless tube.

These semifinished product forms are manufactured with a mean grain size of 30 to 600 μm. Preferred ranges are:

• • 40 μm to 300 μm.
  • 60 μm to 300 μm.
  • 75 μm to 200 μm.

The parts needed for the component are cut or sheared out of the semifinished product, appropriately machined and then joined by means of a fusion welding technique under shield gas.

Optionally, the fusion welding technique may be, for example, one of the following techniques:

• • Tungsten inert gas welding (TIG)
  • Metal inert gas welding (MIG)
  • Metal active gas welding (MAG)
  • Plasma welding
  • Electron-beam welding
  • Laser-beam welding
  • Gas welding or autogenous welding
  • Electrode welding/manual arc welding

The shield gas may preferably be argon or else argon and hydrogen or else argon and nitrogen.

However, even submerged arc welding and other techniques are conceivable.

The range of values for the element chromium lies between 18 and 33%, wherein preferred ranges may be adjusted as follows:

• • 20 or >20 to 33 or <33%
  • 22 or >22 to 33 or <33%
  • 24 or >24 to 33 or <33%
  • 24 or >24 to 32 or <32%
  • 25 or >25 to 32 or <32%
  • 26 or >26 to 32 or <32%
  • 27 or >27 to 32 or <32%
  • 28 or >28 to 32 or <32%
  • 28 or >28 to 31 or <31%
  • 28 or >28 to 30 or <30%
  • 29 or >29 to 31 or <31%

The aluminum content lies between 1.8 and 4.0%, wherein here also, depending on service area of the alloy, preferred aluminum contents may be adjusted as follows:

• • 1.8 or >1.8 to <4.0%
  • 1.8 or >1.8 to 3.2 or <3.2%
  • 2.0 or >2.0 to 3.2 or <3.2%
  • 2.0 or >2.0 to 3.0 or <3.0%
  • 2.0 or >2.0 to 2.8 or <2.8%
  • 2.0 or >2.2 to 2.8 or <2.8%
  • 2.0 or >2.2 to 2.6 or <2.6%

The iron content lies between 0.1 and 7.0%, wherein, depending on the area of application, preferred contents may be adjusted within the following ranges of values:

• • 0.01 to 4.0 or <4.0%
  • 0.01 to 3.0 or <3.0%
  • 0.01 to 2.5 or <2.5%
  • 0.1-2.0 or <2.0%
  • 0.1-1.0 or <1.0%

The silicon content lies between 0.001 and 0.50%. Preferably, Si in the alloy can be adjusted within the range of values as follows:

• • 0.001-0.20 or <0.20%
  • 0.001-0.10 or <0.10%
  • 0.001-0.05 or <0.05%

The same is true for the element manganese, which may be contained in proportions of 0.001 to 2.0% in the alloy. Alternatively, the following range of values is also conceivable:

• • 0.005 to 0.50 or <0.20%
  • 0.005 to 0.20 or <0.20%
  • 0.001 to 0.20 or <0.20%
  • 0.005 to 0.10 or <0.10%
  • 0.005 to 0.05 or <0.05%

The titanium content lies between 0.00 and 0.60%. Preferably, Ti in the alloy can be adjusted within the range of values as follows:

• • 0.001 to 0.60 or <0.60%
  • 0.001 to 0.50 or <0.50%
  • 0.001 to 0.30 or <0.30%
  • 0.001 to 0.25 or <0.25%
  • 0.001 to 0.10 or <0.10%
  • 0.001 to 0.02 or <0.02%
  • 0.01 to 0.10 or <0.10%

Magnesium and/or calcium is/are also present in contents of 0.00 to 0.05%. Preferably, the possibility exists of adjusting these elements in the alloy as follows:

• • 0.00 to 0.03 or <0.03%
  • 0.00 to 0.02 or <0.02%

The alloy contains 0.005 to 0.12% carbon. Preferably this may be adjusted within the range of values in the alloy as follows:

• • 0.01 to <0.12%
  • 0.01 to 0.10 or <0.10%
  • 0.01 to 0.08 or <0.08%
  • 0.01 to 0.05 or <0.05%
  • 0.02 to 0.10 or <0.10%
  • 0.03 to 0.10 or <0.10%

This is true in the same way for the element nitrogen, which is present in contents between 0.0005 and 0.05%. Preferred contents may be specified as follows:

• • 0.001-0.04%.

Furthermore, the alloy contains oxygen in contents between 0.0001 and 0.020%, especially 0.0001 to 0.010%

Furthermore, the alloy contains phosphorus in contents between 0.001 and 0.030%. Preferred contents may be specified as follows:

• • 0.001-0.020%

The element sulfur is present in the alloy as follows:

• • max. 0.010%

Molybdenum and tungsten are contained individually or in combination in the alloy with a content of respectively at most 2.0%. Preferred contents may be specified as follows:

Mo max. 1.0%
W  max. 1.0%
Mo max. <1.0%
W  max. <1.0%
Mo max. <0.50%
W  max. <0.50%
Mo max. <0.10%
W  max. <0.10%
Mo max. <0.05%
W  max. <0.05%

Nickel is the rest. Preferably, the rest can be specified as follows:

• • >50% or ≥50%
  • >55% or ≥55%
  • >60% or ≥60%
  • >65% or ≥65%
  • >67% or ≥67%

For highly corrosive conditions, but especially for a good “metal dusting” resistance, it is advantageous when the following relationship is satisfied between Cr and Al.

$\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 28& \left(1a\right)\end{array}$

wherein chromium and aluminum are the concentrations of the elements in question in mass-%. Preferred ranges may be adjusted as follows:

$\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 29& \left(1b\right)\end{array}$ $\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 30& \left(1c\right)\end{array}$ $\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 31& \left(1d\right)\end{array}$

Beyond this, the following relationship must be satisfied in order that an adequate phase stability is ensured:

$\begin{array}{cc}\mathrm{Fp}\le 39.9\mathrm{with}& \left(2a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3a\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti, Mo, W and C are the concentrations of the elements in question in mass-%.

Preferred ranges may be adjusted as follows:

$\begin{array}{cc}\mathrm{Fp}\le 38.4& \left(2b\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}\le 36.6& \left(2c\right)\end{array}$

Optionally, the element niobium may be adjusted to contents of 0.0 to 1.10% in the alloy. Preferably, niobium may be adjusted within the range of values in the alloy as follows:

• • 0.001 to <1.10%
  • 0.001 to 1.0 or <1.0%
  • 0.001 to 0.70 or <0.70%
  • 0.001 to 0.50 or <0.50%
  • 0.001 to 0.30 or <0.30%
  • 0.01 to 0.30 or <0.30%

If niobium is contained in the alloy, the formula (3a) must be supplemented by a term for niobium as follows:

$\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+1.26*\mathrm{Nb}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3b\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti, Nb, Mo, W and C are the concentrations of the elements in question in mass-%.

Optionally, the zirconium content may lie between 0.0 and 0.20%. Preferably, zirconium may be adjusted within the range of values in the alloy as follows:

• • 0.0 to 0.15 or <0.15%
  • 0.0 to 0.10 or <0.10%
  • 0.001 to 0.07 or <0.07%
  • 0.001 to 0.04 or <0.04%
  • 0.01 to 0.04 or <0.04%

Optionally, the element yttrium may be adjusted to contents of 0.0 to 0.20% in the alloy. Preferably, yttrium may be adjusted within the range of values in the alloy as follows:

• • 0.0 to 0.15 or <0.15%
  • 0.0 to 0.10% or <0.10%
  • 0.0 to 0.08 or <0.08%
  • 0.001 to <0.045%
  • 0.01 to 0.04 or <0.04%

Optionally, the element lanthanum may be adjusted to contents of 0.0 to 0.20% in the alloy. Preferably, lanthanum may be adjusted within the range of values in the alloy as follows:

• • 0.0 to 0.15 or <0.15%
  • 0.0 to 0.10 or <0.10%
  • 0.0 to 0.08 or <0.08%
  • 0.001 to 0.04 or <0.04%
  • 0.01 to 0.04 or <0.04%

Optionally, the element cerium may be adjusted to contents of 0.0 to 0.20% in the alloy. Preferably, cerium may be adjusted within the range of values in the alloy as follows:

• • 0.0 to 0.15 or <0.15%
  • 0.0 to 0.10% or <0.10%
  • 0.0 to 0.08 or <0.08%
  • 0.001 to 0.04 or <0.04%
  • 0.01 to 0.04 or <0.04%

Optionally, in case of simultaneous addition of cerium and lanthanum, cerium mixed metal may also be used in contents of 0.0 to 0.20%. Preferably, cerium mixed metal may be adjusted within the range of values in the alloy as follows:

• • 0.0 to 0.15 or <0.15%
  • 0.0 to 0.10% or <0.10%
  • 0.0 to 0.08 or <0.08%
  • 0.001 to 0.04 or <0.04%
  • 0.01 to 0.04 or <0.04%

Optionally, the element hafnium may be adjusted to contents of 0.0 to 0.20% in the alloy. Preferably, hafnium may be adjusted within the range of values in the alloy as follows:

• • 0.0 to 0.15 or <0.15%.
  • 0.0 to 0.10 or <0.10%.
  • 0.001 to 0.07 or <0.07%.
  • 0.001 to 0.04 or <0.04%
  • 0.01 to 0.04 or <0.04%.

Optionally, 0.001 to 0.60% tantalum may also be contained in the alloy. Preferred tantalum contents may be specified as follows:

• • 0.001 to 0.50 or <0.50%
  • 0.001 to 0.40 or <0.40%
  • 0.001 to 0.30 or <0.30%
  • 0.001 to 0.20 or <0.30%
  • 0.001 to 0.10 or <0.10%
  • 0.01 to 0.10 or <0.10%

Optionally, the element boron may be contained in the alloy as follows:

Boron 0.0001-0.008%

Preferred contents may be specified as follows:

Boron 0.0005-0.008%
Boron 0.0005-0.004%

Furthermore, the alloy may if necessary contain between 0.0 and 5.0% cobalt, which beyond this may still be limited as follows:

• • 0.01 to 5.0 or <5.0%
  • 0.01 to 2.0 or <2.0%
  • 0.10 to 2.0 or <2.0%
  • 0.01 to 0.5 or <0.5%
  • 0.01 to 0.1 or <0.1%

Furthermore, if necessary at most 0.5% copper may be contained in the alloy.

Beyond this, the content of copper may be limited as follows:

Cu max. <0.20 or 0.20%
Cu max. <0.10 or 0.10%
Cu max. <0.05 or 0.05%
Cu max. <0.015%

If copper is contained in the alloy, the formula (3a) must be supplemented by a term for copper as follows:

$\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+0.477*\mathrm{Cu}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3c\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti, Cu, Mo, W and C are the concentrations of the elements in question in mass-%.

If niobium and copper are contained in the alloy, the formula (3a) must be supplemented by a term for niobium and a term for copper as follows:

$\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+1.26*\mathrm{Nb}+0.477*\mathrm{Cu}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3d\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti Nb, Cu, Mo, W and C are the concentrations of the elements in question in mass-%,

Furthermore, optionally at most 0.5% vanadium may be contained in the alloy.

Beyond this, the content of vanadium may be limited as follows:

V max. <0.10%

Finally, as impurities, the elements lead, zinc and tin may also be present in contents as follows:

• • Lead max. 0.002%
  • Zinc max. 0.002%
  • Tin max. 0.002%

The annealing of the welded seams and of the heat-affected zones may be carried out, for example, by heating mats, infrared radiators, lasers or inductive electric heating. Optionally, partial regions of the rest of the construction may be cooled at the same time.

For homogenization of the welded seams and/or for reduction of stresses of an annealing, it is likewise possible, after the welding, to subject the entire component to an annealing between 90° and 1250° C. for times from 0.05 minutes to 24 hours, followed by a cooling in stationary shield gas or air, moving (blown) shield gas or air, with the consequence that the creep strength and the creep ductility of the welded seams are improved by this annealing (if necessary, after the solution annealing of the semifinished product).

Furthermore, it is possible, after the welding and the homogenization of the welded seams, to repeat this annealing, either by annealing of the entire component containing the welded seams or by an annealing only of the welded seams and of the heat-affected zones, optionally by annealing of the entire component containing the welded seams or by an annealing only of the welded seams and heat-affected zones. For this purpose, the sequence of a partial annealing only of welded seams and heat-affected zones and the annealing of the entire component is optional.

Preferably, an annealing after the welding takes place between the following temperatures:

• • 1000 or >1000 to 1200 or <1200° C.
  • 1000 or >1000 to 1175 or <1175° C.
  • 1025 or >1025 to 1150 or <1150° C.
  • 1050 or >1050 to 1130 or <1130° C.
  • 1080 or >1080 to 1130 or <1130° C.

Preferably, the annealing after the welding takes place in the following time intervals:

• • 0.05 minutes to 16 hours
  • 0.05 minutes to 8 hours
  • 0.05 minutes to 4 hours
  • 0.1 minutes to 1 hour
  • 1 minute to 1 hour
  • 5 minutes to 1 hour

The shield gas for the annealing after welding may, if not under air, preferably consist of the following gases:

• • Argon
  • Argon—air mixtures
  • Hydrogen
  • Argon—hydrogen mixtures
  • Argon—nitrogen mixtures

The improvement of the creep strength and of the creep ductility of the welded seams takes place in particular in the range of γ′ formation, which comprises the temperature range lower than or equal to 750° C.

After the annealing of the component containing the welded seams or after the annealing only of the welded seams and of the heat-affected zones of the component, or of the welded seams and of the heat-affected zones of the installation in the plant, the surface may optionally be cleaned or machined by brushing, pickling, abrasive blasting, grinding, turning, scalping and/or milling. Optionally, one or more of such machining operations may also be carried out immediately after the welding. In particular, a material-removing machining by grinding, turning, scalping, milling after the last annealing improves the corrosion resistance, especially the “metal dusting” resistance, of the annealed surfaces, especially of the welded seams and of the heat-affected zones.

After the grinding of the welded seam and of the heat-affected zones, it is advantageous when roughness values Ra of 0.01 to 15 μm are attained, since this improves the corrosion resistance and especially the metal dusting resistance and raises them almost to the value of the parent material.

The components manufactured according to the invention are intended to be used preferably in regions in which highly corrosive conditions prevail, such as, for example, strongly carburizing conditions, atmospheres that cause “metal dusting”, as in the case, for example, of components in the petrochemical industry. Beyond that, they are also suitable for furnace construction.

EXAMPLES

Manufacture of a Component from Semifinished Product

FIG. 1a left shows sketches of the semifinished products in the form of sheet (1), strip (2), bar (3), tube (4), wire and weld filler in wire form (5) in top view and cross section. FIG. 1a right shows, for example, how a product or precursor product (7) is formed, in that two components (1a, 1b), for example, are cut out of the semifinished product sheet and chamfered (ii) and then joined (iv) by fusion welding (iii) with a weld filler in the form of wire (5) by means of a V-groove seam (6a). FIG. 1b left shows a further example for the formation of a product or precursor product (7), in which two tubular components (4a, 4b), for example, are cut out of the semifinished product tube and chamfered (ii) and then joined (iv) by fusion welding (iii) with a weld filler in the form of wire (5) by means of a V-groove seam (6b). FIG. 1b middle shows yet another example for the formation of a product or precursor product (7), in which a hole is milled (ii) through a sheet semifinished product or a sheet component (1c), a tube semifinished product or a tube component (3c) is inserted therein and joined (iv) by fusion welding (iii) with a weld filler in the form of wire (5) by means of a fillet weld (6c). FIG. 1b right shows a further example for the formation of a product or of a precursor product (7), in which two strip portions (2a, 2b) are sheared out from the semifinished product strip (ii), machined to make the edges fit and then joined (iv) abuttingly (9) by fusion welding (iii) with one part of the edges as weld filler (6d).

Tests Performed

The phases occurring in equilibrium were calculated for the various alloy variants with the JMatPro program of Thermotech. The TTNI7 database of Thermotech for nickel alloys was used as the database for the calculations.

The creep strength is determined in an uninterrupted uniaxial creep test with elongation measurement under tensile load according to DIN EN ISO 204. For this purpose, the specimen is mounted in a creep-testing machine and subjected to a constant test force. In the process, the time to fracture t_u and the creep elongation at break A_u^b are determined. The time to rupture is a measure for the creep strength and the creep elongation at break is a measure for the creep ductility. The tests were performed on round specimens with a diameter of 10 mm in the measurement region and an initial reference length L_r0 of 50 mm. The sampling took place in a manner transverse relative to the direction of forming of the semifinished product.

Description of the Properties

The nickel-chromium-aluminum alloy named Alloy NiCrAl—H used in this invention has, besides an excellent corrosion resistance under highly corrosive conditions, in this case, for example, an excellent “metal dusting” resistance, a good phase stability and creep strength.

Phase Stability

In the nickel-chromium-aluminum-iron system having additions of titanium and/or niobium, various embrittling TCP (topologically close packed) phases can be formed, depending on alloying contents, such as, for example, the Laves phase, sigma phase or μ-phase or even the embrittling η-phase or ε-phase. The calculation of the equilibrium phase proportions as a function of temperature for N06690, Batch 111389, for example (see Table 2 for the compositions used here), theoretically shows the formation of α-chromium (BCC phase in FIG. 3) below 720° C. (T_{s BCC}) in high quantitative proportions. The formation of this phase is generally made difficult, since it is analytically very different from the parent material. However, if the formation temperature T_{s BCC} of this phase is very high, it can definitely occur, as described, for example, in “E. Slevolden, J. Z. Albertsen. U. Fink, “Tjeldbergodden Methanol Plant: Metal Dusting Investigations,” Corrosion/2011. paper no. 11144 (Houston, TX: NACE 2011), p. 15″ for a variant of Alloy 693 (UNS 06693). FIG. 4 and FIG. 5 show the phase diagrams of the Alloy 693 variants (from U.S. Pat. No. 4,882,125 Table 1) for Alloy 3 and Alloy 10 from Table 2. This phase is brittle and leads to an undesired embrittlement of the material. Alloy 3 has a formation temperature T_{s BCC} of 1079° C. and Alloy 10 of 939° C. In “E. Slevolden, J. Z. Albertsen. U. Fink, “Tjeldbergodden Methanol Plant: Metal Dusting Investigations,” Corrosion/2011. paper no. 11144 (Houston, TX: NACE 2011), p. 15″ the exact analysis of the alloy in which α-chromium (BCC) forms is not described. It is to be assumed, however, that α-chromium (BCC phase) can be formed among the examples cited in Table 2 for Alloy 693 or analyses that theoretically have the highest formation temperatures Is BCC (such as, for example, Alloy 10). In a corrected analysis (with reduced formation temperature T_{s BCC}), α-chromium was subsequently still observed only close to the surface in “E. Slevolden, J. Z. Albertsen. U. Fink, “Tjeldbergodden Methanol Plant: Metal Dusting Investigations,” Corrosion/2011. paper no. 11144 (Houston, TX: NACE 2011), p. 15″. In order to avoid the occurrence of such an embrittling phase, the formation temperature T_{s BCC} in the nickel-chromium-aluminum alloy (Alloy NiCrAl—H) used in this invention should be lower than or equal to 939° C.—the lowest formation temperature T_{s BCC} among the examples for Alloy 693 in Table 2 (from U.S. Pat. No. 4,882,125 Table 1).

This is the case in particular when the following formula is satisfied:

$\begin{array}{cc}\mathrm{Fp}\le 39.9\mathrm{with}& \left(2a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+1.26*\mathrm{Nb}+0.477*\mathrm{Cu}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3d\right)\end{array}$

wherein Cr, Al, Fe, Si, Ti, Nb, Cu, Mo, W and C are the concentrations of the elements in question in mass-%. Table 2 containing the alloys shows that Fp is greater than 39.9 for Alloy 8, Alloy 3 and Alloy 2 and is exactly 39.9 for Alloy 10. For all other alloys, T_{s BCC} is lower than 939° C. and so Fp is <39.9.
Example for the Manufacture of Components with Welded Seams and their Properties

Tables 3a and 3b show the analyses of industrially smelted batches of Alloy NiCrAl—H alloys, used in this invention, from which sheets and welding rods (weld filler in the form of wire) were made. Formula (2a) Al+Cr≥28 is satisfied for these batches, as is therefore the requirement that was imposed on the “metal dusting” resistance. For the compositions in Tables 3a and 3b, the value for Fp was also calculated according to formula (3a). As required, Fp is lower than 39.9.

These batches were openly smelted in quantities of 16 metric tons, followed by a treatment in a VOD unit. Then electrodes were cast and subjected to ESR. Thereafter the alloy was annealed at temperatures between 900° C. and 1270° C. for 0.1 to 70 hours and hot-formed with intermediate annealings between 900° C. and 1270° C. for 0.1 hours to 70 hours. Hot-rolled sheets with the thickness of 25 and 16 mm were manufactured from Batch 319144. Hot-rolled wire was made from Batch 318385. After the hot rolling, the sheets were solution-annealed at 1100° C. for 40 minutes and then cooled in air. Then they were abrasive blasted, pickled and ground for removal of the oxide layer. The 25-mm-thick sheets had a grain size of approximately 89 μm, the 16-mm-thick sheets a grain size of approximately 82 μm. The 25-mm sheet and the 16-mm sheet therefore have a comparable grain size. The rolled wire was likewise abrasive blasted, pickled and ground and then cold-drawn to final thickness with intermediate annealings between 80° and 1250° C. for 0.05 minutes to 70 hours. Then the wire is solution-annealed under hydrogen in the temperature range of 800 to 1250° C. for 0.05 minutes to 70 hours and processed to welding rods of 2.0 and 2.4 mm diameter.

Sheet portions were cut out of the 25-mm-thick solution-annealed semifinished product sheet and annealed at 980° C. for 3 hours with subsequent air cooling. For creep tests, specimens transverse to the rolling direction were made from the sheets or sheet portions that were only solution-annealed and from those that were additionally annealed. Table 5 shows the results from the creep tests according to DIN EN ISO 204.

Sheet portions measuring 150×500 mm were cut out of the 16-mm-thick semifinished product sheet. Respectively 2 pieces were welded with a 70° V-groove seam by means of manual TIG welding under pure argon using the 2.0-mm-thick and 2.4-mm-thick welding rods from Batch 318385 as weld filler together with the welding parameters indicated in Table 4. The welded seam and the heat-affected zone were brushed immediately after the welding. Some partial pieces of the welded sheets manufactured in this way were annealed at 980° C. for 3 hours with subsequent air cooling, others at 1100° C. for 40 minutes with subsequent air cooling and yet others at 1100° C. for 3 hours with subsequent air cooling. Some were not subjected to any annealing. Seams or seam portions were also manufactured that were ground or not subsequently treated at all. For creep tests, specimens transverse to the welded seam were made from the welded sheets and from the welded and annealed sheets or sheet portions. Table 5 shows the results from the creep tests according to DIN EN ISO 204.

From Table 5, it is evident that, in creep tests at 600° C., an additional annealing of a solution-annealed sheet at 980° C. for 3 hours followed by an air cooling (specimen 19 45B and 19 46B) resulted in a shortening of the time to fracture t_u in comparison with the sheet that was only solution-annealed (specimen 19 23B and 19 7B). A creep test (specimen 302W and 303W) transverse to the welded seam without a further annealing (prior art T) likewise exhibits a significantly shortened time to fracture t_u in comparison with a creep test on the sheet that was only solution-annealed. An additional annealing, at 980° C. for 3 hours followed by an air cooling (specimen 247W and 249W), of the welded portion from which the creep specimens are made, has no marked impact on the time to fracture t_u, but it markedly reduces the creep elongation at break A_u^b. In contrast, an additional inventive annealing, at 1100° C. for 40 minutes followed by an air cooling (specimen 250W and 503W), of the welded portion from which the creep specimens are made, respectively causes a significant prolongation of the time to fracture t_u by a factor of approximately 3 as well as an increase of the creep elongation at break A_u^b, in some cases above the value of the specimens that were not annealed after the welding (specimen 302W and 303W). In particular, an additional inventive annealing, at 1100° C. for 3 hours followed by an air cooling (specimen 511W and 506W), of the welded portion from which the creep specimens are made, respectively causes a further prolongation of the time to fracture t_u as well as an increase of the creep elongation at break A_u^b, almost up to the value of the creep test on the sheet that was only solution-annealed but not welded (specimen 19 23B and 19 7B).

At 700° C., a creep test (specimen 306W) transverse to the welded seam without a further annealing (prior art T) likewise exhibits, just as at 600° C., a shortened time to fracture t_u in comparison with the creep test on the sheet that was only solution-annealed (specimen 30 34B). An additional annealing, at 980° C. for 3 hours followed by an air cooling (specimen 248W), of the welded portion from which the creep specimens are made, again causes a slight shortening of the time to fracture t_u in comparison with the specimen 306W that was not annealed after the welding. In contrast, an additional inventive annealing, at 1100° C. for 40 minutes followed by an air cooling (specimen 251W), of the welded portion from which the creep specimens are made, causes a prolongation of the time to fracture t_u by the factor of approximately 2 as well as an increase of the creep elongation at break A_u^b, significantly above the value of the specimen (306W) that was were not annealed after the welding. In particular, an additional inventive annealing, at 1100° C. for 3 hours followed by an air cooling (specimen 253W), of the welded portion from which the creep specimens are made, causes a further prolongation of the time to fracture t_u beyond the time to fracture of the creep test on the sheet that was only solution-annealed.

At 800° C., an additional annealing, at 980° C. for 3 hours followed by an air cooling (specimen 19 49B), of a solution-annealed sheet resulted in a similar time to fracture t_u in comparison with the sheet that was only solution-annealed (specimen 19 22B). A creep test (specimen 309W) transverse to the welded seam without a further annealing (prior art T) also has a similar time to fracture t_u in comparison with the creep test on the sheet that was only solution-annealed (specimen 19 22B). An additional annealing, at 1100° C. for 40 minutes followed by an air cooling (specimen 519W), of the welded portion from which the creep specimens are made, likewise causes a similar time to fracture t_u in comparison with the creep test on the sheet (specimen 19 22B) that was only solution-annealed.

This means that a heat treatment of the welded seam after the welding at 1100° C. for at least 40 minutes significantly improves, according to the invention, the time to fracture and the creep elongation at break of a creep specimen transverse to the welded seam at temperatures of 600 and 700° C. and thus in the range of γ′-formation. At higher temperatures above the γ′-solvus temperature, the non-annealed welded seam has a similar to better time to fracture t_u in comparison with the sheet that was only solution-annealed. An annealing of the welded seam influences only negligibly the time to fracture at service temperatures above the γ′-solvus temperature.

The claimed limits for the method and the nickel-chromium-aluminum alloy named Alloy NiCrAl—H used in the invention can therefore be justified in detail as follows:

Too low chromium contents mean that the chromium concentration underneath the oxide layer during use of the alloy in a corrosive atmosphere decreases very rapidly below the critical limit, so that a closed chromium oxide layer can no longer be formed. Therefore 18% chromium is the lower limit for chromium.

Too high chromium contents worsen the phase stability of the alloy, especially at the high aluminum contents of ≥1.8%. Therefore 33% chromium is to be regarded as the upper limit.

The formation of an aluminum oxide layer underneath the chromium oxide layer reduces the oxidation rate. Below 1.8% aluminum, the aluminum oxide layer is too incomplete to develop its effect fully. Too high aluminum contents impair the processability of the alloy. Therefore an aluminum content of 4.0% forms the upper limit.

The costs for the alloy increase with the reduction of the iron content. Below 0.01%, the costs rise disproportionally, since special precursor material must be used. For cost reasons, therefore, 0.01% iron is to be regarded as the lower limit. With increase of the iron content, the phase stability is reduced (formation of embrittling phases), especially at high chromium and aluminum contents. Therefore 7% Fe is a practical upper limit in order to ensure the phase stability of the alloy according to the invention.

Silicon is needed for the manufacture of the alloy. A minimum content of 0.001% is therefore necessary. Too high contents in turn impair the processability and the phase stability, especially at high aluminum and chromium contents. The silicon content is therefore limited to 0.50%.

A minimum content of 0.001% manganese is necessary for improvement of the processability. Manganese is limited to 2.0%, since this element reduces the oxidation resistance.

Titanium increases the high-temperature strength. Above 0.60%, the oxidation behavior may be impaired, which is why 0.60% is the maximum value.

Even very low magnesium contents and/or calcium contents improve the processing by the binding of sulfur, whereby the occurrence of low-melting NiS eutectics is avoided. At too high contents, intermetallic Ni—Mg phases or Ni—Ca phases may occur, which again greatly worsen the processability. The magnesium content and/or calcium content is therefore limited to at most 0.05%.

A minimum content of 0.005% carbon is necessary for a good creep resistance. Carbon is limited to at most 0.12%, since above such a content this element reduces the processability by the excessive formation of primary carbides.

A minimum content of 0.0005% nitrogen is necessary, whereby the processability of the material is improved. Nitrogen is limited to at most 0.05%, since this element reduces the processability due to the formation of coarse carbonitrides.

The oxygen content must be ≤0.020%, in order to ensure the manufacturability of the alloy. A too low oxygen content increases the costs. The oxygen content is therefore ≥0.0001%.

The content of phosphorus should be smaller than or equal to 0.030%, since this surface-active element impairs the oxidation resistance. A too low phosphorus content increases the costs. The phosphorus content is therefore ≥0.001%.

The contents of sulfur should be adjusted as low as possible, since this surface-active element impairs the oxidation resistance. Therefore at most 0.010% sulfur is specified.

Molybdenum is limited to at most 2.0%, since this element reduces the oxidation resistance.

Tungsten is limited to at most 2.0%, since this element likewise reduces the oxidation resistance.

Nickel is the element comprising the rest. A too low nickel content reduces the phase stability, especially at high chromium contents. Nickel must therefore be higher than or equal to 50%.

For highly corrosive conditions, but especially for a good “metal dusting” resistance, it is advantageous when the following relationship is satisfied between Cr and Al:

$\begin{array}{cc}\mathrm{Cr}+\mathrm{Al}\ge 28& \left(1a\right)\end{array}$

wherein Cr and Al are the concentrations of the elements in question in mass-%. Only then is the content of oxide-forming elements high enough to ensure a sufficient “metal dusting” resistance.

Beyond this, the following relationship must be satisfied in order that an adequate phase stability is ensured:

$\begin{array}{cc}\mathrm{Fp}\le 39.9\mathrm{with}& \left(2a\right)\end{array}$ $\begin{array}{cc}\mathrm{Fp}=\mathrm{Cr}+0.272*\mathrm{Fe}+2.36*\mathrm{Al}+2.22*\mathrm{Si}+2.48*\mathrm{Ti}+0.374*\mathrm{Mo}+0.538*W-11.8*C& \left(3a\right)\end{array}$

wherein Cr, Fe, Al, Si, Ti, Mo, W and C are the concentrations of the elements in question in mass-%. The limits for Fp and the possible incorporation of further elements have been justified in detail in the foregoing text.

If necessary, the oxidation resistance may be further improved with additions of oxygen-affine elements, such as, for example, yttrium, lanthanum, cerium, cerium mixed metal, zirconium, hafnium. They do this by being incorporated in the oxide layer, where they block the paths of diffusion of the oxygen to the grain boundaries.

Yttrium increases the oxidation resistance. For cost reasons, the upper limit is set to 0.20%.

Lanthanum increases the oxidation resistance. For cost reasons, the upper limit is set to 0.20%.

Cerium increases the oxidation resistance. For cost reasons, the upper limit is set to 0.20%.

Cerium mixed metal increases the oxidation resistance. For cost reasons, the upper limit is set to 0.20%.

If necessary, niobium may be added, since niobium also increases the high-temperature strength. Higher contents very greatly increase the costs. The upper limit is therefore set at 1.10%.

If necessary, the alloy may also contain tantalum, since tantalum also increases the high-temperature strength and the oxidation resistance. Higher contents very greatly increase the costs. The upper limit is therefore set at 0.60%. A minimum content of 0.001% is necessary in order to achieve an effect.

If necessary, the alloy may also contain zirconium. Zirconium increases the high-temperature strength and the oxidation resistance. For cost reasons, the upper limit is set to 0.20% zirconium.

If necessary, the alloy may also contain hafnium. Hafnium increases the high-temperature strength and the oxidation resistance. For cost reasons, the upper limit is set to 0.20% hafnium.

If necessary, boron may be added to the alloy, since boron improves the creep resistance. Therefore a content of at least 0.0001% should be present. At the same time, this surface-active element worsens the oxidation resistance. Therefore at most 0.008% boron is specified.

Cobalt up to 5.0% may be contained in this alloy. Higher contents markedly reduce the oxidation resistance.

Copper is limited to at most 0.5%, since this element reduces the oxidation resistance.

Vanadium is limited to at most 0.5%, since this element reduces the oxidation resistance.

Lead is limited to at most 0.002%, since this element reduces the oxidation resistance. The same is true for zinc and tin.

Too small grain sizes of smaller than 30 μm lead to a poor creep strength at higher temperatures. Too large grain sizes of larger than 600 μm lead to a very low creep ductility at temperatures in the range of γ′-formation.

A homogenization of the welded seams and/or for reduction of stresses by an annealing between higher than 980 and 1250° C. for times from 0.05 minutes to 24 hours, followed by a cooling in stationary shield gas or air, moving (blown) shield gas or air, improves the creep strength and the creep ductility of the welded seams in the range of γ′-formation. At a too low temperature lower than or equal to 980° C., the temperature is too low, so that homogenization cannot be carried out economically because of the long times required. At a temperature well above the solution annealing, marked grain growth takes place, which reduces the hot strength at low temperatures as well as the elongation in the creep test in the range of γ′-formation. Short times of less than 0.05 minutes are not adequate even at very high temperatures. Times longer than 24 hours are uneconomical, especially for larger components.

An annealing under shield gas reduces the oxidation of the material during the annealing and thus a material loss.

DESCRIPTION OF THE FIGURES

FIG. 1a: Left: Sketches in top view and cross section of the semifinished product forms sheet (, strip 2, bar 3 tube 4 and wire or weld filler in wire form 5. Right: Examples of manufacture of a component 7 by chamfering of two sheets 1a, 1b and joining by fusion welding with a weld filler in wire form 5 by means of a V-groove seam 6a.

FIG. 1b: Examples of manufacture of components 7. Left: Chamfering of two tubes at one end 4a, 4b and joining by fusion welding with a weld filler in wire form 5 by means of a V-groove seam 6b. Middle: Milling of a hole in a sheet 1c and inserting and joining of a tube by fusion welding with a weld filler in wire form 5 by means of a fillet weld 6c. Right: Joining of two strip portions 2a, 2b machined to make the edges fit abuttingly by fusion welding 6d with one part of the edges as weld filler 9.

FIG. 2: Loss of metal due to “metal dusting” as a function of the aluminum and chromium content in a strongly carburizing gas containing 37% CO, 9% H₂O, 7% CO₂, 46% H₂ as well as the following activities a_c=163 and p(O₂)=2.5×10⁻²⁷. (from (Hermse, C. G. M. and van Wortel, J. C.: Metal dusting: relationship between alloy composition and degradation rate. Corrosion Engineering, Science and Technology 44 (2009), pp. 182-185).

FIG. 3: Quantitative proportions of the phases in thermodynamic equilibrium as a function of the temperature of Alloy 690 (N06690) for the example of Batch 111389.

FIG. 4: Quantitative proportions of the phases in thermodynamic equilibrium as a function of the temperature of Alloy 693 (N06693) for the example of Alloy 3 from Table 2.

FIG. 5: Quantitative proportions of the phases in thermodynamic equilibrium as a function of the temperature of Alloy 693 (N06693) for the example of Alloy 10 from Table 2.

TABLE 1
Some alloys according to ASTM B 168-11. All values in mass-%.
Alloy	Ni	Cr	Co	Mo	Nb	Fe	Mn	Al	C	Cu	Si	S	Ti	P	Zr	Y	B	N	Ce
Alloy	72.0	14.0-				6.0-	1.0		0.15	0.5	0.5	0.015
600 -	min	17.0				10.0	max.		max.	max.	max.	max.
N06600
Alloy	58.0-	21.0-				Rest	1.0	1.0-	0.10	0.5	0.5	0.015
601 -	63.0	25.0					max.	1.7	max.	max.	max.	max.
N06601
Alloy	44.5	20.0-	10.0-	8.0-		3.0	1.0	0.8-	0.05-	1.0	0.5	0.015	0.6				0.006
617 -	min	24.0	15.0	10.0		max.	max.	1.5	0.15	max.	max.	max.	max.				max.
N06617
Alloy	58.0	27.0-				7.0-	0.5		0.05	0.5	1.0	0.015
690 -	min	31.0				11.0	max.		max.	max.	max.	max.
N06690
Alloy	Rest	27.0-			0.5-	2.5-	1.0	2.5-	0.15	0.5	0.5	0.01	1.0
693 -		31.0			2.5	6.0	max.	4.0	max.	max.	max.	max.	max.
N06693
Alloy	Rest	24.0-				8.0-	0.15	1.8-	0.15-	0.1	0.5	0.010	0.1-	0.020	0.01-	0.05-
602CA -		26.0				11.0	max.	2.4	0.25	max.	max.	max.	0.2	max.	0.10	0.12
N06025
Alloy	45	26.0-				21.0-	1.0		0.05-	0.3	2.5-	0.010		0.020					0.03-
45 -	min	29.0				25.0	max.		0.12	max.	3.0	max.		max.					0.09
N06045
Alloy	Rest	24.0-				8.0-	0.15	2.4-	0.20-	0.50	0.5	0.010	0.01-	0.020	0.01-	0.01-
603 -		26.0				11.0	max.	3.0	0.40	max.	max.	max.	0.25	max.	0.10	0.15
N06603
Alloy	Rest	28.0-		1.0-		2.0-	1.0		0.15	1.5-	1.0-	0.010	1.0
696 -		32.0		3.0		6.0	max.		max.	3.0	2.5	max.	max.
N06696

TABLE 2
Compositions of some alloys according to ASTM B 168-11. All values in mass-%.
Alloy	Batch	C	S	Cr	Ni	Mn	Si	Mo	Ti	Nb
Alloy	164310	0.07	0.002	15.75	73.77	0.28	0.32		0.2
600
N06600
Alloy	156656	0.053	0.0016	22.95	59.58	0.72	0.24		0.47
601
N06601
Alloy	111389	0.022	0.002	28.45	61.95	0.12	0.32		0.29
690
N06690
Alloy	Alloy 10 *)	0.015	≤0.01	29.42	60.55	0.014	0.075		0.02	1.04
693
N06693
Alloy	Alloy 8 *)	0.007	≤0.01	30.00	60.34	0.11	0.38		0.23	1.13
693
N06693
Alloy	Alloy 3 *)	0.009	≤0.01	30.02	57.79	0.01	0.14		0.02	2.04
693
N06693
Alloy	Alloy 2 *)	0.006	≤0.01	30.01	60.01	0.12	0.14		0.01	0.54
693
N06693
Alloy	163968	0.170	≤0.01	25.39	62.12	0.07	0.07		0.13
602
N06025
Alloy	52475	0.225	0.002	25.20	61.6	0.09	0.03		0.16	0.01
603
N06603
Alloy	UNS Mitte	0.080	≤0.01	30.00	61.20	0.1	1.5	2	0.1
696
N06696
									Ts BCC	Cr +
	Alloy	Cu	Fe	P	Al	Zr	Y	B	in ° C.	Al	Fp
	Alloy	0.01	9.42	0.009	0.16			0.001		15.9	19.1
	600
	N06600
	Alloy	0.04	14.4	0.008	1.34	0.015	0	0.001	669	24.3	31.2
	601
	N06601
	Alloy	0.01	8.45	0.005	0.31		0	0	720	28.8	32.7
	690
	N06690
	Alloy	0.03	5.57		3.2			0.002	939	32.6	39.9
	693
	N06693
	Alloy	0.03	4.63		3.08			0.002	979	33.1	41.3
	693
	N06693
	Alloy	0.03	5.57		4.3			0.002	1079	34.3	44.5
	693
	N06693
	Alloy	0.03	5.80		3.27			0.002	948	33.3	40.3
	693
	N06693
	Alloy	0.01	9.47	0.008	2.25	0.08	0.08	0.005	690	27.6	31.8
	602
	N06025
	Alloy	0.01	9.6	0.007	2.78	0.07	0.08	0.003	707	28.0	32.2
	603
	N06603
	Alloy	2	3						792	30.0	35.1
	696
	N06696
	*) Alloy composition from U.S. Pat. No. 4.88.125 Table 1.

TABLE 3a
Composition of the industrially smelted batches (G) of the nickel-chromium-aluminum
alloy named Alloy NiCrAl—H used in this invention, Part 1. All values in mass-%.
																Ts BCC	Cr +
	Name	Batch	C	N	Cr	Ni	Mn	Si	Mo	Ti	Nb	Cu	Fe	Al	W	in ° C.	Al	Fp
H	G	25 mm	319144	0.020	0.017	29.47	67.96	<0.01	0.05	<0.01	0.01	0.14	<0.01	0.11	2.11	<0.01	778	31.58	34.5
		sheet
H	G	16 mm	319144	0.020	0.018	29.47	67.89	0.01	0.05	<0.01	0.01	0.14	<0.01	0.12	2.16	<0.01	783	31.63	34.7
		sheet
H	G	Welding	318385	0.022	0.026	28.92	68.20	0.01	0.06	<0.01	0.02	0.14	0.01	0.47	2.03	<0.01	757	30.92	33.9
		rods
(H: Examples of the nickel-chromium-aluminum alloy named Alloy NiCrAl—H used in this invention, G: industrially smelted).

TABLE 3b

Composition of the industrially smelted batches (G) of the nickel-chromium-aluminum

alloy named Alloy NiCrAl—H used in this invention, Part 2. All values in mass-%.

Name

Batch

S

P

Mg

Ca

V

Zr

Co

Y

La

B

Hf

Ta

Ce

O

H

G

25 mm

319144

<0.002

0.002

0.005

<0.001

<0.01

0.03

<0.01

<0.01

<0.01

0.002

<0.01

<0.01

<0.01

0.001

sheet

H

G

16 mm

319144

<0.002

0.002

0.006

<0.001

<0.01

0.03

<0.01

<0.01

<0.01

0.002

<0.01

<0.01

<0.01

0.001

sheet

H

G

Welding

318385

<0.002

0.002

0.009

<0.001

<0.01

0.02

<0.01

<0.01

<0.01

0.003

<0.01

<0.01

<0.01

0.001

rods

(Pb: max. 0.002%, Zn: max. 0.002%, Sn: max. 0.002% are applicable for all alloys; meaning of H, G: see Table 3a).

TABLE 4
Welding parameters for the welding of the 16-mm-thick sheets (Batch 319144) with welding rods from
Batch 318385, of the nickel-chromium-aluminum alloy named Alloy NiCrAl—H used in this invention.
	Energy per
	Weld filler		Welding speed	unit length
	diam. in mm		Filler and	in cm/min	in kJ/cm
Thickness	Welding		Filler and	Root pass	top passes		Filler and	Filler and	Shield
in mm	technique	Root	top passes	I in A	U in V	I in A	U in V	Root	top passes	top passes	gas
16	m-TIG	2.0	2.0-2.4	120	15	190	17	5	8	20-25	Ar 4.6/
											pure Ar

TABLE 5
Results of the creep tests according to DIN EN ISO 204 on i) 25-mm-thick solution-annealed sheets (1100°
C./40 min/air cooling, grain size 89 μm) of Batch 319144 (BM) and on ii) 16-mm-thick solution-annealed sheets
(1100° C./40 min/air cooling, grain size 82 μm) of Batch 319144 (S), welded with welding rods from Batch 318385.
			Annealing after	Temperature	Initial stress	Time to fracture	Creep elongation
	Specimen	Material	welding	T in ° C.	σ0 in MPa	tu in h	at break Aub in %
	19 23B	BM	none	600	315	4832	7.3
	19 45B	BM	980° C./3 h/AC	600	315	1175	8.1
T	302W	S	none	600	305	626	3.5
	247W	S	980° C./3 h/AC	600	305	593	1.5
E	250W	S	1100° C./40 min/AC	600	305	1927	3.4
E	511W	S	1100° C./3 h/AC	600	305	2013	7.4
	19 7B	BM	none	600	270	15744	4.5
	19 46B	BM	980° C./3 h/AC	600	270	4978	5.1
T	303W	S	none	600	270	1222	2.7
	249W	S	980° C./3 h/AC	600	270	1224	1.6
E	503W	S	1100° C./40 min/AC	600	270	3816	4.1
E	506W	S	1100° C./3 h/AC	600	270	5256**)	**)
	30 34B	BM	none	700	90	3665	3.8
T	306W	S	none	700	90	2450	0.7
	248W	S	980° C./3 h/AC	700	90	1727	0.8
E	251W	S	1100° C./40 min/AC	700	90	3812	2.6
E	253W	S	1100° C./3 h/AC	700	90	4987	*)
	19 22B	BM	none	800	39	1878	16.9
	19 49B	BM	980° C./3 h/AC	800	39	1656	16.6
T	309W	S	none	800	39	2936	4.8
	519W	S	1100° C./40 min/AC	800	39	1896**)	**)
AC = air cooling
*) Fracture at the shoulder, not measurable,
**)Specimen still in testing.
E = According to the invention;
T = Prior art

Claims

1: A method for the manufacture of a component with one or more welded seams and/or for installation of a component in a plant with one or more welded seams, which comprise a nickel-chromium-aluminum alloy, containing (in mass-%) more than 18 to 33% chromium, 1.8 to 4.0% aluminum, 0.01 to 7.0% iron, 0.001 to 0.50% silicon, 0.001 to 2.0% manganese, 0.00 to 0.60% titanium, respectively 0.0 to 0.05% magnesium and/or calcium, 0.005 to 0.12% carbon, 0.0005 to 0.050% nitrogen, 0.0001-0.020% oxygen, 0.001 to 0.030% phosphorus, max. 0.010% sulfur, max. 2.0% molybdenum, max. 2.0% tungsten, the rest nickel, greater than or equal to 50% and the common process-related impurities, wherein the component is partly or completely made up of semifinished products of this nickel-chromium-aluminum wrought alloy and, after the welding, only the welded seams of this nickel-chromium-aluminum wrought alloy and the heat-affected zones surrounding the welded seams are subjected, for homogenization of the welded seams and/or for reduction of stresses, to an annealing between 98° and 1250° C. for times of 0.05 minutes to 24 hours, followed by a cooling in stationary shield gas or air, moving (blown) shield gas or air, with the consequence that the creep strength and the creep ductility of the welded seams are improved by this annealing, Cr + Al ≥ 28 ⁢ and ( 1 ⁢ a ) Fp ≤ 39.9 with ( 2 ⁢ a ) Fp = Cr + 0.272 * Fe + 2.36 * Al + 2.22 * Si + 2.48 * Ti + 0.374 * Mo + 0.538 * W - 11.8 * C ( 3 ⁢ a )

wherein the following relationships must be satisfied:

wherein Cr, Fe, Al, Si, Ti, Mo, W and C are the concentrations of the elements in question in mass-%.

2: The method according to claim 1, wherein the component contains welded seams and, after the welding, the entire component containing the welded seams is subjected, for homogenization of the welded seam and/or for reduction of stresses, to a further annealing between higher than 980 and 1250° C. for times from 0.05 minutes up to 24 hours, followed by a cooling in stationary shield gas or air, moving (blown) shield gas or air or in water, with the consequence that the creep strength and the creep ductility of the welded seams are improved by this.

3: The method according to claim 1, wherein, after a process of grinding of the welded seam and of the heat-affected zone, it is advantageous when roughness values Ra of 0.01 to 15 μm are attained, since this improves the corrosion resistance and especially the “metal dusting” resistance and raises them almost to the value of the parent material.

4: The method according to claim 1, wherein the semifinished product has a grain size of 30 to 600 μm.

5: The method according to claim 1, with a chromium content of 20 to 33%.

6: The method according to claim 1, with an aluminum content of 1.8 to 3.2%.

7: The method according to claim 1, with an iron content of 0.01 to 4.0%.

8: The method according to claim 1, if necessary with a content of niobium of 0.0 to 1.1%, wherein the formula (4a) is supplemented by a term for Nb: Fp = Cr + 0.272 * Fe + 2.36 * Al + 2.22 * Si + 2.48 * Ti + 1.26 * Nb + 0.374 * Mo + 0.538 * W - 11.8 * C ( 3 ⁢ b )

and Cr, Fe, Al, Si, Ti, Nb, Mo, W and C are the concentrations of the elements in question in mass-%.

9: The method according to claim 1, optionally with a content of zirconium of 0.0 to 0.20%.

10: The method according to claim 1, optionally with an yttrium content of 0.001 to 0.20%.

11: The method according to claim 1, optionally with a lanthanum content of 0.001 to 0.20%.

12: The method according to claim 1, optionally with a cerium content of 0.001 to 0.20%.

13: The method according to claim 1, optionally with a content of cerium mixed metal of 0.001 to 0.20%.

14: The method according to claim 1, optionally with a content of hafnium of 0.001 to 0.20%.

15: The method according to claim 1, optionally with a content of tantalum of 0.001 to 0.60%.

16: The method according to claim 1, optionally with a content of boron of 0.0001 to 0.008%.

17: The method according to claim 1, further optionally containing 0.0 to 5.0% cobalt.

18: The method according to claim 1, further optionally containing at most 0.5% copper, wherein the formula (4a) is supplemented by a term for Cu: Fp = Cr + 0.272 * Fe + 2.36 * Al + 2.22 * Si + 2.48 * Ti + 0.477 * Cu + 0.374 * Mo + 0.538 * W - 11.8 * C ( 3 ⁢ c )

and Cr, Fe, Al, Si, Ti, Cu, Mo, W and C are the concentrations of the elements in question in mass-%.

19: The method according to claim 1, further optionally containing at most 0.5% vanadium.

20: The method according to claim 1, wherein the impurities are adjusted to contents of max. 0.002% Pb, max. 0.002% Zn, max. 0.002% Sn.

21. (canceled)