Alloying Effects of Reactive Elements in Ferritic Stainless Steels

Alloying Effects of Reactive Elements in Ferritic Stainless Steels P. R. Borneman Allegheny Ludlum Steel Corporation, Research Center, Brackenridge, PA 15014, USA

ABSTRACT Niobium and tantalum additions through what appears to be precipitation hardening by Laves phases can improve creep life in ferritic stainless steels. Amounts over and above of those required for stabilization are necessary to obtain this effect. The stabilization, however, can be totally or partially accomplished with a more reactive element such as titanium, leaving the entire niobium or tantalum content available for precipitation hardening. Higher than normal annealing temperatures are required to dissolve the Laves phase for precipitation hardening. Other reactive elements investigated produced no strong precipitation hardening effect. Titanium stabilization in ferritic stainless steels has been found to adversely affect wettability when copper brazing. Niobium stabilization has little effect on braze wetting and thus niobium-stabilized stainless steels with restricted titanium content can be easily brazed. A special heat treatment has been developed to cause AI2O3 whisker-like filaments to form perpendicular to the metal surface of Fecralloy for use as catalyst substrates. Use of mischmetal additions to replace yttrium also provides scale adherence and is amenable to the whisker growing treatment. KEYWORDS Ferritic stainless steels; Laves phase; titanium; niobium; tantalum; zirconium; brazing; catalyst substrates; rare earths; yttrium. INTRODUCTION Implementation within the past decade of melting technologies such as Argon Oxygen Jjecarburi zation and Vacuum Oxygen JJecarburi zation have made common a previously unavailable low level of interstitial elements in steel compositions. Among the chief beneficiaries of this technology have been the ferritic stainless steels where control of interstitial elements has provided an 307

308

P. R. Borneman ability to increase chromium content while retaining acceptable toughness at ambient temperatures. These high chromium ferritic stainless steels now represent some of the most corrosion resistant materials available and are particularly suited for service in oxidizing corrosion media. EXTENDED CREEP LIFE Ferritic stainless steels, because they have a lower coefficient of thermal expansion than austenitic stainless steels, are also candidate materials for high temperature low stress applications such as heat exchangers and automotive exhaust parts including catalytic converters. In many of these applications the elevated temperature strength required is little more than that necessary for the material to support its own weight or to withstand stresses imposed by thermal expansion acting upon the component design. The sensitization of ferritic stainless steels by precipitation of grain boundary chromium carbides causing a chromium depletion in the adjacent areas has long been recognized and the use of elements highly reactive with carbon to remove carbon from the alloy matrix as stable carbides is well known. Those elements with a great affinity for carbon, however, also tend to be strongly attracted to other interstitial elements such as nitrogen and therefore many specifications for stabilization elements require a minimum content of the stabilizing element which is some multiple of the combined carbon and nitrogen contents. Decreasing the total interstitial element content thus lowers the requirement for reactive stabilizing element content to be withdrawn from the alloy matrix by combination and provides an opportunity to use small amounts of these stabilizing, reactive elements as matrix alloy additions. Titanium and niobium, for a number of reasons including cost and carbon affinity, are the principal elements utilized for stabilization. A number of other reactive elements, however, which have a stabilization effect to a greater or lesser degree can also be considered for alloy additions when interstitial contents are lowered. Among these are tantalum, zirconium, hafnium, and vanadium. As a method of evaluating the effects of stabilizing elements in alloying the ferritic stainless steels, creep tests at 871°C were loaded to 8.274 MPa until elongation greater than 2% was obtained. Figure 1 shows the curve form obtained. In the usual case, the three stages of creep were almost indiscernible. First stage creep is essentially absent while the second and third stages appear to blend together in a smooth transition as the curve changes from horizontal to near vertical. For convenience in comparing a large number of alloys and treatments, the time in hours to 1% and 2% creep elongation provide convenient criteria. Experimental alloys were melted as 23 kg vacuum induction furnace heats tapped into three 7.7 kg ingots with an addition of the element under investigation made between ingot tappings. Ingots were hot rolled from 1121°C to 2.5 cm χ 10 cm sheet bar, conditioned and hot rolled from 1232°C to 0.25 cm χ 10 cm band and water quenched. This hot rolled band was conditioned and cold rolled to 0.102 cm χ 10 cm strip without annealing. To simulate continuous mill annealing of the cold rolled strip, tensile and creep sample blanks were sheared and then drilled to accommodate hanging by wire from an austenitic stainless steel bar. This bar was placed atop a U-shaped piece of austenitic stainless plate in the annealing furnace allow-

Alloying

Fig.

1.

Effects

Curve form generated in creep t e s t i n g of ferritic stainless s t e e l s at low s t r e s s

levels

of Reactive

F i g . 2.

309

Response surface of creep l i f e t o 2% e l o n g a t i o n a t 8 7 1 ° C - 8 . 2 7 4 MPa i n 16% c h r o m i u m s t a i n l e s s s t e e l s as a f u n c t i o n o f e f f e c t i v e n i o b i u m c o n t e n t and a n n e a l i n g t e m p e r a ture.

(6.89-8.29 W a )

and h i q h e r

Elements

temperatures

(732-87ΓΟ.

i n g t h e s e p a r a t e d s p e c i m e n s t o be s u s p e n d e d b e t w e e n t h e arms o f t h e U . One dummy s a m p l e i n e a c h l o t w a s f i t t e d w i t h a t h e r m o c o u p l e a t t a c h e d t o a h i g h speed r e c o r d e r , a l l o w i n g m o n i t o r i n g o f t h e t e m p e r a t u r e p r o f i l e o f t h e s p e c i mens. A c o n s t a n t f u r n a c e t i m e o f 3 m i n u t e s was used f o r each specimen l o t a n d g e n e r a l l y t h e s p e c i m e n s r e a c h e d f u r n a c e t e m p e r a t u r e f o r a p e r i o d o f 45 seconds t o one minute. A s a f r a m e o f r e f e r e n c e T a b l e 1 s h o w s c r e e p l i f e i n h o u r s t o 1% a n d 2% e l o n g a t i o n o b t a i n e d w i t h m a t e r i a l f r o m t w o c o m m e r c i a l h e a t s o f 11.5% c h r o m i u m stainless steels with titanium overstabilization ( i . e . , titanium content g r e a t e r t h a n t h a t t h e o r e t i c a l l y n e c e s s a r y t o consume a l l o f t h e carbon and n i t r o g e n as simple c a r b i d e s o r n i t r i d e s ) . Because these s t a i n l e s s s t e e l s are e s s e n t i a l l y a l l o y s o f chromium and i r o n , elements w h i c h have s i m i l a r atomic w e i g h t s , t h e r e w i l l be l i t t l e v a r i a t i o n i n a v e r a g e atomic w e i g h t o f the a l l o y s over t h e range o f chromium contents s t u d i e d . It therefore is TABLE

1

12% C h r o m i u m A l l o y s

With Titanium Hours

Effective T i Content Alloy Wt.(X) A Β

Effective Atom F r a c t i o n Ti

0.19 0.52

0.40 1.08

Overstabilization

t o C r e e p a t 8 7 Γ 0 - 8 . 2 7 4 MPa

Annealing Temperature 1010 1038 1Ö66 !?~~2% 1?~1* 3 25

5 41

2 18

5 30

(°C) 1093

3 5 5 20 34 10

12 20

Compositions Alloy

Ç

A 0.026 Β 0.015

N 0.016 0.012

M

n 0.37 0.39

P 0.027 0.015

S

Si

Çr

Ni

Al

Mo

Çu

ii

0.011 0.004

0.42 0.43

11.54 11.62

0.21 0.15

0.020 0.026

0.073 0.024

0.14 0.096

0.35 0.62

Nb

îa

0.01 NA 0.01 ΝΑ

P. R. Borneman

310

convenient to compare the effects of different stabilizing elements by expressing the degree of overstabilization as an atom fraction by dividing the weight percent of the element over and above that required for theoretical stabilization by one hundredth of the atomic weight. Table I shows a small increase in creep life with increased titanium overstabilization but no identifiable effect of annealing temperature. Table 2 shows creep life to 1% and 2% elongation obtained with 11% chromium alloys having varying degrees of tantalum overstabilization. Here a marked effect of tantalum on creep life can be seen with an indication of an optimum in the vicinity of 0.75 effective atom fraction. An effect of annealing temperature can also be seen but is less well defined. Several samples fractured at relatively long times before reaching 1% or 2% elongation. It should be noted that essentially all of the stabilization in these alloys was provided by the tantalum thus increasing the alloy cost. TABLE 2 11% Chromium Alloys With Tantalum Stabilization and Low Titanium Content Hours to Creep at 871°C-8.274 MPa Effective Effective Ta Atom Content Fraction Alloy (Wt %) Ta+Nb 0.,010 0.,255 1.,364 1.,366 2.,24 3.,35

c

D Ε F G H

0.0055 0.14 0.748 0.755 1.17 1.85

1%

1010

Annealing Temperature (°C) 1038 1066 2% 2% 2% 1% 1%

1%

1093

2%

5 8 5 8 5 8 5 8 260 t237 245 245 270 590 660 1495 tl542 t5050 t984 1170 1425 423 455 1400 1690 — 2020* 2150 2275 368 501 315 530 412 546 702 950 640(?) 920(?) 92 161 325 467 515 657

t = Prior Fracture Hours; * Extrapolated Value--Transducer Sticking. Compositions Al121

Ç

C D Ε F G H

0.016 0.0165 0.017 0.0205 0.0195 0.020

Ν 0.0175 0.0165 0.016 0.016 0.015 0.0135

Μη

Ρ

s

Si

Çr

Ni

Al

0.57 0.37 0.34 0.34 0.32 0.31

0.042 0.040 0.037 0.031 0.032 0.032

0.00 3 0.003 0.002 0.00 5 0.005 0.004

0.38 0.39 0.40 0.38 0.39 0.40

10 . 98 10.98 10.98 10.94 10.83 10.70

0.05 0.05 0.05 0.21 0.22 0.22

0.017 0.011 0.008 0.001 0.001 0.001

Mo 0.047 0.045 0.045 0.043 0.043 0.04 3

Cu 0.066 0.067 0.067 0.038 0.05 4 0.073

Ti 0.004 0.004 0.004 0.001 0.00 1 0.00 1

Nb <0.01 <0.01 <0.01 0.006 0.007 0.00 8

lâ 0.46 0.70 1.80 1.87 2.60 3.81

Table 3 presents creep life to 1% and 2% elongation values for a series of 16% chromium steels with titanium content high enough to provide theoretical stabilization and with tantalum additions ranging from about 0.1 to 1.4 effective atom fraction. Once again a similar strong maximum in the effect of tantalum content on creep life can be. seen and a superimposed effect of annealing temperature is evident. Unfortunately, tantalum is a costly alloying element, and although rather large increases in creep life are produced by tantalum additions the cost effective applications available to utilize such properties are severely limited.

Alloying Effects of Reactive Elements

311

TABLE 3 16% Chromium Alloys With Tantalum Additions and Titanium Stabilization Hours to Creep at 871°C - 8.274 MPa

Alloy

Effective Ta Content (Wt %)

Effective Atom Fraction Ta+Nb

I J Κ L M Ν

0.16 0.34 0.57 0.90 1.68 2.56

0.099 0.199 0.327 0.508 0.939 1.425

Annealing Temperature (°C) TÖIÖ 1Ö38 TÖ66 1093 1% 2% 1% 2% 1% 2% 1% 2% 22 93 258 468 485 258

39 129 333 1050 585 333

55 100 174 247 697 174

94 138 265 333 997 265

26 37 52 73 145 263 20 36 135 277 95 151 575 1110 1465 2450 790 1105 1757 2585 135 277 95 151

Compositions Alloy

Ç

Ν

Μη

P

I J K L M N

0.025 0.024 0.024 0.023 0.024 0.024

0.010 0.010 0.010 0.014 0.017 0.018

0.39 0.39 0.39 0.47 0.47 0.47

0.033 0.032 0.032 0.034 0.033 0.024

S S i 0.004 0.004 0.004 0.004 0.004 0.004

0.35 0.35 0.35 0.42 0.44 0.46

Çr

Ni

Al

16.12 16.16 16.08 15.64 15.09 15.10

0.25 0.25 0.25 0.29 0.29 0.23

Mo

0.025 0.025 0.025 0.020 0.018 0.010

Ç u T i

0.050 0.050 0.050 0.043 0.043 0.037

<.001 <.001 0.004 0.015 0.029 0.072

0.43 0.43 0.43 0.25 0.18 0.16

Nb

Ta

0.010 0.010 0.011 0.010 0.010 0.010

0.16 0.34 0.57 0.90 1.68 2.56

Table 4 presents creep life to 1% and 2% elongation values for a series of 16% chromium steels with sufficient titanium content to provide theoretical stabilization and with niobium additions ranging from about 0.5 to 1.5 effective atom fraction. A lower degree of improvement in creep life is evident TABLE 4 16% Chromium Alloys With Niobium Additions and Titanium Stabilization Hours to Creep at 871°C - 8.274 MPa

Hoy

Effective Nb Content (Wt %)

Effective Atom Fraction Nb

0 Ρ Q R S Τ

0.42 0.61 0.80 1.00 1.20 1.40

0.45 0.66 0.86 1.08 1.29 1.51

Annealing Temperature (°C) 1010 1038 1066 1093 1% 2% 1% 2% 1% 2% 2% 1% 33 148 107 113 51 23

43 195 123 140 69 33

42 155 222 230 69 21

60 208 274 279 77 29

60 130 130 240 56 28

80 172 153 298 75 40

13 65 158 272 56 36

25 89 231 325 78 42

Compositions Alloy

C

Ν

Μη

P

S

Si

Çr

Ni

Al

Mo

Çu

Ti

Nb

0 Ρ Q R

0.028 0.029 0.030 0.026 0.027 0.028

0.011 0.015 0.015 0.011 0.011 0.011

0.39 0.39 0.38 0.37 0.37 0.37

0.028 0.028 0.027 0.030 0.030 0.030

0.003 0.003 0.003 0.003 0.003 0.003

0.41 0.39 0.39 0.38 0.38 0.38

16.19 16.27 16.16 16.11 16.03 16.01

0.27 0.27 0.27 0.26 0.26 0.26

0.029 0.025 0.026 0.032 0.024 0.022

0.031 0.031 0.031 0.041 0.041 0.040

0.020 0.020 0.020 0.020 0.020 0.020

0.36 0.32 0.30 0.36 0.35 0.33

0.42 0.61 0.80 1.00 1.20 1.40

s T

312 P. R. Borneman in these data and the superimposed effect of annealing temperature can again be seen. Figure 2 shows a graphic representation of the response surface generated by the time to 2% creep elongation values when plotted against effective niobium content (wt %) and annealing temperature. Although lower creep life values are obtained, the combination of lower cost and lower atomic weight provides greater cost effectiveness for niobium as an alloying element than tantalum. Table 5 presents creep life data for 16% chromium steels with sufficient titanium content to provide theoretical stabilization and with zirconium additions ranging from about 0.5 to 1.2 effective atom fraction. Very little effect is noted from either zirconium content or annealing temperature. A similar study with vanadium additions again showed little effect. TABLE 5 16% Chromium Alloys With Zirconium Additions and Titanium Stabi1ization Hours to Creep at 871 °C - 8.274 MPa

Alloy U V W X

Effective Zr Content (Wt %)

Effective Atom Fraction Zr

0.49 0.78 0.97 1.10

Annealing Temperature (°C) TÖIÖ TÔ38 TÖ66 1093 1% 2% 1% 2% 1% 2% 1% 2%

0.537 0.855 1.06 1.21

6 5 11 13

10 12 35 30

7 5 7 10

12 12 16 23

8 5 10 10

15 12 25 20

8 15 10 18 15 34 47 105

Compositions Alloy

Ç

Ν

Μη

P

S

Si

Cr

Ni

Al

U V W X

0.025 0.027 0.027 0.027

0.014 0.012 0.060 0.011

0.35 0.40 0.33 0.40

0.029 0.030 0.026 0.030

0.0057 0.006 0.0058 0.006

0.38 0.40 0.45 0.40

16.09 16.05 15.84 15.93

0.23 0.27 0.23 0.28

0.038 0.26 0.034 0.27

Kb

Cu

Ti

Zr

0.042 <0.001 0.22 0.49 0.073 0.001 0.34 0.78 0.042 <0.001 0.29 1.01 0.095 0.002 0.33 1.10

Whittenberger, Oldrieve and Blankenship (1978) have shown that 16% chromium, 2% aluminum ferritic stainless steels with tantalum additions of 0.45% and 1.25% exhibited improved 800°C stress rupture and creep properties with increasing tantalum when compared to a similar composition without the tantalum addition. The same properties at 1000°C, however, were not improved by tantalum addition. All three materials were annealed at 997°C and all three of these alloys contained sufficient titanium to accomplish stabilization. Takeyama (1970) and co-workers found age hardening in Fe-Ta and Fe-Nb system alloys with small amounts of tantalum or niobium. The hardening resulted from precipitation of Fe2Ta or Fe2Nb (MgZn? type) Laves phases when aging at temperatures from 600°C to 700°C after quenching from 1300°C. Speich (1962) evaluated age hardening from precipitation of Laves phases from iron-niobium and iron-titanium solid solutions. Solubility of titanium in α-iron being much higher than that of niobium, the titanium content necessary for effective age hardening is so great that the alloys would be difficult to melt commercially. Dunning (1980) also shows Laves phase precipitation hardening in Fe-Ti, Fe-Mo and Fe-W alloys to require large amounts of the secondary element,

Alloying Effects of Reactive Elements 313 thus leading to practical melting and fabrication difficulties. Fe-Ta, Fe-Nb and Fe-Hf alloys, however, require only smaller secondary element additions to produce precipitation hardening. Pearson (1967) shows no simple iron-vanadium Laves phase to form and the rather questionable existence of an iron-zirconium Laves phase which might not produce precipitation hardening as indicated in the phase diagram presented by Shunk (1969). It can safely be concluded that the improved creep properties demonstrated result from Laves phase precipitation at the testing temperature. The annealing temperature effect noted appears to be that of the temperature necessary to produce near complete solution of the Laves phase. A practical limit on this solution temperature is imposed by very rapid grain growth which takes place at solution temperatures near 1100°C and which adversely affects cold formability of the steel products. BRAZING WETTABILITY Fabrication of heat exchangers and exhaust system components may often involve brazing. One of the more common and least expensive brazing filler metals is oxygen-free copper. Attempts to furnace braze components made from the ferritic stainless steels with copper have revealed problems which are attributed to poor wettability. One solution involving copper plating of the stainless steel has major cost disadvantages. Using a cold wall vacuum furnace and a 1.27 cm thick plate fixture to provide temperature uniformity, 3.8 cm dia. χ 0.51 mm thick discs of various ferritic stainless steels were tested with a small piece of 2.57 mm dia. oxygen-free copper wire placed on end at their center. After exposure to a brazing cycle with a maximum temperature of 1121°C, wettability was evaluated by measuring the diameter(d) and height(h) of the drop obtained and calculating a parameter 2 or tand /h. Values of d^/h were found to be largely unaffected by niobium 2 talum content when used as stabilizing elements. Values of d /h, however, decreased with increasing titanium content when this element was added for stabilization. An essentially non-wettable surface oxide film develops with titanium contents of 0.12% or above and lower titanium contents of about 0.01% are required for desirable wettability in these ferritic stainless steels. CATALYST SUBSTRATES Automotive exhaust system emissions in the U.S. are presently controlled through the use of catalytic converters in which the noble metal catalysts are washed onto the surface of ceramic substrates. Development of a metallic substrate for the catalysts presents important advantages in fabricability, vibration and impact tolerance, and weight savings. Using an alloy commonly known as Fecralloy (16%, Cr 5% Al, 0.3Y) which produces an adherent alumina scale, Chapman (1982) and others (1981) at General Motors Corp. have developed treatments to cause the alumina surface oxide to form as thin whiskers growing roughly perpendicular to the metal surface. Figure 3 shows an example of the whisker configuration obtained. Unfortunately, yttrium is extremely expensive, has a limited production and because of its reactivity presents grave recovery problems in melting. In fact, no method is presently available to recover yttrium values in remelting of scrap. Consequently, development of less expensive alloys with adher-

P. R. Borneman

314

Fig. 3. AI2Ο3 whiskers grown on the surface of Fe-Cr-Al-rare earth alloy. ent whisker growing capabilities is desirable. Alloys meeting this requirement were developed using the described whisker growing techniques in combination with a resistance heated oxidation test modified from ASTM specification B78-59T only in that a strip sample 0.051 mm χ 4.76 mm in section was used instead of a round wire and a test temperature of 1260°C was applied. Samples were heated to temperature for two minutes and cooled to room temperature for two minutes alternately until failure. It was found that alloys with from 12-23% chromium, 3-8% aluminum and small additions of cerium and/or lanthanum would grow the desired adherent Al2Ο3 whisker configuration and exhibited useful oxidation resistance. A useful melting addition for the rare earths is the relatively inexpensive mixture known as mischmetal and an optimum dissolved total rare earth metal content appears to be about 0.02% by weight. Only a few of the possible property effects of the reactive elements as alloy additions have been explored and further developments in this field may be expected from the interactions of melting and alloy technologies. REFERENCES Chapman, L. R., C. W. Vigor, J. F. Watton (1981). Oxide whisker-coated foil and method of preparation thereof. U. K. Patent Application GB2063723A. Chapman, L. R. (1982). Enhanced oxide whisker growth on cold-rolled aluminum-containing stainless steel foil. U. S. Patent 4,318,828. Dunning, J. S. (1980). Iron-based alloys strengthened by ternary Laves phases. U. S. BuMines Report of Investigations (1980) RI-8411. Pearson, W. B. (1967). Handbook of Lattice Spacings and Structures of Metals and Alloys - Vol. 2. Pergamon Press. 254-256, 945. Shunk, F. A. (1969). Constitution of Binary Alloys, Second Supplement. McGraw-Hill. 356. Speich, G. R. (1962). Precipitation of Laves phases from iron-niobium (columbium) and iron-titanium solid solutions. Trans. Met. Soc. AIME, 224, 850-858. Takeyama, M., M. Hesegawa, M. Okamoto, and K. Tokoro (1970). Precipitation hardening in the Fe-W, Fe-Nb, and Fe-Ta system alloys. J. Iron and Steel Inst, of Japan, 56, 173-182. Whittenberger, J. D., R. E. Oldrieve, and C. P. Blankenship (1978). Elevated-temperature mechanical properties and cyclic oxidation resistance of several wrought ferritic stainless steels. Metals Technology, 5, 365-371.