Cast iron welding electrodes

Metal filler compositions based on manganese and nickel for the welding of cast iron are disclosed. The filler composition preferably contains about 15 to 50% manganese and 15 to 35% nickel. The filler compositions may be incorporated into a welding rod for Shielded Metal Arc (SMA) or into the welding wire for Gas Metal Arc (GMA) welding or added as metal powders in Flux Cored Arc (FCA) or submerged Arc (SA) welding. The compositions are particularly useful for welding gray and ductile (nodular) cast irons.

This invention relates to metallic filler compositions used for welding 
cast irons. 
The cast irons are a family of iron alloys containing 1.8 to 4.5% carbon. 
The family includes gray iron (ca. 3.4% C), malleable iron (ca. 2.5% C), 
and ductile iron (ca. 3.4% C). The cast irons, especially ductile iron, 
possess many of the physical properties of steel, such as strength and 
toughness. These irons are increasingly being used to replace steel 
castings since they are cheaper and require less energy to produce. In 
order to realize their full potential, however, it is necessary to develop 
new and acceptable methods of welding cast iron parts. 
PRIOR ART STATEMENT 
Because of their high carbon content, two major problems arise in the 
fusion welding of cast irons: (a) the formation of massive carbides in 
regions of the parent metal that are melted or partially melted during the 
weld pass, and (b) the formation of martensite in regions of the parent 
metal that are heated to a temperature above the eutectiod but below the 
eutectic. Both carbide and martensite formation result in weld zones 
having properties different from those of the base metal. Thus, the weld 
zone may be lower in strength, lower in ductility, and most significantly, 
more brittle than the surrounding metal. 
Two approaches to the fusion welding of cast irons have been used to 
achieve sound welds. In the first approach, nodular graphite is produced 
in the weld which resembles the graphite contained in the base metal. This 
is accomplished by adding graphitizing agents, such as silicon, and 
nodularizing agents, such as magnesium or rare earth metals, to the weld 
metal from the welding rod or flux. In this way, a weld metal is produced 
which has a microstructure, mechanical properties, and thermal expansion 
properties similar to those of the base metal. 
In the second approach, nickel or copper is added as filler materials to 
the weld pool to produce an austenitic weld metal. The austenitic weld 
metal is tough, relatively soft, and exhibits other favorable properties. 
Satisfactory welds are produced by this approach because the eutectoid 
transformation to martensite is avoided and because of the ability of 
austenite to absorb carbon rejected by the melted cast iron, thus reducing 
the formation of carbides. 
Nickel works successfully in this second approach because it is an 
austenite phase stabilizer. When present in austenite, it shifts the 
eutectoid point so as to suppress the transformation of austenite into 
pearlite. Nickel is therefore classified as an austenite former. 
At present, nickel is typically introduced into the weld pool as an 
ingredient of the welding rods. Nickel-base covered electrodes are 
available for the arc welding of cast irons. These electrodes are 
classified as "pure" nickel, containing 90 to 95% Ni, nickel-iron, 
containing about 55% Ni, nickel-copper, containing about 60% Ni. The 
"pure" nickel and nickel-iron electrodes have emerged as the most 
satisfactory thus far for welding cast iron. 
The use of nickel as filler material for welding cast iron presents several 
problems. First, nickel is expensive. Second, the thermal properties of 
nickel are significantly different from those of cast iron and give rise 
to thermal expansion mismatch between base metal and weld metal. This can 
result in stresses high enough to cause cracking. Third, phosphorus has 
low solubility in nickel. This too can result in cracking when nickel-base 
electrodes are used to weld irons high in phosphorus. 
Accordingly, it is an object of the present invention to produce a filler 
material with substantially reduced nickel content for welding cast iron. 
It is also an object of the present invention to reduce the mismatch 
between the thermal properties of the fusion zone and the base metal. 
It is also an object of the present invention to produce a fusion zone of 
acceptable strength and toughness. 
SUMMARY OF THE INVENTION 
These and other objects are achieved by replacing nickel in the filler 
material with manganese. More specifically, it has been found that filler 
materials containing 15 to 50% manganese and 10 to 35% nickel provide 
acceptable cast iron weldments. A filler material containing 20% manganese 
and 20% nickel has been found to provide optimal results.

DETAILED DESCRIPTION OF THE INVENTION 
Manganese like nickel, acts as an austenite phase stabilizer and promotes 
formation of austenitic phase fusion zones. These zones absorb large 
amounts of carbon and other interstitial elements from melted cast iron. 
Carbide formation, and the brittleness associated with it, are thereby 
suppressed. 
Among the advantages of replacing nickel with manganese are the following: 
(a) Filler materials based on manganese-nickel alloys have a lower melting 
temperature than nickel-based alloys. Thus, less heat is required during 
welding. This results in savings and also decreases the cracking 
susceptibility in the partially melted region of the heat affected zone. 
(b) Substantial manganese additions reduce the hardness of the fusion zone 
and achieve better mechanical properties than found with high nickel 
filler materials. 
(c) Manganese-nickel filler materials achieve excellent thermal expansion 
compatibility with the base metal and better machinability than high 
nickel filler materials. 
(d) Manganese is cheaper than nickel. 
It has been demonstrated that 55% manganese metal powder additions to a 
neutral commercial submerged arc weld flux can produce lower fusion zone 
hardness in ductile iron weldments than found with iron-55% nickel filler 
metal wire. Metallic powder and flux compositions containing 10 to 35% 
nickel and 15 to 50% manganese provide even further improvements in cast 
iron weldment properties. For example, a mixture of 50% manganese and 10% 
nickel achieved the lowest weld hardness. It has also been shown that 
mixtures of manganese and nickel are better than either manganese or 
nickel alone for austenite stabilization. A filler composition containing 
20% manganese metal powder, 20% nickel metal powder and flux was found to 
be optimal for reducing the propensity to crack, increasing machinability, 
and achieving the greatest economy. 
The filler compositions of the present invention maybe used to weld cast 
iron to base metals made of either cast iron or other materials, such as 
steel, stainless steel, high nickel alloys. 
All methods and processes which may be used to introduce filler materials 
to the weld pool are considered part of the present invention. The filler 
materials of the present invention may be added as part of the flux. 
Alternatively, the filler materials may be incorporated into a welding rod 
or wire. The filler materials may be coated onto or be contained within 
such welding rods and wires. The filler materials may be introduced into 
the weld pool through a combination of welding rod and flux additions. 
The filler compositions of the present invention are useful for all methods 
of fusion welding cast iron. For example, the compositions may be used 
advantageously in the practice of submerged arc welding, gas metal arc 
welding shielded metal arc welding and flux cond metal arc welding. The 
compositions may also be used advantageously in the practice of gas 
welding, such as oxy-acetylene welding. 
EXPERIMENTAL RESULTS 
A series of tests were conducted to demonstrate the usefulness of the 
filler compositions within the scope of the present invention. These tests 
should not be construed as limiting the scope of the invention. Plates 
made of ductile iron containing 3.80% carbon were prepared for testing. 
Fifty 1000 gm flux samples were then prepared by mixing various amounts of 
99.9% pure Mn and Ni powder (100 mesh average) with Hobart H-700 SAW flux. 
(hereinafter the non-metallic flux). This non-metallic flux is desirable 
to enhance arc stability, to control the bead morphology and to protect 
the weld pool and bead from oxidation. Typically, a non-metallic flux 
contains major percentages of SiO.sub.2, MnO and CaO with smaller amounts 
of CaF.sub.2, TiO, Al.sub.2 O.sub.3 and MgO. The non-metallic fluxes may 
be enhanced for welding cast iron by adding thereto 5 to 10% by weight of 
graphite. While the graphite is not essential, it is desirable, 
particularly when welding bead on bead, without base metal dilution. The 
filler metal came from a Page Corporation type El-121/8 diameter low 
carbon steel welding wire. 
The compositions of the various samples prepared and treated are set forth 
in the following Table I: 
TABLE I 
______________________________________ 
Flux Commercial Flux 
Manganese Nickel 
No % grm % grm % grm 
______________________________________ 
50 100 1000 -- -- -- -- 
51 90 900 -- -- 10 100 
52 80 800 -- -- 20 200 
53 70 700 -- -- 30 300 
54 60 600 -- -- 40 400 
55 50 500 -- -- 50 500 
56 40 400 -- -- 60 600 
57 90 900 5 50 5 50 
58 80 800 5 50 15 150 
59 70 700 5 50 25 250 
60 60 600 5 50 35 350 
61 50 500 5 50 45 450 
62 40 400 5 50 55 550 
63 90 900 10 100 -- -- 
64 80 800 10 100 10 100 
65 70 700 10 100 20 200 
66 60 600 10 100 30 300 
67 50 500 10 100 40 400 
68 40 400 10 100 50 500 
69 80 800 15 150 5 50 
70 70 700 15 150 15 150 
71 60 600 15 150 25 250 
72 50 500 15 150 35 350 
73 40 400 15 150 45 450 
74 80 800 20 200 -- -- 
75 70 700 20 200 10 100 
76 60 600 20 200 20 200 
77 50 500 20 200 30 300 
78 40 400 20 200 40 400 
79 70 700 25 250 5 50 
80 60 600 25 250 15 150 
81 50 500 25 250 25 250 
82 40 400 25 250 35 350 
83 70 700 30 300 -- -- 
84 60 600 30 300 10 100 
85 50 500 30 300 20 200 
86 40 400 30 300 30 300 
87 60 600 35 350 5 50 
88 50 500 35 350 15 150 
89 40 400 35 350 25 250 
90 60 600 40 400 -- -- 
91 50 500 40 400 10 100 
92 40 400 40 400 20 200 
93 50 500 45 450 5 50 
94 40 400 45 450 15 150 
95 50 500 50 500 -- -- 
96 40 400 50 500 10 100 
97 40 400 55 550 5 50 
98 40 400 60 600 -- -- 
______________________________________ 
The tests were conducted with the submerged arc welding process which 
covers or submerges the weld zone with a layer of flux. A low carbon steel 
welding rod is then advanced through the flux to the weld zone to create 
the arc. The relative percentages of metal additives may be varied simply 
by mixing different percentages by weight of commercial flux, Ni powder 
and Mn powder. Thus a wide range of percentages of metal additives could 
be tested without having to cast rods with various metal additives. The 
term flux, as used hereinafter to describe the results of these tests, 
refers to the combined commercial flux and metal powders described in 
Table I above. 
The welding parameters used in this investigation were selected to produce 
a constant heat input of 88 kilojules/inch. The welding voltage was 30 
volts with a travel speed of 9.6 inches per minute, and the current was 
approximately 500 amperes. Using these parameters and the fluxes already 
described, single-pass bead-on-plate specimens were made on the pearlitic 
ductile iron base metal plate which has been ground prior to welding to 
remove any oxide scale. The weldments were made in the flat position using 
direct current reverse polarity. The specimens were allowed to cool to 
room temperature before the flux cover was removed. To investigate the 
temperature profiles produced by welding with some of these fluxes, ten 
thermocouples were placed at intervals across certain specimens. 
Following the tests, selected samples were then analyzed to determine the 
metallurgical and chemical composition of the resulting weldment. 
1. Effects of Manganese and Nickel on the Fusion Zone 
Fifty weld specimens were prepared and microscopically examined to observe 
the effects of nickel and manganese powder additions to the welding fluxes 
on the microstructure and properties of the fusion zone. 
FIG. 1 shows the volume percent of carbides in the fusion zone as a 
function of flux composition. With no manganese or nickel additions to the 
flux, the structure is carbides and very fine pearlite, as compared to the 
pearlite of the ductile iron base metal. The maximum amount of carbides in 
the fusion zone resulted with a 10% Mn powder addition to the flux. Small 
amounts of nickel in the flux are more effective than small amounts of 
manganese in reducing the amount of carbides, but for large additions this 
trend is reversed, as indicated by the close spacing of the lines on the 
lower part of the manganese side of this diagram. The smallest amount of 
carbides are found in the welds made with fluxes containing both manganese 
and nickel. 
The reduction in amount of carbides causes a decrease in carbide 
continuity, and eventually carbide-free zones are produced in the flux 
composition range which is illustrated as cross-hatched in FIG. 1. Fluxes 
which produced fusion zones with discontinuous carbides occupy a larger 
range, lying roughly within the 0-10% carbide region. 
Furthermore, micrographs of the samples suggest that the manganese 
additions are not only more effective than nickel in reducing amounts of 
carbides, but are also more effective in making the carbides discontinuous 
when comparable carbide content are considered carbide content and 
morphology is the strong effect on machinability. Machinability was 
simultated by drilling tests which were done on various fusion zone 
samples. These tests showed that, qualitatively, drilling was faster and 
easier as carbide contents and continuity decreased. 
FIG. 2, which relates the amount of retained austenite in the weld metal to 
flux composition, shows some similarities to FIG. 1. The amount of 
retained austenite increases with the alloy content of the flux. Also, 
fusion zones produced with high alloy fluxes contain either austenite or 
carbide, so the lower section of the diagram illustrates that a decrease 
of carbides is accompanied by an equal increase in austenite. Therefore, 
the carbide-free zone in FIG. 1 is matched by a 100% austenite area in 
FIG. 2. However, this is not true in the low alloy regions where the 
microconstituents contain significant amounts of other microconstituents, 
such as pearlite and martensite. 
Again, the relative effectiveness of manganese and nickel changes as total 
alloy content increases. Nickel is the more effective austenite stabilizer 
in low amounts, since at least 10% Ni in the fusion zone is needed to 
produce stable austenite; but FIG. 3 shows that manganese is more 
effective in large additions. The two figures also show that combinations 
of manganese and nickel are better than either manganese or nickel alone 
for austenite stabilization. 
Fluxes containing small amounts of manganese produced more martensite than 
those with no alloy additions. This is probably due to the increase in 
hardenability provided by small amounts of manganese. In some low-alloy 
welds, hard nodules were found. 
The graphitizing effect of nickel was evident from the micrographs which 
showed that graphite formed in the spaces between dendrites in a type D 
flake morphology. Residual amounts of nodularizing agents caused some 
spheres of graphite to form with the flakes. Type D graphite is damaging 
to mechanical integrity, and, all other things being equal, it would be 
better to produce no graphite by using Mn than to produce deleterious 
graphite with nickel. However, the graphite does help to reduce thermal 
expansion. 
Since hardness is strongly influenced by microstructure, FIG. 3 is very 
similar to FIG. 2 and even more so to FIG. 1. The isolated regions of high 
hardness are caused by the tendency toward increased hardenability and 
carbide formation when small amounts of manganese are used. Conversely, 
the region of very low hardness at the bottom of FIG. 3 is caused by the 
lack of carbides in the samples. Between these two extremes, the fusion 
zone hardness decreases as the amount of carbide is reduced by increasing 
alloy additions. The higher hardness values of samples in which nickel is 
the main alloy element may be caused in part by the greater continuity of 
the carbide networks in these samples, as discussed above. 
2. Effects of Manganese and Nickel on the Heat Affected Zone 
One problem which is common to both filler metal approaches is that as the 
base metal is heated above the eutectoid temperature, the dissolution of 
the graphite spheroids increases the carbon content of the surrounding 
matrix in the heat-affected zone. Upon cooling, this carbon-enriched 
matrix may transform to produce carbides and martensite. This concern over 
continuous carbide formation in the heat affected zone (HAZ) suggests 
using welding procedures with low heat input. 
The heat affected zone varied significantly with flux composition, as shown 
in FIG. 4. Also shown in this figure are the best estimates of the 
temperatures of the weld pools produced by these fluxes. As indicated, the 
HAZ widths show no completely consistent correlation with the pool 
temperatures. 
The hardness profiles in FIG. 5 show large variations in peak fusion line 
hardness as the flux compositions change. These variations may be 
influenced by the maximum fusion zone temperature, but the hardenability 
and austenite stabilizing effects of the alloy additions were probably 
more significant in causing the variations. Flux composition also had a 
strong influence on fusion zone hardness as discussed earlier. 
By comparing FIG. 5 with micrographs, it is seen that the changes in peak 
hardness reflect vastly different fusion line and HAZ microstructures for 
different filler compositions. Fusion lines resulting from 30% Mn -30% Ni 
and 20% Mn - 20% Ni fluxes are narrow and relatively soft. The lack of 
carbides in these fusion lines suggests that the fusion zone solidified 
below the base metal eutectic temperature; conversely, the large amount of 
carbides in the fusion zone which solidified at a temperature much higher 
than that of the base metal eutectic. The hardest HAZ, made with 20% Mn 
flux, had this same type of microstructure. Heat affected zones 
intermediate in carbide content and carbide continuity result when the 
fusion zone melting point is slightly above the base metal eutectic. 
Sequential micrographs demonstrate many of the fusion line and heat 
affected zone concepts discussed previously. Away from the fusion zone, 
the maximum local base metal temperature is not high enough to effect any 
phase transformations and the base metal is virtually unchanged. Moving 
toward the fusion zone, and thus toward higher maximum temperatures, the 
unaffected base metal gives way to HAZ. In this region, temperatures 
between the eutectoid and eutectic range are reached. Diffusion of carbon 
from the graphite nodules into the surrounding austenite and the 
subsequent austenite transformation upon cooling give rise to a variety of 
microstructures in the HAZ. In low-temperature regions, carbon diffuses 
away from the nodules during heating and back toward the nodules during 
cooling. This leaves a band of high-carbon austenite which transforms to 
make a ring of pearlite colonies within the ferrite ring. Higher 
temperature regions contain martensite because the more extensive 
diffusion of carbon produced a higher-carbon austenite, and because the 
higher cooling rates in these regions suppressed the pearlite reaction. 
Where temperatures are high enough to be in the liquid plus austenite 
ranges, incipient melting occurs. The nodules partially decompose in the 
eutectic reaction to produce a shell of high-carbon liquid. This liquid 
transforms to austenite plus carbide because of the high cooling rates. 
Bordering on the fusion zone, the carbide shells connect to each other or 
open into the fusion zone to form fusion lines of various morphologies. 
Finally, the fusion zone is reached, and consists of austenite and 
carbides in amounts determined by the fusion zone composition. 
3. Effects of Manganese and Nickel On The Coefficient of Thermal Expansion 
The average coefficient of thermal expansion was measured as a function of 
temperature for selected samples, and the results are plotted in FIGS. 6 
and 7. The curves for base metal and welds with no flux additions show 
changes in slope due to the pearlite to austenite transformation. As 
shown, the 60% Ni flux produces welds with lower coefficients of thermal 
expansions than that of the base metal, while welds made with 60% Mn flux 
have larger coefficients of thermal expansion (.alpha.). Large amounts of 
carbides in the weld, as in the 40% Mn sample, reduce the coefficient of 
thermal expansion. 
The effects of flux composition on thermal expansion are presented in a 
more general way in FIG. 8 which shows the value of .alpha. at 850.degree. 
C. Also shown, are symbols which indicate the types of cracking found in 
each zone of .alpha. values. As a rule, the welds with high value of 
showed longitudinal fusion zone cracks because of residual tensile 
stresses in the bead. On the other hand, welds with very low values of 
thermal expansion tend to develop compressive stresses on the bead, which 
can result in fusion line cracks at the very bottom of the fusion zone. 
By showing the general trends in .alpha., FIG. 8 suggests ways of 
optimizing the flux composition. To reduce cracking tendencies, the 
coefficient of expansion of the fusion zone should match that of the base 
metal (18.4.times.10.sup.-6 /.degree.C. at 850.degree. C.). By reducing 
the nickel content of fluxes containing only nickel, .alpha. increases 
towards the value for the base metal. At the same time, however, excessive 
amounts of carbide cause hardness and machinability problems. Similarly, 
reducing the manganese content of fluxes with only manganese will cause 
.alpha. to decrease towards the value for the base metal, but excessive 
carbides would again cause problems. Suitable values of .alpha. and 
acceptable hardness can be produced with certain fluxes along the 60% 
total alloy line. Satisfactory results may also be obtained using similar 
alloy additions when the flux composition is near the 20% Mn-20% Ni point. 
4. Chemical Analysis of the Fusion Zone 
Several samples were chemically analyzed to determine fusion zone 
composition. The following Table II lists the samples that were chemically 
analyzed. FIG. 10 correlates the chemical composition of the resulting 
weldment with its metallurgical properties as previously described with 
respect to FIGS. 1-4 and 8-9. The sample points illustrate that the 
resulting weldment has a different chemical composition than the initial 
flux composition used to create the weldment. These differences are due 
primarily to differences in Ni and Mn recovery rates and dilution from the 
base metal and filler rod. From this, the recovery of the alloy elements 
from the flux was calculated. It was found that the recovery of nickel was 
high than that of manganese. On the average, the recovery of manganese was 
64% and the recovery of nickel was 83%. This recovery is a combination 
figure that includes both additive consumption and base metal dilution. 
The chemical analysis and recovery for each sample are presented in Table 
II. 
TABLE II 
______________________________________ 
Sam- Flux Com- % Mn in % Ni in 
% C in 
% Mn % Ni 
ple position Fusion Fusion Fusion 
Recov- 
Recov- 
Point 
Mn--Ni Zone Zone Zone ered ered 
______________________________________ 
50 0%--0% 1.22 0.05 2.06 -- -- 
52 0%--20% -- 12.6 2.83 -- 63 
54 0%--40% -- 23.3 1.90 -- 58 
56 0%--60% -- 39.4 1.80 -- 66 
64 10%--10% 6.2 9.6 1.80 62 96 
76 20%--20% 13.1 15.9 1.12 66 80 
86 30%--30% 20.2 34.3 1.21 67 114 
74 20%--0% 13.4 -- 2.10 67 -- 
90 40%--0% 18.0 -- 2.15 45 -- 
98 60%--0% 50.9 -- 0.26 85 -- 
73 15%--45% 3.6 21.4 1.56 24 48 
94 45%--15% 43.1 20.7 0.12 96 138 
______________________________________ 
It will be observed that carbon is absorbed into the fusion zone from the 
base metal. In almost all cases, the fusion zone contained more than 1% 
carbon. 
Knowing the fusion zone composition allows determination of the composition 
of welding rods and wires which would give the same results as the alloyed 
fluxes. It should be mentioned that the low recovery of manganese and 
nickel in some cases, present no real practical problems, since this type 
of recovery is not encountered in other wire processes which are more 
suitable than the submerged arc welding process for most production 
welding of cast iron. 
5. Optimizing the Weld Metal Composition 
The determination of an optimum fusion zone composition was based on 
considerations of reducing cracking propensity, increasing machinability, 
and achieving economy. Fusion line cracking is aggravated by the presence 
of continuous carbides and the hard fusion lines which they create. Fusion 
zone cracking can be caused by excessive amounts of continuous carbide and 
by thermal expansion mismatch, which also can initiate fusion line cracks. 
Cracking problems, then, can best be avoided by producing welds with 
carbide-free fusion lines, soft fusion zones with discontinuous carbides, 
and coefficients of thermal expansion which match that of the base metal. 
Machinability also is increased when the fusion lines are soft and when 
the fusion zone has small amounts of discontinuous carbides in a soft 
austenite matrix. The best economy is achieved by reducing total alloy 
content and by substituting manganese for nickel. 
The discussion of fusion zone microstructure, heat affected zone width, and 
thermal expansion have shown that the least cracking and best 
machinability will be found in welds with at least 90% austenite, the 
narrowest heat affected zone (0.75-00.80 mm), and a thermal expansion 
coefficient of about 18.4.times.10.sup.-6 /.degree.C. at 850.degree. C. 
The ranges of flux compositions which produce each of these desirable 
properties are shown in FIG. 9. 
The area common to these separate ranges is the range of filler 
compositions which produce the soundest welds. This area relating to the 
soundest welds is also shown as the cross-hatched area of FIG. 10. Also 
shown in FIG. 10 is the point A representing a weldment formed with a flux 
having a 20% Mn-20% Ni composition. This composition would result in a 
weld metal which is about 90% austenite and only 10% carbides, has a heat 
affected zone of about 0.75 mm, and has a coefficient of thermal expansion 
of about 19.3.times.10.sup.-6 /.degree.C. at 850.degree. C. A second point 
B, and a third point C describe a triangle within the cross hatched area 
that defines the preferred metallurgical structure and chemical 
composition of the resulting weldments. When economy is taken into 
account, the 20% Mn-20% Ni composition is optimal, though higher alloy 
compositions could be useful in some specific applications, for example, 
when dilution is very high. 
6. Application to Other Welding Processes 
The results of the foregoing experiment indicate the invention is 
applicable to a wide range of welding processes and techniques. The 
selection of the preferred composition of the welding rod or welding wire 
for various other processes may be back calculated, taking into account 
the amount of expected additive consumption and base metal dilution that 
would result from the process and welding heat input selected. 
For example, the invention is believed to be particularly applicable to 
both the submerged arc (SA) and Flux Cored Arc (FCA) processes wherein the 
Ni and Mn additives would be added or mixed with the non-metallic flux. 
Alternately, the Ni and Mn additive would be alloyed with the Fe in the 
welding wire or rod for use in the Shielded Metal Arc (SMA) (known in 
Europe as manual metal arc welding) or Gas Metal Arc (GMA) processes. Each 
of these welding techniques is well known to those skilled in the art, and 
are described in standard reference materials used in the art. The 
following table lists the pages of each of two reference works wherein the 
foregoing processes are described. 
AWS Handbook, Volume 2, 7th Edition, (Published by the American Welding 
Society, Miami, Fla.). 
______________________________________ 
Pages 
______________________________________ 
Shielded Metal Arc Welding (SMA) 
43-76 
Gas Metal Arc Welding (GMA) 
113-152 
Flux Cored Arc Welding (FCA) 
153-187 
Submerged Arc Welding (SA) 
189-223 
______________________________________ 
ASM Handbook, Volume 6, 9th Edition, (Published by the American Society for 
Metals, Metal Park, Ohio 44073). 
______________________________________ 
Pages 
______________________________________ 
Shielded Metal Arc Welding (SMA) 
76-95 
Gas Metal Arc Welding (GMA) 
153-213 
Flux Cored Arc Welding (FCA) 
96-113 
Submerged Arc Welding (SA) 
114-152 
______________________________________ 
The foregoing pages of these references are incorporate herein by reference 
thereto. 
As noted above in Table II, the recovery of Mn in the resulting weldment 
averaged 64% while the recovery of Ni in the weldment was 83%. These 
recovery rates are a combination of metal additive consumption and base 
metal dilution. These recovery rates are lower than would be encountered 
in commercial practice for two reasons. First, the addition of Mn as a 
fine sieve metal powder induces high oxide losses as the Mn is evaporated 
and consumed in the arc. Secondly, the relative amount of heat used in the 
test was relatively high, at 88 kilojoules/inch. 
Typically, a much lower consumption rate would be found in the metal powder 
additive when used in the Flux Cored Arc welding process. When the Mn and 
Ni is alloyed in the rod, the recovery (from additive consumption) would 
be much higher, ranging from 75 to 85% for Mn and 90-95% for Ni. These 
recovery figures would be typical of the Shielded Metal Arc (SMA) and Gas 
Metal Arc (GMA) welding processes. 
In addition to the recovery rate, the base metal dilution must also be 
considered. This is the dilution of the resulting weldment by melted base 
metal entering the weld pool. Dilution is primarily a function of heat 
input, but varies some what with various welding techniques. The Submerged 
Arc welding process would have the highest dilution ranging from 25 to 40% 
with the Flux Cored Arc (FCA) welding process the lowest, averaging from 
10-25%. The two alloy wire or rod processes, SMA and GMA, would run from 
15 to 30%. 
While the preferred width of the heat affected zone described in FIGS. 4, 9 
and 10 is not one of the defining curves for the cross-hatched areas of 
FIGS. 9 or 10, it is desirable, as a matter of practice in welding cast 
iron, to keep this zone as small as possible and to use the minimum heat 
input necessary to achieve a true weld. 
The heat input is measured in Joules or kilojoules per inch or per mm, 
wherein 
EQU H.sub.1 =(VI/S) 
where H.sub.1 is heat input, V is volts, I is amperage and S is weld speed 
in inches-per-minute or mm-per-second. The tests of the present invention 
were conducted at 88 kilojoules per inch to insure adequate alloying of 
the weldment with the metal powder additives. It is preferable, however, 
to weld cast iron with a lower heat input, which may be as low as 20 to 25 
kilojoules per inch for a commercial welding process using an alloyed 
electrode. 
To determine the range of compositions desirable in each of the foregoing 
processes to achieve the preferred weldment described in FIG. 10, one must 
then consider, for each welding process, the recovery and dilution to be 
expected. 
For example, in the Shielded Metal Arc welding process, the metal additive 
would be alloyed in the welding rod. The preferred range of weld metal 
composition at points A, B and C illustrated in FIG. 10, one first adjusts 
for the expected recovery (addition consumption) rates of 75% for Mn and 
95% for Ni (typical of SMA welding process). The resultant intermediate 
number is then adjusted for base metal dilution to derive the desired 
composition of the consumable welding rod. These calculations for SMA 
welding process are 
______________________________________ 
Weld Metal Adjusted for 
Adjusted for 
Composition* Recovery Max Dilution 30% 
Points Ni Mn C Ni Mn Ni Mn 
______________________________________ 
A 16 13 1.1 17 17 25% 25% 
B 34 20 1.2 36 27 51% 30% 
C 17 19 1.7 18 25 26% 36% 
______________________________________ 
*balance substantially Fe. All compositions are in weight percent. 
The carbon referred to above typically was from the base metal, which may 
run as high as 3.5% for gray and ductile cast iron. Alternately 5-10% 
graphite may be added to the non-metallic flux surrounding the consumable 
to supply the C for the weld pool. This would be particularly desirable 
for weld deposits (refacing) wherein beads are laid upon beads without 
base metal dilution. The above table is an arithmetic projection based on 
standard recovery and dilution values. As a formulating range, the alloy 
steel core should consist of 17-50% Ni, 20-40% Mn, with the balance 
substantially Fe. 
The calculations for the consumable product to be used for Gas Metal Arc 
welding are essentially the same as for SMA since the recovery percentages 
and dilution percentages are essentially the same. The wire electrode, 
however, is not coated with a flux, since the gas supplied during the 
welding process inhibits the formation of oxides. It is therefore 
desirable to add C to the consumable electrode to achieve approximately 
1.5% C in the alloy. The preferred range for compositions of the alloyed 
wire electrode used for Gas Metal Arc welding would therefore be 20 to 50% 
Ni, 20 to 40% Mn, 1.5% C and the balance substantially Fe. This welding 
electrode would produce weld deposits within the cross-hatched area of 
FIG. 10. The preferred composition from an economic point of view would be 
the alloy at point A since it requires less alloying elements that other 
ranges in FIG. 10. 
The calculations of the outside range of compositions suitable for Flux 
Cored Arc welding electrodes are somewhat more variable because Mn powder 
consumption increases as heat input increases. Dilution, however, is lower 
(10-25%) with the FCA welding process then with other processes. The 
preferred range, however, may be calculated in a similar manner. With 
reference again to the weld metal compositions delivered in FIG. 10: 
______________________________________ 
Weld Metal Adjusted for 
Adjusted for 
Composition 
Recovery Max Dilution 
Points Ni Mn C Ni Mn Ni Mn 
______________________________________ 
A 16 13 1.1 17 17 23% 23% 
B 34 20 1.2 36 27 48% 36% 
C 17 19 1.7 18 25 24% 33% 
______________________________________ 
In the preferred method for using the flux cored rod, the metal Mn and Fe 
powders are mixed with non-metallic flux and 5-10% graphite powder. This 
flux core is surrounded by a hollow steel electrode. The above table is an 
arithmetic projection based on standard recovery and dilution values. As a 
formulating range, the metal powder should consist of 15-35% Ni and 15-30% 
Mn, with the balance of the composition a non-metallic welding flux. 
The section on chemical analysis showed that the 20% Mn-20% Ni flux 
produces a fusion zone with 13.1% Mn and 15.9% Ni. The wire composition 
needed to produce an equivalent weld will vary according to the amount of 
dilution by the base metal; but for 30% dilution, which is typical of 
submerged arc and gas metal arc welding processes, the wire would have to 
contain 18.7% Mn and 22.7% Ni. For simplicity this wire composition of 60% 
Fe, 20% Ni and 20% Mn is called 20-20 casting filler metal. This is 
substantial reduction of both Ni and total alloy content from the 55 Ni 
rod commonly used at present, which contains at least 50% Ni. 
While the invention has been described by referring to specific 
embodiments, it should be obvious to one of ordinary skill in the art that 
many variations can be made without departing from the scope of the 
present invention.