The Formation of Nanodispersed Composite Metal Structure with Electroslag Surfacing

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Modern Applied Science; Vol. 9, No. 9; 2015

ISSN 1913-1844 E-ISSN 1913-1852
Published by Canadian Center of Science and Education

## The Formation of Nanodispersed Composite Metal Structure with Electroslag Surfacing

Gennady Nikolaevich Sokolov 1, Ilya Vasilevich Zorin 1, Alexander Alekseevich Artem'ev 1, Yuriy Nikolaevich

Dubtsov 1 & Vladimir Ilyich Lysak 1

1 Volgograd State Technical University, Volgograd, Russia

Correspondence: Gennady Nikolaevich Sokolov, Volgograd State Technical University, Volgograd, Russia.
E-mail: naplavka34@gmail.com

Received: January 22, 2015   Accepted: February 17, 2015   Online Published: August 30, 2015

doi:10.5539/mas.v9n9p333     URL: http://dx.doi.org/10.5539/mas.v9n9p333

_The reported study was partially supported both by RFBR (research project No. 13-08-01282 and 14-08-00868)_
_and Council on Grants of the President of the Russian Federation (research project No. MK-4265.2014.8)._

Abstract

Deposited metal requirements for cycled thermal-force operation were formulated. The influence of metal
droplets and slag emulsion circular flow on the formation of an electroslag deposited metal structure is revealed.
Micromorphology, ultimate composition and high-temperature properties of refractory deposited metal based on
iron, nickel and nickel aluminide Ni3Al were studied. Nano-sized particles of intermetallic compounds of various
origins are shown to increase the performance characteristics of refractory coatings in a cycled thermal-force
operation at temperatures of 700 to 1100 °С.

Keywords: electroslag surfacing, composite structure, aluminide Ni3Al, hot hardness, nano-sized particles

1. Introduction

In order to harden pressing dies and other tools for cycled thermal-force operation at temperatures up to 1100 °С
without shock loads it is efficient to use composite structure surfacing alloys (Reed, 2008; Sims, Stoloff, &
Hagel, 1995). These high-strength materials can have both an artificial and a natural composite structure. We
know (Dilthey, Woeste, & Aretov, 2006; Kivineva, Olson, & Matlock, 1995) that wear-resistant alloys of an
artificial composite structure are produced by adding carbide, nitride and boride particles and other refractory
chemical compounds into the welding pool. A natural composite structure is formed under the influence of high
gradient directed solidification (Petrushin, Chabina, & Nazarkin, 2012; Povarova, et al., 2008; Povarova, et al.,
2011). In this case the dendrites in the metal in the direction of solidification front the function as the alloy
matrix. The eutectic solid phases at the dendrites’ boundaries provide better deformation resistance to the metal.
This more efficient cycled thermal-force operation type of deposited composite metal can be produced by
electroslag surfacing methods (Hradecky, 1976; Kuskov, Skorohodov, & Ryabtsev, 2001).

The object of the article is to determine the principles and develop a method of electro-slag surfacing for the
formation of a creep- and wear-resistant alloy based on iron, nickel and nickel aluminide with nanodispersed
composite structur.

2. Materials and Methods

The research was performed with powder wires and electroslag surfacing technology using a current
solidification mould with a hollow electrode, electrically neutral surfacing materials being supplied into the slag
through a hole in the hollow electrode (Figure 1).

Both the powder wires and the surfacing technology were developed by the authors. Experimental samples were
applied to the surface of steel Х20 using АNF-6 flux; the surface of the slag pool was protected with argon.
Three types of deposited metal were studied based on: iron Х 200 CrMoNi 28-7-2, nickel Х 250 CrNiMoNb
22-66-4-2 and nickel aluminide Ni3Al (Х 50 CrNiAlMoZr 22-65-6-3-3).

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Figure 1. Electroslag surfacing diagram in a current supplying a solidification mould: 1 – hollow graphite
electrode; 2 – current supply section; 3 – insulator in a process cut; 4 – slag pool; 5 – electrically-insulating
layer; 6 – forming section; 7 – molten metal bath; 8 – product; 9 – thermocouple; 10 – graphite refractory; 11 –

composite wire

The structure, micromorphology and element composition of the deposited metal were studied with optical,
electron and atomic-powered microscopy. The composition and distribution of alloy elements in the structural
components were determined with energy-dispersive analysis in the secondary electron signal mode. Deposited
metal deformation resistance was estimated by its hardness at high temperature. Deposited metal
high-temperature scratch-hardness tests were performed in an inert atmosphere with a diamond cone, with a cone
vertex angle of 120° and a curve radius of 200 µm at a load of 0.5-1.0 N. Wear resistance criteria was the
indicator I = 1 / lg _Vd_, where _Vd_ is the deformed metal volume at a constant temperature at the track segment of
10 mm long remaining after the indenter movement.

3. Theoretical Justification of the Deposited Metal and the Process of Its Development

_3.1 Deposited Metal Requirements_

Computations and experiments (Schumon, 1968) show that thermo-force significantly increases the diffusion
rate of alloy elements in the thin surface layer of deposited metal. This leads to metal homogenization and
reduces its wear resistance. For example, Figure 2 shows that in the initial period of pressing in the die operation,
when stresses are at a maximum, the diffusion rate of chromium and molybdenum increases by several orders.

Figure 2. Influence of the contact time t of the die surface with surfacing by Х 200 CrMoNi 28-7-2 alloy and

steel forgings heated to 700 °С on the diffusion rate _D_ of chromium and molibdenum in the contact layer of

metal (100…150 µm)

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Hence, in order to increase the lifetime of deforming hot metal tools it is better to use surfacing alloys that resist
homogenization. Equation (1) (Schumon, 1968) shows that the metal homogenization time _t_ can be increased in
two ways:

#### ln( С − С 0
## ),

#### 2 ln( С − С 0 2 t

#### l 2 ln( С − С D ≈ π2 t

## ≈ l 2 ln( С 2− С 0 ),                   (1)

#### _t_

where _D_ is the diffusion coefficient of alloying elements; _l_ is the distance between hardening phases; _t_ is the total
time of the deposited metal and the hot alloy contact (the alloy-homogenization time); _С–С0_ is the alloying
elements concentration gradient.

The first method creates a certain degree of structural and chemical inhomogeneity in the metal characterized
with the coefficient _nc_

#### = l 2 ln  С π−2 С 0

####  С π−2 С 0 

####  

#### nс = l 2 ln  С −2 С 0                    (2)

The second method reduces atom diffusion mobility in the alloy matrix and in the hardening phases. Thus, the
formation of a surfacing alloy composite structure for cycled thermal-force operation shall be considered as
diffusional, structural and technological characteristics based on the following principles:

- in deposited metal, identification is required of the best atomic per cent ratio of carbide-forming elements
and carbon responsible for the formation of a strong framework of carbide eutectics and for dilution
resistance at high temperatures of mono- and complex carbides;

- within an operation temperature range in the deposited metal matrix, reduction is required in the atom
diffusion mobility of the basic and alloying element depending upon the matrix crystal lattice as well as the
number, morphology, distribution and type of sub-dispersed hardening phases;

- in order to provide deposited metal resistance to hot cracking and thermal fatigue cracking, it is necessary
to establish the best ratio of matrix alloy and disintegrated carbide eutectic in the metal that does not form
closed-link chains around primary dendrites;

- the surfacing process shall provide a minimal bending deflection of slag (gas) –metal interface to form a
flat solidification front in a welding pool to produce directional positioning of primary dendrite.

In this respect, within the range of maximum temperatures of 650–700 ºС, it is efficient to use comparatively
inexpensive carbonic chromium and molybdenum alloys of the Fe-Cr-Mo-C system as surfacing alloys, while at
temperatures up to 750–950 °С the nickel-based deposited metal of the Ni-Cr-Mo-C system is effective for
cycled thermal-force operation (Whelan, 1979; Wu, & Redman, 1994). For surfacing tools operating at
temperatures of 950–1100 ºС it is efficient to use refractory alloys with matrices based on nickel aluminide
Ni3Al (Lippold, Kiser, & DuPont, 2009; Liu, 1993; Povarova, K. B., et al., 2008; Stoloff, 1989); these alloys are
used in the aerospace industry as a construction material for gas turbine vanes. Powder wires and composite
wires (Kivineva, Olson, & Matlock, 1995) have been developed to produce such alloys in deposited metal.

_3.2 Electroslag Surfacing Process Study_

We know (Dilawary, 1978; Liu, Zang, Jiang, Geng, & Yao, 2012; Palti, & Shevtsov, 2002) that in the process of
electroslag welding and surfacing, due to technological factors, the welding pool melt front of solidification has
the deflected shape which influences the deposited metal quality. The flat solidification front providing dendrite
directional positioning can be produced by surfacing with a current supplying a solidification mould and a
two-circuit DC power supply (Figure 1) (Sokolov, Ryabchuk, & Lysak, 2008).

It is experimentally proved that this surfacing technology enables the energy environment in the slag pool to
change significantly. The slag areas with the highest potential are close to the hollow electrode in the bottom of
the current supply section which results in a high concentration of current lines (Figure 3). This contributes to
high-temperature generation in the spherical hollow space of the graphite electrode with its maximum up to 3500
°С. However the maximum volume electromagnetic force does not change abnormally the shape of the
slag–metal interface.

A mathematical model dealing with slag gravity flow caused by a significant (up to 1500 °С) temperature drop
between the overheated near-electrode area of the slag pool and the slag body adjacent to the slag–metal

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interface can explain the physics of the above phenomenon. The heat convection influence induced by the
temperature drop on hydrodynamic processes in the slag pool was previously neglected. This is due to the
temperature drop in a conventional electroslag process not usually exceeding 300–400 °С, and the Archimedes
forces density in the area of the main current is one order less than the electromagnetic force density (Palti, &
Shevtsov, 2002).

An electroslag process model demonstrated slag circulation flow with metal (emulsion) droplets in a
solidification mould (Sokolov, Ryabchuk, & Lysak, 2008). Liquid slag was believed to be the dispersion phase
while molten metal droplets were considered to be the dispersed phase. Two important parameters were
considered: temperature drop and the emulsion ‘thinning’ effect, the rheological state of the emulsion being
described by Oswald de Ville’s law (Bernadskaya, Valentinova, Mishta & Ryabchuk, 2005). In the vertical plane
(Figure 4), circulation is conditioned by two factors – the temperature difference between the areas near the
upper and lower electrodes (central circulation flow) and the temperature difference between the areas near the
side and bottom electrodes (near wall flow), in addition to the electric and magnetic field density. In the
horizontal plane, the rotational flow is conditioned by electric and magnetic fields only. Since metal pool surface
formation quality is mainly influenced by emulsion flow in the vertical plane, we will consider only these
movements.

(a) (b)

Figure 3. Distribution of the potential field (a) and current field (b) for the direct current two-circuit power
supply of a slag pool: 1 is a hollow electrode; 2 is a current supply section; 3 is an insulator; 4 is a forming
section; 5 is a product; ––––– are voltage equipotential lines; – – – are current lines. Electrolyte flows initiated

by power supply section electric field and hollow electrode (A), heat convection flows (B), hydraulic pressure

head direction (C)

In order to determine the velocity of the emulsion circulation flow caused by heat convection, let us consider the
following problem. Suppose a cylinder rod with diameter _d_ is located along the axis of a cylindrical vessel with
diameter _D_, filled with emulsion, with _d_ being much smaller than _D_, and the rod surface temperature _trod_ = const
higher than the cylinder wall temperature (Figure 4). The emulsion far from the cylindrical rod is motionless (no
gravitational flow), its temperature far from the cylindrical rod is constant and equals _t0._ Under the statement of
the problem, _trod_ is greater than _t0._ The emulsion volume is so large that free gravitational flow emerging near the
other bodies in the same volume does not influence the flow near the cylindrical rod, this flow being conditioned
by heat convection and electromagnetic force.

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Figure 4. Diagram for determining flow velocity of slag emulsion and metal droplets in a cylindrical coordinate

system (ref. to the text)

In order to simplify the solution of the problem, let us assume the following: inertia forces are negligible
compared to the gravity and viscous friction force; the pressure gradient is zero; liquid thermophysical
parameters (except density) are constant; density is a linear function of temperature.

We will consider the physics of the most intensive slag gravitation flow near the cylindrical rod in the cylindrical
coordinate system. With regard to the above assumptions, the Z-component of the gravitational flow equation
can be represented as follows:

####   

_n_

−

#### ϕ   = − g   r   

####   

####  k ∂  ∂ Vz  ∂ Vz ∂ r  ∂ r  ∂ r

####  ∂  ∂ Vz  ∂ V ∂ r  ∂ r  ∂

#### ∂

####     

#### z  z

####    

####  ∂ Vz  2 +   ∂  ∂ r   ∂

,          (3)

####    

#### 2 2 n 2−1 
## +  ∂ V ϕ    = − g (ρ− ρ0
####      ∂ r    

####   ∂ V ϕ    ∂ r 

#### 

#### _V_

## (ρ− ρ0 )

where _g_ is gravity acceleration, m/s 2, _ρ0_ is slag density far from the cylindrical rod, kg/m 3, _ρ_ is slag variable
density in the gravity flow area, kg/m 3, _Vz, Vφ_ are respectively vertical and tangential components of velocity,
m/s.

Solving this equation considering the temperature change in the emulsion moving layer and the emulsion density
linear dependency of temperature, we obtain the following gravity flow equation:

#####         

_n_

−1

##### 

##### _V_

##### ∂

#####   

#####       

##### ∂    ∂ Vz  n   ∂ r     * k    1 + 2   = −ρ0 g β t rod ∂ r  ∂ r    ∂ V   

##### ∂    ∂  ∂ r  ∂

##### n     2 z    ∂ r     * r    1 + 2   = −ρ0 g β t rod  1 − ,         (4)

#####  1 +   

#####  ∂ r     * r  1 + 2   = −ρ0 g β t rod  1 −   ∂ Vz     δ 

##### ∂

#####         

##### 

#####  ∂ Vz   ∂ r 

#####   ∂ V ϕ    ∂ r 

##### 

#####  ∂ Vz   ∂ r 

##### ∂

##### _r_

#####       

##### ρ0 g β t rod  1 −  δ

##### * r  2 0 g β t rod  1 − 

##### * r g β t rod  1 −

##### ∂ ∂

##### _Vz_ ∂ r

_z_

#####  

Solving this equation we obtain gravity flow velocity in the following form:

####  n       2 n + 3  

- _n_ _n_ +
####  n  ρ0 g β trod  n z =   * δ      2 n + 3  3 k  

#### n  ρ0 g β* t *   2 n + 3  3 k 

####   

#### n  ρ0 g β* t   n + 3  3 k 

####  n V z = 

####  ρ0 g β trod * 

####     3 k  

- _n_
#### =  n  ρ0 g β* trod  δ      2 n + 3  3 k  

+

1

_n_ _n_ +1

,            (5)

where _k_ _*_ is the reduced characteristic of consistence.

In order to determine the unknown value δ (Figure 3) we will use the integral equation of continuity and the heat
balance equation. Solving them and integrating the result we find:

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#####     

_n_

### (2 n + 3)(3 n +1)⋅ G ⋅ 3 n k n λ z
##### ⋅
# n (2 n +1) c р ρ nn +1 ( g β t rod * )

3 _n_ +1

_n_

##### n + 3)(3 n +1)⋅ G

### )(3 n +1)⋅ G ⋅ 3 n k n λ z
##### ⋅
### n (2 n +1) n +1

##### 2 n + 3)(3 n +1)⋅ G ⋅ 3
### n (2 n +1)

1 1
##### _n k n_

###  (2 n + 3)(3 n +1)⋅ G ⋅ 3 n k n λ z 
##### δ =  ⋅ n +1 1           (6)

###  (2 n + 3)(3 n +1)⋅ G ⋅
##### = 

+

##### ⋅ +1)

# c р ρ n ( g β t

1 1

# ( g t ) n

#####     

_р_

# ρ n ( g β t

_n_ +1 1
# n ( g β t rod * ) n

The above dependencies were used to calculate the velocity of metal droplets and slag emulsion circular flow
caused by heat convection. Input data for the calculation were thermophysical constants and indicators for slag
АNF-6 (70% CaF2 and 30% Al2O3), and metallic melt of nickel aluminide Ni3Al at temperatures typical of the
analysed electroslag surfacing process. The value of molten slag movement velocity calculated with no regard to
the temperature drop influence is 0.042 m/s (Sokolov, Ryabchuk, & Lysak, 2008). In our calculations, the
average velocity of slag circulation is 0.046 m/s. Since these are counter-current flows, the resultant velocity in
the slag pool is ten times less.

The analysis of the movement model for slag with metal droplets proved that a moderate resultant force is
formed in two toroidal slag flows rotating round the electrode axis; this force influences the emulsion movement
and enables the life of metal droplets to be increased, increasing their metallurgical processing efficiency. Due to
the pressure quadratic dependence of liquid flow velocity, the emulsion flow pressure on the molten metal
surface decreases almost one hundred-fold. This results in the formation of a flat slag–metal interface without
catering for the typical conventional electroslag welding and surfacing which influences the deposited metal
quality.

4. Results and Discussion

The survey of Х 200 CrMoNi 28-7-2 metal surfaced with the new electroslag process showed that the composite
structure of elecroslag surfaced metal with an iron based matrix consists of a molybdenum saturated solid
solution (a microhardness of 3800–3900 MPa). The metal contains sub-dispersed x-phase particles, with the
x-phase kinetics being similar to the σ-phase and composed of Fe12Cr36Mo10. There are also carbides (Mo, Cr,
Fe)7C3 (a micro-hardness of 8000–10000 MPa) and 20–30 volume per cent of carbide eutectic composed of
carbides (Мо, Сг, Fe)23C6 and МоС with a micro-hardness of 7300–7500 MPa (Figure 5).

a) overall view of the structure (×400) b) solid solution with intermetallic

compounds FeCrMo (×20K)

Figure 5. Deposited metal structure Х 200 CrMoNi 28-7-2

It was established (Figure 6, a) that, as a result of deposited metal hardening due to the formation of a strong
carbide framework, and due to the formation of nano-sized (80–100 µm) iron, chromium and molybdenum

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intermetallic compounds in a plastic hard solution, the deposited metal hardness at 700 °С after 5–6 hours of
testing does not noticeably reduce.

(a) (b)

Figure 6. Dependence of hot Brinell hardness of deposited metal Х 200 CrMoNi 28-7-2 of hold time at a
temperature of 700 ºС (а) and of test temperature (b): deposited metal; 2 – steel Х30 CrWVSi 2-8; 3 – steel Х40

CrWVMo 2-5

In comparison, the hardness of chromium and tungsten heat-resistant tool steel, used for hot deformation die
production, decreases fast after 1.5–2 hours of work. The ‘hot’ hardness studies at wide temperature ranges
(Figure 6, b) showed that the experimental alloy of the Fe-Cr-Mo-C system significantly exceeds the existing
tool steels in this parameter.

The research studied the structure of nickel-based deposited metal with uniformly distributed carbide eutectic
with a micro-hardness of 4000 MPa; nickel and carbides based alloyed solid solutions NbC, Mo2C (Figure 7).

(a) (b)

Figure 7. Microstructure of deposited metal (×600) Х 250 CrNiMoNb 22-66-4-2 (а) and alloying element

distribution (b). Columbium content 1.82 (% mass)

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Columbium sub-dispersed carbides, separating mainly on chromium and nickel γ-solid solution crystal lattice
defects, make the solution harder. They also influence relatively uniform distribution of silicon and carbide
eutectic (Figure 7, b). Molybdenum is also evenly distributed in the deposited metal, confirming its active
chemical bonding with nickel, and is one of the reasons for the increased heat resistance of such alloys compared
with nickel alloys containing tungsten instead of molybdenum.

Comparative ‘hot’ hardness test data for the deposited metal of various structural classes (Figure 8) indicate that
the considered material is highly resistant to cycled thermal-forces at temperatures up to 900 °С.

Figure 8. The influence of test temperature T on high-temperature Brinell hardness of deposited metal:

1 − Hastelloy C; 2 − Х 250 CrNiMoNb 22-66-4-2; 3 − Х30 CrWSiV 3-9

Metallographic research of nickel-aluminide based deposited metal showed that the metal has a complex
heterophased structure. Its structure (Figure 9) consists mainly of γ'-Ni3Al alloyed aluminide dendrites with a
micro-hardness of 2800–3200 MPa (Figure 9, а). The dendrites contain disordered γ-hard nickel-based solution,
alloyed with chromium, tungsten, molibdenum, tantalum, titanium and iron; dispersed secondary γ'secondary-phases;
intermetallic compounds CrNiMoTiZr (κ-phase) (Figure 9, b); and refractory carbides Ta2C, WC, Mo2C.
Interdendritic spaces contain irregularly shaped aluminides γ'eutectic of eutectic origin with κ-phase precipitations
(a microhardness of 4600–4800 MPa).

Carbide eutectic made up of Cr7C3 and Mo2C carbides as well as γ'secondary-phase (microhardness 5500-5700 MPa)
are uniformly distributed among primary crystallites γ'. The total content of γ'-Ni3Al phases of various origin in
the deposited metal is 85–90 vol. %.

The surface electronic probing of deposited metal proved that alloy composition is characterized by significant
chemical micro-inhomogenity which affects the nickel aluminium ratio, thus influencing the formation of
aluminide γ' (Figure 9, c–d). This correlation is very close to the stoichiometric in primary dendrite crystallite
areas located in nickel oversaturated axial volumes and in the spaces among them.

The scanning of a metallographic sample surface in a dendrite γ' cross-section area has revealed nano-sized
particles γ'secondary distributed uniformly in solid solution based on γ'+γ; these particles condition metal structure
dispersion (Figure 10).

It stands to reason that the increased heat resistance of deposited metal in question can be explained by a
concerted action of two types of structural hardening. The first type of hardening is determined by the stable
composite structure preserved for an extended period of time, composed of a strong framework of γ'secondary-phase
and carbide eutectic with a plastic matrix in the dendrite-shaped solid solution form alloyed with refractory
elements; the solution itself is made of eutectic based on γ'+γ. The second type of structural hardening is
characterized by the considerable volumetric portion in the alloy of thermodynamically stable nano-sized
particles of γ'secondary-phase and κ-phase micro-particles of stable size, morphology and distribution.

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(a) (b)

typical areas of surface electronic probing (c, d)

Figure 10. 3D image of surfaced metal surface in dendrite γ׳-Ni3Al cross-section

The test results of the comparative analysis for various structural classes of the surfaced metal (Figure 11 and
Figure 12) indicate that the produced electroslag surfaced alloy based on γ'-Ni3Al has an increased wear
resistance at temperatures of 950–1100 °С.

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Figure 11. Change of deposited metal hardness _НВ_ at high temperatures _Т_ : 1 - stellite Х160CrCoWNi 28-60-8; 2

- Hastelloy-C; 3 – Х 250 CrNiMoNb 22-66-4-2; 4 – deposited metal with matrix based on γ'-Ni3Al

Figure 12. Dependence of wear resistance indicator _I_ during metal scleroscope tests of test temperature _Т_ : 1 - Х
30 CrWSiV 3-9; 2 - Х 260 CrMoNiNb 28-6-2; 3 – Х 240 CrNiMoNb 25-65-4-2; 4 – deposited metal with matrix
based on γ'-Ni3Al

Full-scale industrial tests proved that the developed heat-resistant alloys deposited on various metallurgical tools
and dies operating at 700–1100 °С increased their wear resistance by 2–2.5 times compared to forged tools made
of chromium-tungsten-molibdenum refractory steels.

5. Conclusions

1. The new electroslag surfacing process using a conductive solidification mould and its two-circuit power supply
enables high-temperature metal of high quality with a natural composite structure to be achieved.

2. The natural composite is a combination of carbide eutectic’s strong framework and plastic matrix with evenly
spread hard-melting nano-sized phases. It enables wear resistance of tools and machine parts exposed to cycled
thermo-force to be increased.

3. For long-term processing at temperatures over 1000 ºC it is preferable to use deposited metal based on
intermetallic compound Ni3Al with a composite structure stabilized with nano-sized pro-eutectoid constituents.
This neutralizes the annealing processes developing in alloys at cycled thermo-force exposure.

Acknowledgments

The reported study was partially supported both by RFBR (research project No. 13-08-01282 and 14-08-00868)
and Council on Grants of the President of the Russian Federation (research project No. MK-4265.2014.8).

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