E.O.

Paton Electric Welding Institute of the National Academy of Science of Ukraine

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The Paton Welding Journal 01

2024
International Scientific-Technical and Production Journal ◊ Founded in January 2000 (12 Issues Per Year)
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editor-in-chief
I.V. Krivtsun E.O. Paton Electric Welding Institute, Kyiv, Ukraine
Deputy Editor-in-Chief
S.V. Akhonin E.O. Paton Electric Welding Institute, Kyiv, Ukraine
Deputy Editor-in-Chief
L.M. Lobanov E.O. Paton Electric Welding Institute, Kyiv, Ukraine
Editorial Board Members
O.M. Berdnikova E.O. Paton Electric Welding Institute, Kyiv, Ukraine
Chang Yunlong School of Materials Science and Engineering, Shenyang University of Technology, Shenyang, China
V.V. Dmitrik NTUU «Kharkiv Polytechnic Institute», Kharkiv, Ukraine
Dong Chunlin Guangzhou Jiao Tong University, Guangzhou, China
M. Gasik Aalto University Foundation, Finland
A. Gumenyuk Bundesanstalt für Materialforschung und –prüfung (BAM), Berlin, Germany
J. Kleiman Integrity Testing Laboratory, Markham, Canada
V.V. Knysh E.O. Paton Electric Welding Institute, Kyiv, Ukraine
V.M. Korzhyk E.O. Paton Electric Welding Institute, Kyiv, Ukraine
V.V. Kvasnytskyi NTUU «Igor Sikorsky Kyiv Polytechnic Institute», Kyiv, Ukraine
Yu.M. Lankin E.O. Paton Electric Welding Institute, Kyiv, Ukraine
O.V. Makhnenko E.O. Paton Electric Welding Institute, Kyiv, Ukraine
S.Yu. Maksymov E.O. Paton Electric Welding Institute, Kyiv, Ukraine
D.G. Mohan School of Engineering University of Sunderland England, United Kingdom
M.O. Pashchin E.O. Paton Electric Welding Institute, Kyiv, Ukraine
V.D. Poznyakov E.O. Paton Electric Welding Institute, Kyiv, Ukraine
U. Reisgen Welding and Joining Institute, Aachen, Germany
M. Rogante Rogante Engineering, Civitanova Marche, Italy
I.O. Ryabtsev E.O. Paton Electric Welding Institute, Kyiv, Ukraine
C. Senderowski Mechanics and Printing Institute, Warsaw University of Technology, Poland
V.M. Uchanin Karpenko Physico-Mechanical Institute, Lviv, Ukraine
Yang Yongqiang South China University of Technology, Guangzhou, China
Executive Director O.T. Zelnichenko, International Association «Welding», Kyiv, Ukraine
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© E.O. Paton Electric Welding Institute of NASU, 2024
© International Association «Welding» (Publisher), 2024
1
THE PATON WELDING JOURNAL, ISSUE 01, January 2024 ISSN 0957-798X

CONTENTS
ORIGINAL ARTICLES
S.Yu. Maksymov, G.V. Fadeeva, Jia Chuanbao, V.A. Kostin,
A.A. Radzievskaya, D.V. Vasilyev
INFLUENCE OF COOLING RATE ON MICROSTRUCTURE AND PHASE COMPOSITION
OF HAZ OF DUPLEX (DSS) 2205 STEEL IN WET UNDERWATER WELDING* ......................... 3

V.M. Korzhyk, V.Yu. Khaskin, E.V. Illyashenko, S.I. Peleshenko, A.A. Grynyuk,
O.A. Babych, A.O. Alyoshin, O.M. Voitenko
Hybrid laser-plasma welding: efficiency and new possibilities
(Review)* ..................................................................................................................................... 13

V.V. Skryabinsky, V.M. Nesterenkov, M.O. Rusynyk

Impact of heat treatment on mechanical properties
of joints during electron beam welding of 2219 alloy* ...................................... 22

O.Ye. Korotynskyi, M.P. Drachenko, A.M. Zhernosekov, I.V. Vertetska

Welding current formers using artificial long lines* ........................................ 27

S.V. Akhonin, V.O. Berezos, O.G. Erokhin, O.O. Kotenko, M.I. Medvedev, M.G. Lyashenko
Nickel scrap recycling by electron beam melting method** ............................. 32

O.S. Kremenchutskyi, S.S. Polishchuk

Effect of the texture of ferromagnetic Co‒Fe coatings
on their damping capacity** .............................................................................................. 36

V.O. Shapovalov, V.G. Mogylatenko, R.V. Lyutyi, R.V. Kozin

Nitrogen absorption by 04Cr18Ni10 steel in plasma-arc melting
under slag of CaO‒Al2O3 system** .................................................................................... 43

L.M. Lobanov, V.V. Savitsky, O.P. Shutkevych, K.V. Shyian, I.V. Kyianets
Nondestructive method of residual stress determination
in welded joints based on application of high-density current pulses
and speckle-interferometry*** ....................................................................................... 51

*Translated Article(s) from “Automatic Welding”, No. 12, 2023.

**Translated Article(s) from “Electrometallurgy Today”, No. 4, 2023.
***Translated Article(s) from “Electrometallurgy Today”, No. 4, 2023.

2
ISSN 0957-798X THE PATON WELDING JOURNAL, ISSUE 01, January 2024

DOI: https://doi.org/10.37434/tpwj2024.01.01

Influence of cooling rate on microstructure

and phase composition of HAZ
of duplex (DSS) 2205 steel
in wet underwater welding
S.Yu. Maksymov1, G.V. Fadeeva1, Jia Chuanbao2, V.A. Kostin1,
A.A. Radzievskaya1, D.V. Vasilyev1
1
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
2
Institute of Materials Joining, Shandong University
17923 Jingshi Road, Jinan 250061, China
ABSTRACT
The article shows the results of the analysis of changes in the microstructure and volumetric particles of phase components
of HAZ metal in modeling of welding thermal cycles, inherent in wet underwater welding and welding in air, with the use of
the Gleeble-3800 system. The value of cooling rate of different areas of weld metal in wet underwater welding and welding
in air was determined. It is shown that as a result of cooling impact of water environment, the cooling rate in wet underwater
welding is almost by an order higher than that of welding in air (w13/8 = 8.21 °C/s — air, in the middle of the weld, and in wet
underwater welding it is accordingly w13/8 = 81.70 °C/s in the middle of the weld, w13/8 = 165.85 °C/s at the beginning of the
weld and w13/8 = 320.51 °C/s in the weld crater). The change in volumetric particles of phase components of ferrite, austenite
and excess phases (chromium Cr2N nitrides) was determined in the microstructure of HAZ metal depending on the cooling
rate. Phase transformations almost completely occur in the high-temperature heat-affected zone (HHAZ) in the temperature
range T = 1300‒800 °С. Contribution of low-temperature heat-affected zone (LHAZ), temperature range T = 800‒500 °С on
the change in phase components is negligible. The amount of ferritic and austenitic components and especially the morphology
of austenite in the microstructure of HHAZ depend on the cooling rate, as well as the amount of precipitation of excess phases
(probably, chromium Cr2N nitrides)

KEYWORDS: duplex steels, wet underwater welding, input energy, cooling rate, phase composition. austenite, ferrite, micro-
structure in HAZ, HAZ simulation, thermal welding cycle, Gleeble

INTRODUCTION the chemical composition of the steel. There are sev-

The difference in the physical properties of the wa- eral methods to obtain data on the temperature-time
ter environment, first of all, which are 4 times higher dependences in the field of phase transformations.
than heat capacity and 25 times higher than thermal
conductivity compared to air cause a higher cooling
rate both in the high-temperature HAZ, which corre-
sponds to the temperature range of 1300‒800 °C, and
in the low-temperature HAZ, which corresponds to
the temperature range of 800‒500 °C, relative to other
types of welding. Information on the values, that ex-
actly correspond to the cooling rate in wet underwater
welding is almost absent in the literature.
ANALYSIS OF EXISTING PROCEDURES
FOR DETERMINATION OF INFLUENCE
OF THERMAL WELDING CYCLES (TWC)
ON HAZ MICROSTRUCTURE.
PROBLEM STATEMENT
To obtain data on the structure and dynamics of phase
transformations in duplex stainless steels during
welding, it is necessary to investigate the temperature
range, which according to the pseudobinary diagram
of Fe‒Cr‒Ni (Figure 1) [1] corresponds to the tem-
perature range of 1200 (1300)‒800 °C, depending on
Figure 1. Pseudobinary (Fe‒Cr‒Ni) diagram, built with the help
Copyright © The Author(s) of the equivalent Сreq/Nieq ratio [1]

3
S.Yu. Maksymov et al.

tire temperature range from room to liquidus tempera-

ture in a one specimen, thereby significantly reducing a
quantity of specimens required to obtain microstructures
corresponding to certain temperatures [5].
The procedure of using systems of different mod-
ifications of the Gleeble type is well-known, with the
help of which, a number of microstructures necessary
for research is simulated.
In [6], modeling of microstructures was carried
out with the use of the Gleeble system to assess the
effect of cooling time and alloying elements on the
Figure 2. Changes in the content of austenite depending on the HAZ microstructure of duplex stainless steel. In addi-
time of cooling at different concentrations of nitrogen [6] tion, submerged arc welding of steel was performed to
One of the methods is a procedure applied to the ob- compare the HAZ microstructures obtained by simu-
tained data on real TWC by recording the tempera- lation with Gleeble and in real welding. Modeling of
ture and time of the metal being in the temperature microstructures in the Gleeble-1500 thermomechan-
range of phase transformations. With the help of ther- ical simulator was performed according to the heat
mocouples, the temperature and the time spent in the treatment mode. The peak temperature T = 1350 °C,
mentioned temperature range are recorded, to which exposure τ = 1 s, and then the time of cooling from the
a certain area of the welded joint is corresponded. temperature of 800 to 500 °C (∆τ8/5) was 5; 20; 60 and
Based on these data, the average cooling rate in this 100 s, respectively.
range is calculated. The procedure of recording TWC It was determined that the optimal cooling time
parameters obtained with the help of thermocouples, (∆τ8/5) after welding is from 30 to 60 s for duplex stain-
placed directly in the welding zone or in the HAZ, is less steel with the chemical composition: 0.165 %
very time-consuming and associated with great diffi- N‒5.5 % Ni‒22.3 % Сr‒3.2 Mo. It is shown that cool-
culties, especially when recording TWC in the weld ing in the temperature range from 800 to 500 °C in the
metal. Most often, thermocouples are placed in the interval of 30‒60 s ensures the content of austenite in
HAZ of a welded joint. There are publications where the HAZ of not lower than 25 %.
recording of TWC was carried out directly in the weld
Figure 2 shows changes in the austenite content
metal. A new procedure was developed, which uses a
depending on the cooling time and nitrogen concen-
combination of thermocouples placed in a pool of a
tration.
molten weld metal and simultaneously placed on the
back side of a weld through drilled holes [2, 3]. This Analyzing the given data in Figure 2, it can be noted
procedure was applied and improved to obtain data on that the concentration of nitrogen has a greater effect on
the cooling rates of the weld metal in welding multi- the austenite content than the cooling time ∆τ8/5.
layer joints of superduplex 2507 steel [4]. Figure 3 shows the content of austenite, which is
A new heat treatment method has recently been determined in the microstructure during modeling
presented, in which a stationary TIG arc affects a disc- with Gleeble and in the HAZ microstructure in real
shaped specimen. This method allows covering the en- welding of steels with a different nitrogen content.
The difference in the amount of austenite in the
microstructures simulated by Gleeble and real HAZ
obtained during welding is relatively small at the sim-
ilar cooling time. Thus, the cooling time set by Glee-
ble can be successfully used as a reference point when
choosing welding conditions for duplex steels.
When studying the influence of input energy on
the evolution of austenite in the simulated HAZ of du-
plex 2205 steel, experiments on simulation the weld-
ing thermal cycle were carried out in the MMS-200
thermomechanical simulator according to simulated
welding thermal cycles corresponding to different in-
put energies [7]. The values of input energy are com-
parable to those used in submerged arc welding. The
Figure 3. Comparison of austenite content based on the Gleeble
simulation (1) and nitrogen content [6] in the HAZ microstructure morphology of austenite precipitation in the HAZ
in real welding [6] microstructure depending on the input energy and the

4
 INFLUENCE OF COOLING RATE ON MICROSTRUCTURE AND PHASE COMPOSITION OF HAZ

effect of the microstructure on the impact toughness Table 1. Cooling parameters in temperature ranges
were studied. Thermal cycles were simulated accord-
Description of Cooling time in temperature ranges, s
ing to the levels of input energy. Input energy was de-
groups
termined with the use of mathematical models by τ8/5, τ8/5 τ12/8
where τ8/5 is the cooling time for the specimen from 800 1 20; 50; 100; 300 7; 18; 37; 120
to 500 °C. Different values of τ8/5 6; 20; 50; 100; 300 and 2 7; 20; 50; 100 7; 7; 7; 7
600 s were taken to obtain different values of heat input, 3 20; 20; 20 7; 18; 37

which corresponded to the next values: 6.2; 11.3; 17.8; As τ12/8 and τ8/5 grew, the content of ferrite de-
25.2; 43.7 and 61.8 kJ/cm. The calculated levels of in- creased and the content of austenite increased, but τ12/8
put energy correspond to real levels in submerged arc was a more important cooling parameter that affects
welding from low to superhigh input energy. When the final microstructure of thermal modeling of HAZ
determining the share of austenite in the HAZ micro- of 2507 SDSS. At a ferrite content of about 50 %, τ8/5
structure, the following conclusion was reached after was 100 s, whereas τ12/8 was only 37 s. The impact
modeling. The content of austenite lower than 20 % toughness of HAZ grew with an increase in τ12/8, sim-
corresponds to the input energy of 0.62 kJ/mm and ilarly, the resistance to pitting corrosion in the HAZ
austenite/ferrite ratio of 1:1 is achieved when the input during welding increased with an increase in τ12/8 and
energy is increased to 6.18 kJ/mm. The value of impact τ8/5, but the effect of τ12/8 was particularly obvious. The
toughness changes and correlates accordingly with the most optimal properties were provided when τ12/8 was
morphology of austenite precipitation in the HAZ. from 18 to 37 s (τ8/5 = 20 s).
With the use of the Gleeble™-1500 system, a num- In [6], the cooling time in the temperature range
ber of microstructures were simulated, representing from 800 to 500 °C, τ8/5, was chosen as a variable cri-
those available in the HAZ of welded joints of duplex terion. In [7], τ8/5 was also chosen, then, according to
steels [8]. The simulation took place according to the mathematical models, it was converted to the corre-
thermal procedure: heating at a rate of 130 °C/s until sponding values of input energy. The rate of cooling
reaching the peak temperature T = 1300 °C, holding from a temperature of 1300 °С was chosen as a vari-
at the peak temperature τ = 1 s and τ = 10 s, then cool- able parameter in [8]. The authors of [9] also studied
ing at a rate from 90 to 2.0 °C/s. I.e., the cooling rate the cooling time in two temperature ranges, namely
from a temperature of 1300 °C was used as a variable. τ12/8 and τ8/5.
The data obtained as a result of the conducted stud- The results of extensive research on wet underwa-
ies allow assuming that the welding process with low ter welding of duplex steels and the properties of the
and medium heat inputs, which provide a cooling rate produced joints are not found in the literature.
of HAZ in the range from 20 to 50 °С/s, should be As a result of the analysis of studies [6-9] on the
the most effective for ensuring the necessary impact application of thermomechanical simulators of the
toughness of HAZ to –20 °C. This range of cooling Gleeble type to study the influence of various criteria
rate provides a good balance between the grain size on the microstructure and phase balance of HAZ of
and ferrite/austenite ratio. It was also determined that duplex steels, which in turn affect the main techno-
high cooling rates contribute to both the preservation logical properties of duplex steels, such as mechanical
of ferrite as well as higher deposition of nitrides. properties and corrosion resistance, it can be noted,
In [9], the study of the cooling time on the micro- that almost all studies relate to the parameters inher-
structure and properties of HAZ in 2507 steel was car- ent during welding in air using various technologies.
ried out using the Gleeble™-3800 thermomechanical In addition, there was no work that used a real thermal
simulator. The heating rate was 100 °C/s, and the max- cycle in wet underwater welding of duplex steels to
imum temperature was 1250 °C. The specimens were model HAZ.
held for 2 s before cooling. Since the range from 800 Therefore, the aim of this examination was to de-
to 500 °C was the most uncertain temperature range, termine and study the effect of cooling rate on the mi-
and the range from 1200 to 800 °C was a typical range crostructure and phase composition of HAZ of duplex
in which the transformation of ferrite into austenite steels, which is simulated by the Gleeble method, us-
took place, two τ8/5 ranges were chosen — cooling ing real welding thermal cycles corresponding to wet
time from 800 to 500 °C and τ12/8 — cooling time from underwater welding compared to air welding.
1200 to 800 °C to study the effect of cooling time on
RESEARCH METHODOLOGY
the microstructure and properties of 2507 steel. To an-
AND METHODS
alyze and compare the influence of different values of
τ8/5 and τ12/8 on HAZ modeling, three groups of cooling The Gleeble-3800 complex was used to study the
parameters were chosen, as given in Table 1. influence of cooling rate on the microstructure and

5
S.Yu. Maksymov et al.

Table 2. Chemical composition of 2205 steel (certificate data)

Number according Designation according Steel Content of elements, wt.%

to EN standard to EN standard grade C Mn P S Si Ni Cr Mo N

1.4462 Kh2CrNiMoN 22-5-3 2205 0.018 1.936 0.03 0.0008 0.303 4.931 22.146 2.557 0.1515

Table 3. Numerical values of TWC parameters

Temperature range, °С; cooling rate wcool, °C/s; cooling time τ, s

Specimen number 1300‒800 800‒500

w13/8 τ13/8 w8/5 τ8/5

1 — air (middle of the weld) 8.21 60.88 5.02 59.78
2 — water (middle of the weld) 81.70 6.12 50.34 5.96
3 — water (beginning of the weld) 165.85 2.94 100.00 3.00
4 — water (crater) 320.51 1.56 161.29 1.86

phase composition of HAZ of duplex 2205 steel under shown in Figure 4 and according to the data presented
the action of thermal cycle in wet underwater welding in Table 3.
compared to welding in air [10]. It was used to model To study the effect of cooling rates and cooling time
a number of microstructures that correspond to those on the microstructure and phase composition of HAZ,
formed in the HAZ under the influence of welding two ranges of cooling rates w13/8 and w8/5 were chosen,
thermal cycle. The chemical composition of the stud- which correspond to the high-temperature HHAZ and
ied steel is given in Table 2. low-temperature LHAZ. The same applies to the cool-
To model thermal cycles, which are inherent to ing time in the corresponding ranges, τ13/8 and τ8/5. The
those in wet underwater welding and welding in air, temperature range of 1200 (1300)‒800 °С, depending
the curves were used obtained experimentally apply- on the chemical composition of the metal with a typ-
ing thermocouples (Figure 4) [11]. ical range, in which the transformation of ferrite into
Table 3 shows the values of cooling rates and the austenite occurs to the greatest extent during cooling.
time spent by the specimens in different temperature The temperature range of 800‒500 °С was chosen to
ranges during TWC simulation, which correspond to analyze and compare the influence of different ranges
those presented in Figure 4. on HAZ microstructure during modeling.
Numerical values of parameters were obtained as a If we analyze the values of TWC parameters given
result of differentiation of TWC curves. in Table 3, namely, the time spent in the temperature
HAZ simulation was carried out according to the range τ13/8, for different areas of the weld in wet under-
thermal procedure: heating was carried out at a rate water welding, i.e.: the middle of the weld is 6.12 s;
of 100 °С/s to a temperature of Т = 1300 °С, hold- the beginning of the weld is 2.94 s; weld crater is
ing time at a peak temperature τ = 2 s, during cooling
1.56 s, so it is less than 10 s. I.e., no value corresponds
of the specimens, TWC were simulated, which are
to the concept, according to which it is recommend-
ed to cool a welded joint in the temperature range of
1200‒800 °C within 10 s from the time of achieving
the optimal microstructure and properties of the weld-
ed metal [12].
After thermal modeling, the specimens were cut
out in such a way as to cover all areas of the micro-
structures formed after TWC modeling. The further
studies were carried out using optical microscopy
(OM) and analytical scanning electron microscopy
(SEM). The content of austenite and ferrite in simu-
lated HAZ specimens was determined with the use of
Figure 4. Influence of environment on the nature of welding ther- the MIPAR image analysis software.
mal cycle in different weld areas [11]: a — beginning of the weld, To reveal the microstructure, electrolytic etching
specimen No. 3 (water); b — middle of the weld, specimen No. 2 was carried out in a 10 % solution of ammonium sul-
(water); c — weld crater, specimen No. 4 (water); d — middle of
the weld, specimen No. 1 (air) fate at a voltage of 15 V for 20 to 40 s.

6
 INFLUENCE OF COOLING RATE ON MICROSTRUCTURE AND PHASE COMPOSITION OF HAZ

Optical metallography was performed in Ver- the shape of austenite grains is observed, the size of
samet-2 (USA) and Neophot-32 (Germany) mi- austenite and ferrite grains decreases by an average
croscopes. Microhardness was measured in M400 of 1.5 times, especially in a high-temperature HAZ
Leco device. (T = 1300‒800 °С). The morphology of austenite pre-
RESULTS AND DISCUSSION. cipitation also changes. During cooling, grain-bound-
ary austenite (GBA) begins to be formed on the grain
ANALYSIS OF HAZ MICROSTRUCTURES
boundaries of δ-ferrite, and then acicular (Widman-
Figures 5 and 6 show the microstructures in different staetten) austenite (WА) nucleates along the grain
areas of HAZ simulated at different cooling rates. boundaries of δ-ferrite grains and grows in the middle
When analyzing the microstructures shown in of the grain. In addition to grain-boundary austenite
Figures 5 and 6, it should be noted that a change in and acicular (Widmanstaetten) austenite, if there is

Figure 5. Microstructures of high-temperature HAZ simulated at different cooling rates in the temperature range T = 1300‒800 °C: a,
b — specimen No. 1, air, w13/8 = 8.21 °C/s; c, d — specimen No. 2, water — middle of the weld, w13/8 = 81.70 °C/s, e, f — specimen
No. 3, water — the beginning of the weld, w13/8 = 165.85 °C/s; g, h — specimen No. 4, water — weld crater w13/8 = 320.51 °C/s

7
S.Yu. Maksymov et al.

Figure 6. Microstructures of low-temperature HAZ simulated at different cooling rates in the temperature range T = 800‒500 °С: a,
b — specimen No. 1, air, w8/5 = 5.02 °С/s; c, d — specimen No. 2, water — middle of the weld, w8/5 = 50.34 °C/s; e, f — specimen No. 3,
water — the beginning of the weld, w8/5 = 100.00 °C/s; g, h — specimen No. 4, water — weld crater, w8/5 = 161.29 °С/s
enough time for diffusion (depending on the cooling grains grow upon heating to a peak temperature of
rate), intragranular austenite (IGA) may nucleate and T = 1300‒1350 °C (depending on the chemical com-
grow in the middle of δ-ferrite grains. Since GBA and position of metal). This temperature corresponds
WA require less supercooling (which is controlled by to the single-phase region of ferrite (Figure 1) [1].
the cooling rate) for nucleation and growth compared During subsequent cooling with a decrease in tem-
to IGA, they have more time to grow and thus, they perature, ferrite loses its stability and transforms into
dominate in the final microstructure, especially at a austenite in the temperature range T = 1300‒500 °C,
low cooling rate. which corresponds to the two-phase region of austen-
In general, the microstructure in HAZ of du- ite and ferrite. The final phase composition of HAZ
plex stainless steel changes as follows: during heat- microstructure is the result after these two process-
ing, austenite transforms into ferrite, and ferrite es, namely, heating and then cooling. The final HAZ

8
 INFLUENCE OF COOLING RATE ON MICROSTRUCTURE AND PHASE COMPOSITION OF HAZ

Table 4. Size of ferrite grains in the HAZ metal structure depending on the cooling rate

Cooling rate w13/8, °C/s

Temperature range,
Description of zones 8.21 81.70 165.85 320.51
°С
Grain size h×1, μm

I — coarse grain zone 1300‒800 HHAZ 100‒350×150‒450 60‒130×100‒400 125‒250×150‒300 80‒230×120‒300

II — normalization zone 1100‒800 HHAZ 50‒150×100‒160 50‒100×100‒200 50‒155×100‒200 50‒150×100‒200

microstructure mainly depends on the cooling stage, At all cooling rates, the refinement of ferrite grains
which is characterized by the welding thermal cy- in the II temperature zone compared to the I zone
cle. At a cooling rate w13/8 = 8.21 °С/s, the presence is observed. At a cooling rate w13/8 = 320.51 °C/s,
of all types of austenite is observed, a large amount the size of ferrite grains in HHAZ is the lowest
of grain-boundary austenite, almost along all grain and is equal to 80‒230×120‒300 μm (I zone), and
boundaries, as well as intragranular austenite, and acic- 50‒150×100‒200 μm (II zone) (Table 4).
ular (columnar), i.e., Widmanstaetten austenite. The This data reaffirms that the time spent in the tem-
perature range of phase transformations is not the
precipitation of excess fine phases was not detected.
only factor that affects the completeness of phase
With an increase in the cooling rate from w13/8 = 81.70
transformations, as well as depends on the size of fer-
to 320.51 °С/s, the morphology of austenite precip- rite grains, i.e., on the diffusion of both ferrite- as well
itation changes. Precipitation of Widmanstaetten as austenite-forming elements: nitrogen, nickel, man-
austenite is no longer observed and the amount of ganese, and primarily on the nitrogen diffusion coef-
intragranular austenite is decreasing. At lower cool- ficient, since it is higher than other elements: nickel,
ing rates, in addition to it, grain-boundary austenite manganese, chromium and molybdenum.
is present, and at a cooling rate w13/8 = 320.51 °С/s,
PHASE COMPOSITION OF HAZ
almost only grain-boundary austenite is present in the
microstructure. With an increase in the cooling rate Figure 7 shows changes in volumetric particles of
from w13/8 = 81.70 to w13/8 = 320.51 °С/s, the precip- phase components, ferrite, austenite and excess
itation of an excess fine phase is observed, mainly in phase (probably, precipitation of Cr2N chromium ni-
the coarse ferrite grains, as well as sometimes at the trides) depending on the cooling rate in the simulat-
ed HAZ of duplex steel in the temperature range T =
boundaries of austenite and ferrite grains.
1300‒800 °C determined with the use of the MIPAR
When measuring microhardness at a cooling rate
w13/8 = 8.21 °C/s, the most important is austenite in software for image analysis.
HHAZ, both acicular (Widmanstaetten), as well as Figure 8 shows microstructures of the simulated
grain-boundary austenite, HV = 3300‒5150 MPa. A HAZ at different cooling rates corresponding to the
decrease in the microhardness of austenite occurs with temperature range T = 1300‒800 °C, which were used
an increase in the cooling rate from w13/8 = 8.21 to to determine the phase composition by means of the
w13/8 = 320.51 °C/s, as in HHAZ (T = 1300‒800 °C) as MIPAR software.
well as in LHAZ (T = 800‒500 °C) and is in the range Table 5 shows the values of volumetric parti-
of HV = 3360‒4390 MPa. Microhardness of austenite cles of phase components of the simulated HAZ
in BM is mainly HV = 2970‒3090 MPa, sometimes at different cooling rates in the temperature range
it reaches HV = 3300‒3570 MPa. Such a change in
the microhardness of austenite may indicate that the
main element that affects the microhardness of aus-
tenite is nitrogen. Microhardness of austenite depends
on its content content in the lattice. Microhardness of
ferrite at all cooling rates is from w13/8 = 8.21 °C/s to
w13/8 = 320.51 °C/s, both in HHAZ (T = 1300‒800 °C)
as well as in LHAZ (T = 800‒500 °C) is almost the
same and equals to HV = 2300‒2900 MPa. Micro-
hardness of ferrite in the base metal is almost the
same HV = 2300‒2570 MPa, sometimes it reaches
HV = 2970 MPa, i.e., microhardness of ferrite remains
unchanged both in HAZ as well as in the base metal.
Table 4 shows the size of ferrite grains in the sim-
Figure 7. Phase composition of HAZ of duplex steel depending on
ulated HAZ. the cooling rate in the temperature range of 1200 (1300)‒800 °С

9
S.Yu. Maksymov et al.

Figure 8. Microstructures (×100) of simulated HAZ at different cooling rates, processed with the use of the MIPAR software: a, b —
specimen No. 1, air, w13/8 = 8.21 °C/s; c, d — specimen No. 2, water — middle of the weld, w13/8 = 81.70 °С/s; e, f — specimen No. 3,
water — beginning of the weld, w13/8 = 165.85 °C/s; g, h — specimen No. 4, water — weld crater. w13/8 = 320.51 °C/s. Matrix — ferrite;
grain — austenite; small inclusions — nitrides
T = 1300‒800 °С, which were determined with the
Table 5. Phase composition of HAZ simulated at different cooling use of the MIPAR software.
rates in the temperature range T = 1300‒800 oС
If we compare the phase composition of the micro-
Fraction of phases, % structures at different cooling rates in the high-tem-
HAZ cooling perature range of HHAZ (T = 1300‒800 °С), then
rate, °C/s Excess phase
δ, ferrite γ, austenite
(fine)
with an increase in the cooling rate from w13/8 = 8.21 to
320.51 °С/s, the content of austenite decreases more
Base metal 52.000 48.000 –
8.21 57.499 38.674 3.236
than twice, and the content of ferrite, on the contrary,
81.70 64.644 30.268 3.746 increases by 1.2 times.
165.85 67.696 20.965 8.606 The data given in Table 5, indicate that the trans-
320.51 68.848 17.733 13.437 formation of ferrite into austenite occurs almost com-

10
 INFLUENCE OF COOLING RATE ON MICROSTRUCTURE AND PHASE COMPOSITION OF HAZ

pletely in HHAZ in I and II zones in the temperature Conclusions

range T = 1300‒800 °C, and the completeness of the 1. The influence of cooling rate on the microstruc-
transformation depends on the time spent in this range ture and phase composition in the simulated HAZ
and on the size of ferrite grains. At the same time, of duplex 2205 steel with the use of Gleeble-3800
the morphology of austenite precipitation depends was studied. Simulated microstructures with cool-
to a greater extent on the time spent in this tempera- ing rates from w13/8 = 8.21 to 320.51 °C/s, as well as
ture range. At a cooling rate w13/8 = 8.21 °C/s, the from w8/5 = 5.02 to 161.29 °С/s showed a change in
time spent in the temperature range T = 1300‒800 °C the phase composition of austenite and ferrite. To the
is 60.88 s, in the microstructure the presence of all greatest extent, the change in the volumetric particles
types of austenite is observed: grain-boundary, Wid- of phases occurs at the cooling rates w13/8 — from
manstaetten and intragranular. Precipitation of excess 81.70 to 320.51 °С/s, which correspond to the cooling
phases is not observed. When the cooling rate grows rates in wet underwater welding in the high-tempera-
to w13/8 = 320.51 °С/s, the morphology of austenite ture HAZ.
precipitation changes, precipitation of acicular (Wid- 2. When the cooling rate w13/8 grows from 8.21 to
manstaetten) austenite is not observed anymore and 320.51 °C/s, the volumetric fraction of austenite de-
the amount of intragranular austenite decreases. At all
creases by 2.18 times (from 38.67 to 17.73 %), the
cooling rates, grain-boundary austenite is present, and
volumetric fraction of ferrite, on the contrary, grows
at a cooling rate w13/8 = 320.51 °С/s, in the microstruc-
by 1.2 times (from 57.41 to 68.85 %).
ture, mostly only grain-boundary austenite is present.
3. The content of austenite at the cooling rates from
With an increase in the cooling rate from w13/8 = 81.70
w13/8 81.70 to — 320.51 °C/s, which correspond to the
to 320.51 °С/s, precipitation of tiny excess phases is
cooling rates in wet underwater welding, decreases
observed mainly in the coarse ferrite grains, as well as
from 30.27 to 17.73 %. The main part is formed by
sometimes at the boundaries of austenite and ferrite
grain-boundary and intragranular austenite. At the
grains.
cooling rates w13/8 = 8.21 and 5.02 °С/s, which are in-
The analysis of chemical elements in the high-tem-
herent during welding in air, all types of austenite are
perature HAZ (T = 1300‒800 °С) revealed precipi-
observed: grain-boundary, acicular (Widmanstaetten),
tation of excess phases with an increased content
of chromium in both austenite and ferrite. This is and intragranular austenite.
explained by the fact that at a high cooling rate, the 4. The phase transformation of ferrite into austen-
transformation of ferrite into austenite does not occur ite occurs mostly in the high-temperature HAZ, in the
to the full extent, the amount of austenite precipitation temperature range T= 1300–800 °С.
decreases, and an excess phase with a higher chro- 5. The completeness of phase transformation of
mium content is observed in ferrite grains (probably, ferrite into austenite depends on the cooling rate, i.e.,
Cr2N). At a cooling rate w13/8 = 320.51 °С/s in the I the time spent in a given temperature range, and also
zone (T = 1300‒800 °С), in ferrite grains, with the on the sizes of ferrite grains.
use of a scanning microscope (SEM), rod-type in- 6. The amount of excess phase precipitation (prob-
clusions of up to 10 μm in length with a chromium ably, chromium Сr2N nitrides) is directly proportional
content of 23.2‒24.14 % were revealed. Chromium to the cooling rate and also depends on the amount of
Cr2N nitrides found in [13] also have a rod-like ap- the austenitic component. With an increase in the cool-
pearance. This indicates that these are probably Cr2N ing rate, the amount of chromium nitride precipitation
chromium nitrides, since the base metal contains almost four times increases, i.e. from 3.24 to 13.44 %.
0.1515 % of nitrogen and the carbon content is lower 7. With an increase in the cooling rate, a decrease
than 0.02 % (0.018 %). Since the process of transfor- in the size of ferrite grains is observed.
mation of ferrite into austenite is a diffusion process, 8. The amount of ferritic component grows with
the completeness of the transformation of ferrite into an increase in the cooling rate, but it is not critical
austenite, that is, the final phase composition of HAZ and amounts to 68.85 %, i.e., it does not reach 70 %,
microstructure, depends on the diffusion coefficients which is allowed by recommendations and standards.
of ferrite-forming elements and austenite-forming el- Despite the fact that the cooling rate in wet underwater
ements, primarily nitrogen (Table 5, Figure 7). welding is by an order higher than during welding in
The obtained data may indicate that the complete- air, due to the refinement of the microstructure, a crit-
ness of phase transformations depends not only on the ical increase in the fraction of ferrite is not observed.
cooling rate and time spent in the temperature range, 9. The obtained data can be used when choosing
where the transformation of ferrite into austenite oc- the modes and a type of weld metal alloying in wet
curs, but also on the size of ferrite grains. underwater welding.

11
S.Yu. Maksymov et al.

10. Recommendations for the range of welding in- 8. Lippold, J.C., Varol, I., Baeslack, W.A. (1994) The influence
put energy values Q = 0.5‒2.5 kJ/mm for 2205 steel of composition and microstructure on the HAZ toughness of
duplex stainless steels at –20 °C. Welding J., Res. Suppl. I,
developed for welding in air and which contribute to 75–79.
obtaining a balanced phase composition of HAZ of 9. Zhou, Y., Zou, D., Li, K. et al. (2018) Effect of cooling time
duplex stainless 2205 steel, should be subjected to on microstructure and properties of 2507 super duplex stain-
correction in wet underwater welding. less steel welding heat-affected zone. Mat. Sci. Forum, 940,
5358.
11. The results of studies on the influence of cool-
10. 10.Grigorenko, G.M., Kostin, V.A., Orlovsky, V.Yu. (2008)
ing rate on the microstructure and phase composition Current capabilities of simulation of austenite transformations
of HAZ of duplex (DSS 2205) steel simulated by the in low-alloyed steel welds. The Paton Welding J., 3, 22‒24.
Gleeble-3800 method correspond only to those TWC 11. Hasui, A., Suga, Y. (1980) On cooling of Underwater Welds.
and those chemical compositions of the base metal Transact. of the JWS, 11(1). April.
12. Geipl, H. (1989) MAGM-Schweissen von Rorrosions bestӓn-
that are used in this study. dign Duplex-Stahlen 22Cr5(9)Ni3Mo. Entfluss von schutz-
References gas-und werfahrenvarianten. Linde – Sonderdruck, 146, Hӓll-
riegels – kreuth.
1. Verma, I., Taiwade, R., R.V (2017) Effect of welding pro-
13. Hu, Y., Shi, Y., Shen, X., Wang, Zh. (2017) Microstructure,
cesses and conditions on the microstructure, mechanical
pitting corrosion resistance and impact toughness of duplex
properties and corrosion resistance of duplex stainless steel
stainless steel underwater dry hyperbaric flux-cored arc
weldments — A review. J. of Manufacturing Processes, 25,
welds. Materials, 10, 1443, www.mdpi.com/journal/materials
134–152.
2. Bermejo, V.M.A., Hurtig, K., Hosseini, V.A. et al. (2016) ORCID
Monitoring thermal cycles in multi-pass welding. In: Proc. of S.Yu. Maksymov: 0000-0002-5788-0753
the 7th Int. Swedish Production Sym. — SPS-16, (Lund, Swe-
den, 25–27 October). Conflict of interest
3. Bermejo, V.M.A., Hurtig, K., Karlsson, L., Svensson, L.E. The Authors declare no conflict of interest
(2017) A step forward in understanding superduplex multi-
pass welds by monitoring thermal cycles. In: Proc. of the 70th CORRESPONDING AUTHOR
IIW Annual Assembly (Shnghai, China, 28 June 2017). S.Yu. Maksymov
4. Bermejo, M.A.V., Daniel, E., Hurtig, K., Karlsson, L. (2019) A
new approach to the study of multi-pass welds microstructure
E.O. Paton Electric Welding Institute of the NASU
and properties of welded 20-mm-thick superduplex stainless 11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine.
steel. http://www.researchgate.net/publication/331715232 E-mail: maksimov@paton.kiev.ua
5. Hosseini, V.A., Karlsson, L., Engelberg, D., Wessman, S.
(2018) Time-temperature — precipitation and property di- Suggested Citation
agrams for super duplex stainless steel weld metals. Weld. S.Yu. Maksymov, G.V. Fadeeva, Jia Chuanbao,
World, 62, 517–533. V.A. Kostin, A.A. Radzievskaya, D.V. Vasilyev
6. Hsienh, R.-J., Liou, H.-Y., Pan, Y.-Ts. (2001) Effects of cool- (2024) Influence of cooling rate on microstructure
ing time and alloying elements on the microstructure of the
Gleeble-simulated heat-affected zone of 22 % Cr duplex and phase composition of HAZ of duplex (DSS)
stainless steels. J. of Mater. Eng. and Performance, 10(5), 2205 steel in wet underwater welding. The Paton
526–536. Welding J., 1, 3–12.
7. Wu, T.-h., Wang, J.-j., Li, H.-b et al. (2018) Effect of heat in-
put on austenite microstructural evolution of simulated heat Journal Home Page
affected zone in 2205 duplex stainless steel. DOI: https://doi. https://patonpublishinghouse.com/eng/journals/tpwj
org/10.1007/s42243-018-0134-z

Received: 06.09.2023
received in revised form: 13.10.2023
accepted: 16.01.2024

12
ISSN 0957-798X THE PATON WELDING JOURNAL, ISSUE 01, January 2024

DOI: https://doi.org/10.37434/tpwj2024.01.02

Hybrid laser-plasma welding:

efficiency and new possibilities (Review)
V.M. Korzhyk1, V.Yu. Khaskin1, E.V. Illyashenko1, S.I. Peleshenko3, A.A. Grynyuk1,
O.A. Babych2, A.O. Alyoshin2, O.M. Voitenko1
1
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
2
“Foreign Trade Office of China-Ukraine E.O. Paton Institute of Welding” Ltd.
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
3
National Technical University of Ukraine “Igor Sikorsky Kyiv Polytechnic Institute”
37 Prospect Beresteiskyi (former Peremohy), 03056, Kyiv, Ukraine
ABSTRACT
Research papers devoted to development of laser-plasma processes during the last two decades are reviewed. It was found
that the current directions of scientific research of the processes of laser-plasma welding are focused mainly on studying the
peculiarities of simultaneous impact of constricted arc plasma and laser radiation with wave length of 1.03–1.07 μm (first of all,
fiber laser) on steels and alloys, as well as investigations of the physical fundamentals of manifestation of the synergic (hybrid)
effect at such an impact and determination of the possibilities of its practical application. It was determined, in particular, that
increase of the effectiveness of synergic effect manifestation is related to improvement of the plasma arc burning conditions in
the zone of ionized vapour plume, which forms under the impact of focused laser radiation, as well as simplification of laser
keyhole formation due to plasma arc pressure.

KEYWORDS: laser-plasma welding, synergic effect, process efficiency, steels, aluminium alloys, industrial application

INTRODUCTION Over the last 10–20 years CO2 lasers have confident-
Ideas of hybrid application of the laser radiation and ly replaced fiber lasers, the radiation of which practi-
the electric arc for welding and related processes, pro- cally does not interact with arc plasma [5]. This has
posed by W.M. Steen, became developed in theoret- significantly changed the way we look at the hybrid
ical and practical works of such prominent scientists laser-plasma process and prospects for its industrial
as U. Dilthey, K. Paul, F. Ridel, I.V. Krivtsun and oth. application. Modern approaches to development of
Modern hybrid welding processes have become ac- welding and related laser-plasma technologies are
cepted by industry to a certain extent. For instance, based on application of radiation with wave length
they are applied in car- and shipbuilding, production in the range of 1.03–1.07 μm, i.e. of fiber, disc and
of pipes of different diameters, etc. In the opinion of a Nd:YAG-lasers.
number of researchers, laser-plasma welding is quite
PURPOSE AND OBJECTIVES OF RESEARCH
promising among other laser-arc processes. There-
fore, the authors propose a review of its state-of-the- The objective of the work is to analyze the current
art to predict its further development. state of the directions of investigations and industrial
applications of laser-plasma welding processes and to
PROBLEM DEFINITION
assess the efficiency of manifestation of the synergic
The authors of works [2‒4] conducted analytical (hybrid) effect at application of laser radiation with
modeling of the processes of laser-plasma welding 1.03–1.07 μm wave length.
and spraying with application of models of integrat- The following tasks were solved to achieve this
ed coaxial heads. In these works, higher efficiency of purpose:
the coaxial laser-arc discharge was attributed to oc- ● establishing modern directions of investigations
currence of a combined laser-arc discharge through of laser-plasma welding processes;
absorption of the CO2-laser beam, passing through
● determination of the efficiency of synergic effect
the center of the arc column, by constricted-arc plas-
manifestation in laser-plasma welding of steels and
ma. Here, the degree of laser radiation absorption by
alloys;
the arc plasma was indicated as a key parameter of
● analysis of laser-plasma process impact on the
discharge control. Such an approach mainly defined
characteristic welding defects in steels and alloys;
the principles of hybrid welding 20–30 years ago.
● analysis of the state-of-the-art of industrial ap-
Copyright © The Author(s) plication of laser-plasma welding.

13
V.M. Korzhyk et al.

ANALYSIS OF LITERATURE SOURCES the surface; II — weld pool surface and III — inter-
Already at the start of the XXI century action occurring directly under the surface. Such fac-
acad. I.V. Krivtsun stated that the main factor for de- tors, as a common welding source, relative position of
termination of the nature of metal penetration in com- the laser and plasma sources, as well as the role and
bined laser-arc welding is the thermal and dynamic influence of welding parameters, have a major impact
influence of the used heat sources on the weld pool on the extent of synergic effect manifestation.
surface. Therefore, he developed a system of equa- In work [9] it is shown that the arc characteristics
tions to describe the process of metal evaporation practically do not change in cases of interaction of “gas
under the impact of multicomponent plasma forming CO2-laser – helium TIG arc” and “disc Yb:YAG-la-
above the weld pool in laser-plasma welding [6]. Such ser – argon TIG arc”. The reason is that the inverse
a system is the base for calculation of the characteris- bremsstrahlung coefficients differ markedly, because
tics of thermal and dynamic influence of arc, laser or of different electron density of argon and helium arcs
combined plasma on the welding pool surface in the and different wave lengths of CO2 and Yb:YAG la-
respective gas-shielded welding processes. His next sers. Such a study to a certain extent promotes partial
step was to study the features of metal penetration in application of the experience of the use of CO2-laser
laser-arc welding using Nd:YAG-laser [7]. A mathe- in hybrid processes with solid-state laser radiation.
matical model of thermal processes developed for this Work [10] presents the results of investigation of
purpose allowed calculating the penetration profiles the synergic effect in hybrid laser–arc welding. Exper-
at a combined influence of the laser beam and the iments were conducted with Nd:YAG-laser of power
electric arc on the product, taking into account their PL = 500 W in combination with standard TIG-weld-
interaction on the metal surface. Calculations showed ing equipment. Two aspects were studied: heat trans-
the presence of a synergic (hybrid) effect, which is fer efficiency and melting efficiency. Heat transfer
manifested in a non-additive increase of the volume efficiency was determined by calorimetric measure-
of metal remelted by laser-plasma process, compared ments, and melting efficiency — by cross-sections of
to metal volumes, remelted separately by the laser and welds produced in different welding modes. Results
plasma processes. show that laser-arc interaction does not lead to any
In order to analyze the synergic combination ef- noticeable change in heat transfer efficiency, but re-
fect, arising during the process, laser-plasma welding sults in a significant increase in melting efficiency.
can be divided into three zones [8]: I — plasma above Non-additive increase of the cross-sectional area of
welds produced with addition of two heat sources (la-
ser and arc) is indicative of the presence of a synergic
effect and hybrid mode of welding.
Spectral analysis of the hybrid plasma plume and
high-speed photographic analysis of the process in hy-
brid welding revealed the following. First, the princi-
ple of the synergy effect consists in that at interaction
with the nonconsumable electrode constricted arc the
laser transfers the electron energy to a higher level,
and creates the conditions for quantum transition. Due
to that more photons are emitted, which increase the
heat input into the material being welded. The syn-
ergic effect is enhanced with increase of laser power
and is decreased with the arc current. This effect is
proportional to the weld cross-section, particularly
in its upper part. Secondly, the amount of spatter at
hybrid laser-arc welding is much smaller than in arc
welding.
Figure 1. Experimental set-up with separate arrangement of the In work [12] a number of investigations of la-
plasmatron and the laser beam [12]: 1 — plasma torch mounted ser-plasma welding were conducted by the scheme in
at angle α = 35° (forward inclined); 2 — plasma nozzle (distance Figure 1. It is proposed to define welding efficiency
to sample L = 2 mm); 3 — laser beam directed at angle β = 20°
ηW as a ratio of the theoretical power PFZ required for
(backward inclined); 4 — cross-jet (air knife); 5 — high-speed
camera; 6 — protective glass; 7 — sample; 8 — direction of work melting the fusion zone material (FZ index), to the
table (sample) movement total input welding power PW according to

14
 Hybrid laser-plasma welding: efficiency and new possibilities

Figure 2. Transverse sections in welding AISI304 steel (δ = 1 mm) by the laser beam (PL = 200 W, ω0 = 200 μm), plasma welding
(Qp = 1.8 l/min; dW = 5 mm) and laser-plasma welding (laser-plasma) (PL = 200 W, ω0 = 200 μm, Qp = 1.8 l/min; dW = 5 mm) with the
respective efficiency values [14]

P ρwch AFZ ∆hFZ A method and model of efficiency determination

ηW= FZ= , (1) were applied in work [14]. While the laser beam of
PW PW
power PL = 200 W and focal point diameter of 200 μm
where ρ is the density of the material being welded; barely melts the material, the process of plasma welding
wch is the movement speed; AFZ is the cross-sectional with arc power of about 2 kW reaches weld penetration
area in the fusion zone, and δhFZ is the required in- through approximately 2/3 of the blank thickness for the
crease of specific enthalpy for melting. Relationship applied set of parameters (Figure 2). A combination of
(1) can be considered as the basis for determination of both the processes provides complete penetration weld-
relative welding efficiency, which compares the effi- ing. While energy combination efficiency ηC rises by just
ciency of combined laser-plasma process with that of ~10 % compared to arithmetic effectiveness of combin-
individual processes. ing energy ηC of individual processes, melting efficiency
A change in arc voltage at introduction of laser ηM of the combined process is approximately 1.5 times
radiation into the plasma-arc process can be one of higher than that of melting ηM of the plasma-arc process.
the causes for improvement of laser-plasma welding It can be assumed that the heat flow in the weld pool,
efficiency compared to individual processes. In case controlled by the conductive and/or convective transfer
of aluminium welding, a marked voltage drop in the mechanisms, changes favourably to create the resulting
range from ‒2 to ‒3 V is observed at switching on cross-section of the weld with increased penetration due
the laser beam. In welding steel under the same con- to more advantageous thermal and/or hydrodynamic
ditions of a highly focused laser beam, a moderate boundary conditions. The authors of work [14] propose
increase of arc voltage between 0.15 and 0.6 V was to regard it as a clear proof of the hypothesis that the sec-
found. Calculations showed [12] that the efficiency of ondary, i.e. thermal effects are responsible for synergic
laser-plasma welding can change from 1.5 (for 6082 advantages of laser-arc treatment efficiency.
aluminium alloy) to 2.4 for AISI304 steel. In work [15] it was determined that synergic effect
If the synergic effect of hybrid laser-arc treatment manifestation depends on welding speed. At 2 m/min
is interpreted as increase of energy transfer from the
speed of welding AISI304 steel (δ = 4 mm) exceeding
heat source to the material, then the thermal efficiency
of the hybrid penetration cross-sectional area is equal
ηT of the process corresponds to the ratio of power
to the sum of areas produced by laser and plasma pro-
PU, which is required for melting the material being
cesses (~2 kW each), and it reaches 30 %, while for
welded per a unit of time (without losses), to total
4 m/min speed it is ~20 %. In work [16] the quan-
applied power PA [13]. In keeping with equation (2),
this value can be divided into melting efficiency ηm titative evaluation of the synergic effect in laser-arc
(energy consumption within the base material) and hybrid welding was performed using a dimensionless
energy combination efficiency ηC (energy input from parameter of melting energy increment ψ:
the heat source), using power PT, transferred from the SH − ( SL + S A )
heat source to the blank [13]: =ψ ⋅ 100% ,
SL + S A
P P P where SH, SL, SA are the cross-sectional areas of welds
ηT = U =ηM , ηC = U T (2)
PA PT PA in hybrid, laser and arc welding, respectively.

15
V.M. Korzhyk et al.

Figure 3. Penetration formation in sheets of S235JR steel 3 mm thick, due to a change of laser radiation power: a — 0; b — 200; c —
330; d — 440 W (unchanged parameters: I = 150 A; V = 1000 mm/min; QP = 0.8 l/min; L = 8 mm, β = 3°) [17]

Figure 4. Penetration formation in sheets of S235JR steel 4 mm thick due to a change in laser radiation power: a — 0; b — 440; c — 0;
d – 440 W and welding speed: a, b — 200; c, d — 250 mm/min (unchanged parameters: I = 150 A; QP = 0.4 l/min, L = 8 mm, β = 19°) [17]

The larger ψ value, the stronger is the syner- bead formation, while laser radiation power is respon-
gic effect. It was calculated that in hybrid laser-TIG sible for penetration depth.
welding ψ = 59.3–83.6 %, and at laser-MIG weld- To achieve the greatest effect from simultaneous ap-
ing ψ = 1–23 %. It can be anticipated that in case of plication of the laser and plasma, specialists of the Insti-
constricted electric arc application in the hybrid pro- tute for Production Technology together with special-
cess the synergic effect will be even greater than in ists of the Institute for Material and Beam Technology
laser-TIG welding [17]. For laser-plasma welding (Dresden, Germany) developed a hybrid laser-plasma
using Nd:YAG laser this effect can be evaluated by head, designed for up to 100 W radiation power and
transverse microsections, given in Figures 3 and 4, in up to 40 A welding current (Figure 5) [23‒26]. During
keeping with the specified mode parameters. investigation of stainless steel welding by this method
For realization of laser-plasma welding processes, it was found [25] that laser beam activation causes an
the focused laser beam can be aimed at the point of abrupt drop of arc voltage by approximately 1 V (Fig-
interaction with the material at a certain angle, i.e. by ure 6, a). This phenomenon was observed only in the
the paraxial scheme (Figure 1) (for instance, [18]), or case of low arc currents. For higher arc currents this
normal to the surface of the product being welded, i.e. effect disappeared (Figure 6, b).
by the coaxial scheme (for instance, [4, 19]). The non- In work [26] it was found that under stable arc
consumable electrode is usually inclined at a certain burning conditions the measured voltage drop after
(minimal possible) angle to the focused laser beam laser beam activation (100 W), is closely related to
axis [20]. Filler wire can be fed in the direction to- shifting of the arc impact zone from the position be-
wards the plasma jet or not fed at all. Metal and alloy hind the beam focal point to a point irradiated by the
powders can be also used as filler materials [21, 22]. laser. In the case of a pure plasma process, the arc is
Influence of arc current predominantly ensures upper deflected backwards, and the anode region evident-
ly lags behind the arc column axis (Figure 7, a). In
the case of a variant with laser radiation, this lagging
behind becomes smaller, and the arc anode region is
stably rooted in the beam focusing zone (Figure 7, b).
At the same time, an increase of arc voltage by 0.4–
0.6 V was observed. The authors of work [26] believe
that the main mechanism of arc stabilization should
be the surface effect, which is unrelated to changes
in arc plasma volume properties, either through direct
interaction of laser radiation and arc plasma, or due to
a possible change of plasma composition as a result of
laser-induced evaporation.
In work [27] a mathematical model was proposed,
Figure 5. Head for laser-plasma microwelding and cutting of thin
metals (radiation power of 100 W, welding current of 40 A) [24] which showed the potential of laser-plasma process

16
 Hybrid laser-plasma welding: efficiency and new possibilities

Figure 6. Arc voltage during bead deposition on a plate from AISI304 stainless steel with laser beam support and without it under
different welding conditions: a — arc current I = 40 A; laser power P = 100 W, welding speed V = 0.75 m/min and sheet thickness
δ = 1 mm; b — arc current I = 160 A; laser power P = 400 W, welding speed V = 2.00 m/min and sheet thickness δ = 3 mm [25]

in terms of peculiarities of the influence of hybrid In work [33] it was established that in laser-mi-
thermal cycles on the material microstructure. The croplasma welding of 7075 alloy (δ = 1.5 mm) the
model was verified by experiments on laser welding volume fraction of remelted metal defects in the form
of car body steels. Work [28] describes laser-plas- of pores of 15–25 μm size decreases, compared to mi-
ma welding of low-carbon steel plates 6 mm thick at croplasma welding, to the level characteristic for laser
up to 5 kW laser power and up to 150 A arc current, welding (~5 %). Remelted metal hardness decreases
which ensured a 100 % increase of the speed of com- by 15–20 % at HAZ metal hardness close to that of
plete penetration welding or increase of penetration the base metal. For comparison, in the laser process
depth by 25–100 %, compared to application of just the remelted metal hardness is decreased by ~15 %,
the laser. It was also found that complete penetration and in the microplasma process it is ~30 % (relative
in laser-plasma welding leads to considerable energy to base metal). Obtained data confirm the advantage
losses, because of its release through the keyhole root. of the laser-microplasma process, proved in [34]. This
All the advantages of the hybrid process are revealed method reduces the use of laser energy to 40–50 %,
only when the keyhole root is enclosed (in the blank). the time of weld pool existence (0.03–0.05 s) be-
comes close to laser welding, and the risk of alloying
Numerical study of the temperature field during
element burn-out is eliminated.
3D printing of thin-walled metal parts by hybrid la-
The laser-plasma method of material treatment can
ser-plasma method shows that the temperature gradi-
be used for thermal surface modification, alongside
ent directly determines the grain growth rate in the
the welding processes, in particular for alloying. In
HAZ of the built-up wall [29]. In work [30] a real-time
work [35] it is shown that the modes of laser-plas-
observation of the parameters of the vapour-gas chan-
ma alloying promote an increase of strength char-
nel and the weld pool at laser and laser-arc welding acteristics (by 20 % on average), compared to laser
was performed. Authors of works [31, 32] showed the alloying. In work [36] the influence of concentrated
good prospects for application of hybrid laser-plasma energy flows on the materials is considered in the
welding method for joining thin sheet (up to 3–4 mm) case of laser-plasma hardening, and the possibility
stainless steels of austenitic and ferritic grades with- of nanostructured layer formation is established. Su-
out filler material application. perthin coatings can be deposited on the part working

Figure 7. Arc shape before (a) and after (b) beam activation: I = 40 A; P = 100 W, V = 1 m/min; material is AISI304 stainless steel
(δ = 1 mm) [26]

17
V.M. Korzhyk et al.

surfaces by an optical pulsed discharge, created by tion to this liquid metal sagging, the conditions for
laser-plasma method [37]. In work [38] it is shown keyhole formation are greatly improved. It can be as-
that at plate surface exposure to a laser heat source, sumed that formation of the synergic (hybrid) effect
an intensive subsurface melt flow (~50 cm/s) forms in the case of application of laser radiation with λ =
in the molten zone, owing to the dominating impact 1.03–1.07 μm, occurs both through improvement of
of the thermocapillary force, generated as a result of a laser radiation absorption by liquid metal, molten by
high temperature gradient (~7000 °C/cm) on the met- the plasma source, and due to formation of weld pool
al pool free surface. This flow, directed from the pool metal sagging under the plasma source impact.
axial part towards the melting front, intensifies energy According to the results of high-speed filming de-
transfer from the pool overheated axial portion to its scribed in work [26] (Figure 8), after activation of fo-
periphery region, and promotes widening of the melt- cused laser radiation, the plasma arc is shortened due
ed zone. Influence of convective stirring of the pool to approaching the zone of the laser plume ionized by
on penetration depth is essentially smaller due to a metal evaporation (i.e. more electrically conductive
predominantly subsurface melt flow. zone). It promotes a shortening of the plasma arc and
DISCUSSION OF THE RESULTS arc voltage drop described in work [26]. In the case of
OF LITERATURE SOURCE ANALYSIS plasma arc immersion into the laser keyhole, its elon-
Welding of steels and alloys by highly concentrated gation can occur, which will lead to a certain increase
heat sources can lead to formation of such characteris- of arc voltage.
tic defects as hot cracks, internal pores, softening of the In case of application of laser-plasma powder hy-
near-weld zone, weld sagging, undercuts and irregular brid welding, energy losses for heat removal into the
nature of reinforcement bead formation [15, 33, 39]. filler material are eliminated [22]. It promoted intro-
One of the advanced methods to eliminate the above duction of such a technology into shipbuilding [41].
defects is application of hybrid laser-arc and laser-plas- Laser-plasma welding without filler material appli-
ma welding processes [39]. In laser-plasma welding the cation is actively used in automotive manufacturing
penetration depth and root bead formation are predomi- [42]. It is applied for manufacturing tailor welded
nantly ensured by the laser component, and elimination blanks, overlap welding of zinc-coated steel (with a
of undercuts and formation of the upper bead are pro- gap), welding using additional material. Welding of
vided by the plasma component [15]. stainless tubes is an example of industrial application
One of the more important aspects of laser welding of laser-plasma welding without filler material [43].
with deep penetration is formation and containment The future of laser-plasma welding as an indepen-
of the laser vapour-gas channel — the so-called key- dent process is associated with development of an in-
hole [40]. The influence of the plasma-component in tegrated head, which combines two energy sources by
laser-plasma welding can be assessed by Figure 4, c. a coaxial scheme [42]. One of the examples of such
From this Figure one can see that even in the absence an integrated welding head is the coaxial head shown
of laser radiation the arc plasma creates a certain sag- in Figure 8, a, which was developed at the Bremen In-
ging due to its own pressure on the weld pool liquid stitute of Applied Beam Technology (Germany) [44].
metal, which is a certain keyhole nucleus [17]. It is This head was later upgraded and fitted with a system
obvious that in the case of laser radiation penetra- of filler wire feed (Figure 8, b) [45]. Another example,

Figure 8. Integrated head for laser-plasma welding: without (a) [44] and with (b) [45] filler wire feed

18
 Hybrid laser-plasma welding: efficiency and new possibilities

Figure 9. 3D-model (a) and appearance (b) of the head for laser and laser-plasma welding, developed at PWI [46]
developed at PWI, is the coaxial head for laser-plasma er), laser-plasma welding can be used for small-scale
welding (Figure 9) [46]. production of such thin-walled products and structures
Conducted analysis of literature data allows defin- from steels and alloys, as regular and profile pipes,
ing the following main advantages of the hybrid la- body elements of automotive and railway transport,
ser-plasma process, compared to laser one: products for food and chemical industry, etc.
● simultaneous use of laser and plasma energy al- One can assume that it is rational to predominantly
lows reducing the laser power and lowering the equip- focus further investigations of laser-plasma welding
ment cost (estimated up to 40–50 %); on relative influence of the radiation of fiber laser and
● plasma component of laser-plasma welding al- constricted arc on steels and alloys. The prospect here
lows lowering the requirements to preparation and fit- is revealing the features, advantages and disadvantag-
up of the edges to be welded and removing the oxide es of such a process with the purpose of establishing
film (for aluminium alloys); the limits of synergic effect manifestation, possibili-
● improvement of productivity due to increase of ties for enhancing it and ways to further use this effect.
welding speed; Conclusions
● reducing energy consumption of the process due
1. Current directions of investigation of laser-plas-
to increase of its efficiency; ma welding processes are focused predominantly on
● widening of the deposited bead in laser-plasma studying the features of simultaneous impact on steels
surfacing and increase of penetration depth in weld- and alloys of the constricted arc plasma and laser ra-
ing due to a change in hydrodynamic flows in the diation with wave length of 1.03–1.07 μm (first of all,
weld pool. fiber laser), as well as studying the physical funda-
Further prospects for development of laser-plasma mentals of the synergic (hybrid) effect manifestation
welding and related processes are associated with ap- under such an impact, and determination of its possi-
plication of fiber lasers (λ = 1.07 μm), as the most ble practical applications. It was determined, for in-
accessible ones for a wide range of users [47]. The stance, that promotion of the synergic effect manifes-
plasma component characteristics are related to the tation is associated with improvement of the plasma
metal being welded (straight polarity for steels and arc burning conditions in the zone of ionized vapour
multipolar asymmetrical current for aluminium al- plume, which forms under the impact of focused laser
loys) [48]. Compared to laser welding, laser-plasma radiation, as well as facilitation of laser keyhole for-
process promotes lowering of the requirements to mation due to plasma arc pressure.
edge preparation, and compared to plasma welding it 2. It was proposed to define the effectiveness of the
lowers the residual deformations [49]. Considering an synergic effect manifestation in laser-plasma welding
increase in productivity, one can anticipate tendencies of steels and alloys as the ratio of the theoretical mag-
of separate laser and plasma welding processes being nitude of power required for melting the weld materi-
replaced by laser-plasma welding in industry. Due to al, to the total input welding power, or as a ratio of the
ensuring rather high speeds (up to 10 m/min and high- cross-sectional area of the laser-plasma process weld

19
V.M. Korzhyk et al.

to the sum of the cross-sectional areas of welds, made aluminum alloys. Optics & Laser Technology, 120, 105766.
separately by plasma and laser welding. It was estab- DOI: https://doi.org/10.1016/j.optlastec.2019.105766
12. Mahrle, A., Schnick, M., Rose, S. et. al. (2011) Process
lished that the efficiency of laser-plasma welding can characteristics of fibre-laser assisted plasma arc welding.
vary from 1.5 to 2.4. Phys. D: Appl. Phys. 44, 345502. DOI: 10.1088/0022-
3. Application of laser-plasma welding allows pre- 3727/44/34/345502
13. Hipp, D., Mahrle, A., Jäckel, S. et. al. (2018) Füssel U. Meth-
vention of such defects characteristic for laser and
od for high accuracy measurements of energy coupling and
plasma welding of high-strength steels and alloys as melting efficiency under welding conditions. J. of Laser Appli-
hot cracks, internal pores, near-weld zone softening, cations, 30, 032414. DOI: https://doi.org/10.2351/1.5040615
weld sagging, undercuts and irregular formation of 14. Hipp, D., Mahrle, A., Beyer, E. et. al. (2019) Thermal effi-
ciency analysis for laser-assisted plasma arc welding of AISI
the reinforcement bead.
304 stainless steel. Materials, 12, 1460. DOI: https://doi.
4. Industrial application of high-speed laser-plas- org/10.3390/ma12091460
ma welding is associated with reduction of laser ener- 15. Korzhyk, V.M., Khaskin, V.Yu., Grynyuk, A.A. et al. (2021)
gy (to ~50 %, compared with laser welding), lowering Features of laser-plasma welding of corrosion-resistant steel
AISI 304 with laser application. The Paton Welding J., 12,
of the requirements to preparation and fit-up of the 9‒17. DOI: https://doi.org/10.37434/tpwj2021.12.02
edges to be welded, increase of welding speed, and 16. Meng, Y., Gao, M., Zeng, X. (2018) Effects of arc types on
minimizing the process energy input. This technolo- the laser-arc synergic effects of hybrid welding. Optics Ex-
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and structures (first of all, from stainless steels, tita- of plasma-laser-hybrid welding process. Physics Procedia, 12,
nium and aluminium alloys), such as structures from 194–200. DOI: https://doi.org/10.1016/j.phpro.2011.03.025
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Features of welding high-strength alloys based on aluminium
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The Authors declare no conflict of interest
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Received: 06.09.2023
received in revised form: 17.11.2023
accepted: 16.01.2024

21
THE PATON WELDING JOURNAL, ISSUE 01, January 2024 ISSN 0957-798X

DOI: https://doi.org/10.37434/tpwj2024.01.03

Impact of heat treatment

on mechanical properties
of joints during electron beam welding
of 2219 alloy
V.V. Skryabinsky, V.M. Nesterenkov, M.O. Rusynyk
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
ABSTRACT
Plates of 2219 alloy of 40 mm thickness were joined by electron beam welding. The effect of sequence of welding and heat
treatment operations on the mechanical properties of the joints and distribution of hardness in the HAZ were studied. It was
established that the plates of 2219 alloy that were quenched before welding and artificially aged after welding, have the highest
strength. Aging improves the ultimate strength of the joints from 300‒315 to 385‒395 MPa, and hardness of the weld and HAZ
metal increases by 5–10 HRB. In the study of the joint microstructure it was found that a region of high-temperature recovery of
the hardened state is located at a distance of about 1.0 mm from the fusion line in the HAZ. It is characterized by an increase in
hardness after aging to the level of base metal hardness. Measuring the welding thermal cycles it was found that the maximum
heating temperature of this region is about 590 °C. Next the annealing region is located. In this region, a decrease in the metal
hardness by 2–3 HRB is observed. For welding speed of 20 mm/s, the HAZ width is about 8 mm.

KEYWORDS: electron beam welding, aluminium alloy, welded joints, heat treatment, mechanical properties, ultimate strength

INTRODUCTION possible or difficult to quench them, only artificial ag-

2219 alloy of Al‒6Cu alloying system is a heat-hard- ing is performed after welding. In work [5] it is shown
enable alloy, the maximal mechanical properties of that postweld artificial aging improves the structure
which are achieved after quenching and artificial ag- and mechanical properties of welded joints.
ing. Due to its excellent mechanical properties, cor- In welding aluminium alloys, there is a certain
rosion resistance and weldability, it has been used correlation between metal strength and hardness,
to produce welded structures for more than 50 years which, in their turn, are determined by its structure
now. However, improvement of welded joint quality [6]. Artificial aging after welding helps revealing the
and stability of their mechanical properties still re- nature of structural transformations and state of the
main a relevant task [1, 2]. Nonconsumable electrode solid solution in the HAZ in such alloys. The degree
argon-arc welding is usually used to join thin blanks of HAZ metal strengthening at artificial aging allows
of 2219 alloy, and over the recent decades friction stir determination of the regions of high-temperature re-
welding has become widely accepted. Electron beam covery of the quenched state, degree of annealing and
welding is used to join thick parts. presence of low-temperature recovery [7]. Changes in
With any fusion welding process, welded joint the metal structure in these regions will be visible in
strength will be lower than that of the base metal. This microsections, and measurement of welding thermal
is related to the presence of a region of remelted metal cycles allows determination of temperatures, at which
and HAZ adjacent to it. Postweld heat treatment of the these changes occur [8].
joints is performed to improve the joint strength.
Mechanical properties of 2219 alloy welded joints OBJECTIVE OF THE WORK
depend not only on heat treatment temperature and The objective of the work is to determine the sequence
time, namely on the sequence of quenching, artificial of quenching, artificial aging and welding operations,
aging and welding performance. Maximal mechani- at which maximal mechanical properties of welded
cal properties of welded joints are achieved by con- joints on large-sized products from 2219 aluminium
ducting full postweld heat treatment (quenching and alloy plates are achieved.
aging). Such an effect of strength increase is observed
MATERIALS
both for welding of quenched and artificially aged [3],
AND INVESTIGATION PROCEDURE
and annealed semi-finished products [4]. In those cas-
es, when large-sizes products are welded, and it is im- The nature of weld formation, hardness distribution
in the welded joint cross-section, macro- and micro-
Copyright © The Author(s) structure in the metal of the weld and HAZ, as well

22
 Impact of heat treatment on mechanical properties of joints during electron beam

Table 1. Chemical composition of 2219 alloy

Weight fraction of chemical elements, %

Al Cu Mn Mg Fe Si Zn Zr Ti

Base 5.8‒6.8 0.2‒0.4 0.02 0.3 0.2 0.1 0.1‒0.25 0.02‒0.1

Table 2. EBW mode parameters

Welding speed, Accelerating Welding current, Focus Scan Beam scanning Beam scanning
mm/s voltage, kV mA pattern amplitude, mm frequency, Hz

20 60 440 Sharp Circle 0.5 630

as mechanical properties of the joints were studied. Mechanical properties of the samples were studied
Investigations were conducted on 2219 alloy plates for three variants of the sequence of welding and heat
40 mm thick. State of delivery is T-351 (quenching + treatment operations: 1 — quenching — artificial ag-
mechanical deformation + natural aging). The alloy ing — welding; 2 — quenching — artificial aging —
chemical composition is given in Table 1. welding — artificial reaging; 3 — quenching — weld-
Welding was conducted in UL-209M unit with ing — artificial aging. Artificial aging was conducted at
power supply from ELA 60/60 source with 60 kV the temperature of 175±5 °C for 19 h with cooling in air.
accelerating voltage. Welding mode parameters (Ta- Ultimate strength was determined by tensile test-
ble 2) were selected so as to ensure complete pene- ing of standard round samples with 4 mm diameter of
tration of the butt in one pass with formation of weld the working part. Impact bend testing was performed
reinforcement and reverse bead. on Charpy samples with a notch in the weld metal.
Hardness of the weld and HAZ metal was measured Samples for testing were cut out across the weld, plac-
by Rockwell instrument with load on the steel sphere ing the weld in the sample center.
P = 600 N. Hardness measurements of the metal of the
INVESTIGATION RESULTS
weld and HAZ were conducted on transverse sections
AND THEIR DISCUSSION
for four variants of the sequence of welding and heat
treatment operations: 1 — quenching — artificial ag- Transverse section of a joint of 2219 alloy plates is shown
ing — welding; 2 — quenching — artificial aging — in Figure 1. During welding, formation of weld rein-
welding — artificial reaging; 3 — annealing — weld- forcement and reverse bead was guaranteed (Figure 2).
ing; 4 — annealing — welding — artificial aging. EBW process proceeded without liquid metal splashing.
Electron beam is a linear heat source and, there- Slight spatter was observed from the weld root side.
fore, the temperature across the plate thickness is re- Welding mode ensured producing narrow welds
garded to be stable. In this case, the thermal cycles in of approximately 2.0 mm width with parallel fusion
EBW of 40 and 10 mm plates will coincide. In order boundaries. Such a shape of the weld promotes reduc-
to simplify the experiments, the thermal cycles were tion of residual welding deformations of the structure.
recorded at EBW of 2219 alloy plates of 10 mm thick-
ness. An EBW mode was selected, which at welding
speed of 20 mm/s, ensured producing a 2 mm wide
weld with parallel fusion boundaries.
Temperature on the plate surface was measured by
chromel-alumel thermocouples, made from 0.1 mm
wires. Thermocouple junctions were caulked into a re-
cess on the plate surface at 2, 4 and 6 mm distance from
the weld axis, which at 2 mm weld width was equal to 1,
3 and 5 mm from the fusion line, respectively. Thermo-
couple readings were recorded by a high-speed record-
ing voltmeter with 100 mm/s speed of tape pulling.
Welded joint structure was revealed by electrolyti-
cal polishing and additional chemical etching in 25 %
aqueous solution of fluoric acid. The microstructure was
examined in an optical metallographic microscope Neo-
phot-32, fitted with Olympus C-500 digital camera. Figure 1. Transverse section of a joint of 2219 alloy plates 40mm
thick

23
V.V. Skryabinsky et al.

Figure 2. Appearance of the weld of a joint of 2219 alloy plates

40 mm thick from the side of electron beam entrance (a) and exit (b)
Results of welded joint hardness measurements are Figure 4. Hardness distribution in the cross-section of joints of
shown in Figure 3. Base metal hardness in the quenched annealed 2219 alloy plates 40 mm thick; 1 — welded joints; 2 —
and aged states is equal to 96 HRB, and weld metal welded joints artificially aged after welding
hardness is 73 HRB. One can see that metal hardness axis (or 1.0, 3.0 and 5.0 mm from the fusion line) were
at 1.0 mm distance from the fusion line is by 1–2 HRB equal to 590, 440 and 300 °C, respectively.
higher than that of metal hardness at 2.0–3.0 mm dis- Weld metal microstructure (Figure 6, a) is dispersed;
tance from the weld. Artificial aging of the joints after it consists of a matrix — aluminium-based α-solid solu-
welding increases weld metal hardness by 10 HRB, and tion and CuAl2 (θ-phase), precipitating along the bound-
metal hardness in the HAZ rises by 3–5 HRB. After aries and chaotically in the grain bulk. CuAl2 (θ-phase)
welding, HAZ width is equal to approximately 8 mm. is the main strengthening phase in alloys of this system.
The fusion line (Figure 6, b) is well-formed, and no de-
When welding annealed plates, weld metal hard-
fects were found on the fusion line. HAZ width is up
ness (Figure 4) is on the level of base metal hardness
to 10 mm from the fusion line. Low-melting eutectic
(72–73 HRB), and the HAZ metal located at 1.0 mm interlayers form in the HAZ region adjacent to the fu-
distance from the fusion line shows the highest hard- sion line. A region of high-temperature recovery of the
ness. When moving away from the fusion line, metal quenched state is located at 0.5–3.0 mm distance from
hardness becomes lower. Artificial aging of welded the fusion line (Figure 3). It is characterized by a hard-
joints strengthens the weld and HAZ metal, and their ness increase after aging to base metal hardness level.
hardness here increases by 5–10 HRB. The annealing region is next. A lowering of metal hard-
Changes in HAZ metal hardness are the conse- ness by 2‒3 HRB is observed in this region.
quence of metallurgical processes, proceeding in the The influence of the sequence of welding and heat
metal under the impact of welding thermal cycle. The treatment operations on the mechanical properties
welding thermal cycles were recorded, in order to de- of 2219 alloy welded joints was studied. Quenched
termine the temperatures, at which these changes oc- plates and plates after full heat treatment (quenching
cur. Experimentally obtained characteristic curves of and artificial aging) were welded. Quenched plates
temperature change during heating and cooling under and part of the plates after full heat treatment were
the impact of the welding thermal cycle are shown subjected to postweld artificial aging.
in Figure 5. Maximal heating temperatures for points
located at 2.0, 4.0 and 6.0 mm distance from the weld

Figure 3. Hardness distribution in the cross-section of joints of

quenched and artificially-aged 2219 alloy plates 40 mm thick; Figure 5. Thermal cycles of points on the surface of 2219 alloy
1 — welded joints; 2 — welded joints reaged artificially after plate at EBW with 20 mm/s speed (L — distance from weld mid-
welding dle; 1 — 2 mm; 2 — 4; 3 — 6)

24
 Impact of heat treatment on mechanical properties of joints during electron beam

Figure 6. Microstructure of weld (a) and HAZ (b‒d) metal at EBW of 2219 alloy plates (a — weld metal; b — fusion line; c —
high-temperature recovery region; d — annealing region), (×500, reduced 2 times)

Table 3. Mechanical properties of joints of 2219 alloy plates 40 mm thick in different initial states of base metal and at further heat
treatment

Kind of treatment Ultimate strength, Relative elongation, Impact toughness,

Before welding After welding σult, MPa δ, % KCV, kgf∙m/cm2

300.0–315.0 3.0–4.0 4.2–4.7

Quenching and artificial aging Without heat treatment
308.7 3.3 4.5

357.0–367.5 2.6–5.7 1.4–1.7

Quenching and artificial aging Artificial aging
361.7 3.6 1.5

385.0–395.0 3.0–3.0 2.9–3.2

Quenching Artificial aging
388.7 3.0 3.0

Note. The numerator gives the minimal and maximal values from 3 measurements; the denominator gives the average values.

Results of welded joint testing for static tension the metal initial condition. As one can see from Fig-
and impact bending are shown in Table 3. ures 3 and 4, in welding heat-hardened plates of 2219
Electron beam welding is characterized by high alloy the HAZ metal is softened, and in welding of
rates of heating and cooling of the metal of the weld annealed plates, on the contrary, the metal strength in
and HAZ. Such cooling rates in EBW of 2219 alloy the HAZ becomes higher.
will promote formation of copper solid solution in the In case of welding plates which have passed the
weld metal. At further artificial aging precipitation of full heat treatment cycle, the ultimate strength of the
strengthening phases and increase of weld metal hard- joints was equal to 300.0–315.0 MPa. It was possible
ness occur, respectively. to increase the ultimate strength to the level of 357.0–
Hardness increase at 1 mm distance from the fu- 367.5 MPa, having conducted artificial aging. Here,
sion line is due to short-time heating of the metal to the impact toughness decreased from 4.2‒4.7 to 1.4–
1.7 kgf∙m/cm2. Postweld artificial aging operation is
quenching temperature and rapid cooling. Maximal
more favourable, compared to aging before welding.
temperature of metal heating is about 590 °C. This
In this case, the ultimate strength of the joints rises to
zone is usually called the zone of high-temperature 385–395 MPa, and impact toughness decreases only
recovery of the quenched state. After conducting arti- slightly to the level of 2.9–3.2 kgf∙m/cm2. Relative
ficial aging, metal hardness is increased here up to the elongation changes only slightly here.
level of base metal hardness in the state after quench-
ing and artificial aging. This zone was earlier revealed Conclusions
only in arc-welded joints [7, 8]. 1. At EBW of quenched plates from 2219 alloy the
Annealing zone, called the low-temperature recov- maximal mechanical properties of welded joints are
ery zone, is located farther from the fusion line. At achieved by conducting postweld artificial aging.
the beginning of this zone the maximal temperature 2. Artificial aging of welded joints of 2219 alloy
was 440 °C, and in the middle part it was approxi- plates increases the hardness of weld and HAZ metal by
mately 300 °C. Hardness of the metal of the weld and 5–10 HRB.
high-temperature recovery zone at EBW of heat-treat- 3. When measuring HAZ hardness of 2219 alloy
ed and annealed plates is the same and independent of joints produced by EBW, a region of high-temperature

25
V.V. Skryabinsky et al.

recovery of the quenched state with hardness increase 7. Lozovskaya, A. V., Chaika, A. A., Bondarev, A. A. et al.
was found at approximately 1 mm distance from the (2001) Softening of high-strength aluminium alloys in differ-
ent methods of fusion welding processes. The Paton Welding
fusion line. After artificial aging the hardness of this J., 3, 13‒17. https://patonpublishinghouse.com/as/pdf/2001/
region is increased to the level of base metal hardness as200103all.pdf
in the heat-hardened state. 8. Lan-Qiang Niu, Xiao-Yan Li, Liang Zhang, Xiao-Bo Liang,
Mian Li (2017) Correlation between microstructure and me-
References chanical properties of 2219-T8 aluminum alloy Joints by VP-
1. Zhang, D.K., Wang, G.Q., Wu, A.P. et al. (2019) Study on the TIG welding. J. Acta Metallurgica Sinica, 30(5), 438‒446.
inconsistency in mechanical properties of 2219 aluminium al- DOI: https://doi.org/10.1007/s40195-016-0516-9
loy TIG-welded joints. J. of Alloys and Compounds, 777(10),
1044-1053. https://www.sciencedirect.com/science/article/ ORCID
abs/pii/S0925838818338568 V.V. Skryabinsky: 0000-0003-4470-3421,
2. Tianyi Zhao, Yue Zhao, Zhandong Wan et al. (2023) “Anneal” V.M. Nesterenkov: 0000-0002-7973-1986,
softening effect of 2219-T8 aluminum alloy joint during weld- M.O. Rusynyk: 0000-0002-7591-7169
ing and its influence on prediction of welding residual stresses.
J. Mater. Research Technology, 24, 5202‒5214. https://www. Conflict of interest
sciencedirect.com/science/article/pii/S2238785423007871 The Authors declare no conflict of interest
3. Zhang, D.K., Wang, G.Q., Wu, A.P. et al. (2019) Effects of
post-weld heat treatment on microstructure, mechanical CORRESPONDING AUTHOR
properties and the role of weld reinforcement in 2219 alumi- V.V. Skryabinsky
num alloy TIG-welded joints. https://www.amse.org.cn/arti-
cle/2019/1006-7191/1006-7191-32-6-684.shtml
E.O. Paton Electric Welding Institute of the NASU
4. Chen, Y. C., Liu, H. J., Feng, J. C. (2005) Effect of post-weld 11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine.
heat treatment on the mechanical properties of 2219-O fric- E-mail: skriabinski.vv.555@gmail.com
tion stir welded joints. J. Mater. Sci., 41(1), 297–299. https://
www.researchgate.net/publication/227050248 Suggested Citation
5. Malarvizhi, S., Raghukandan, K., Viswanathan, N. (2008) V.V. Skryabinsky, V.M. Nesterenkov, M.O. Rusynyk
Effect of post weld aging treatment on tensile properties of (2024) Impact of heat treatment on mechanical prop-
electron beam welded AA2219 aluminum alloy. Int. J. Adv. erties of joints during electron beam welding of 2219
Manuf. Technol., 37, 294–301. https://link.springer.com/arti-
cle/10.1007/s00170-007-0970-7 alloy. The Paton Welding J., 1, 22–26.
6. Rabkin, D.M., Lozovskaya, A.V., Sklabinskaya, I.E. (1992) Journal Home Page
Metals science of aluminium and its alloys. Kyiv, Naukova
Dumka [in Russian]. https://patonpublishinghouse.com/eng/journals/tpwj

Received: 12.10.2023
received in revised form: 15.12.2023
accepted: 17.01.2024

26
ISSN 0957-798X THE PATON WELDING JOURNAL, ISSUE 01, January 2024

DOI: https://doi.org/10.37434/tpwj2024.01.04

Welding current formers

using artificial long lines
O.Ye. Korotynskyi, M.P. Drachenko, A.M. Zhernosekov, I.V. Vertetska
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
ABSTRACT
Application of artificial long lines in arc welding equipment is considered. These lines allow forming pulsed welding currents
of a regulated shape. The need for such currents is determined primarily by the objectives of pulsed technologies, where the
load current shape ensures the required time law of power input into the technological object. The most characteristic examples
of such technologies are pulsed laser welding, laser heat hardening, laser piercing of holes, etc. The possibility of adjustment
of amplitude-time parameters of load current pulses in a broad range allows determination and further on ensuring the optimal
parameters of pulsed electrophysical processes in order to improve their quality and productivity. Special pulsed current gener-
ators are required to create such energy flows, which are described and proposed in the paper.

KEYWORDS: welding current formers, artificial long line, step-down multiphase converter, pilot-arc power source

INTRODUCTION have been used with success for generation of current

Artificial long lines (ALL) are currently used to con- pulses of a regulated shape in the pulse duration range
struct pulsed formers of arc welding currents. The need from several milliseconds to seconds, at load current
for such currents is determined primarily by the objec- levels of several hundred amperes. Such converters
tives of welding technologies, where the load current use modern power transistors, capable of switching
pulse shape ensures the required time law of power hundreds of amperes of current at frequencies of up to
input into the technological object. The most charac- hundreds of kilohertz.
teristic examples of such technologies are pulsed laser When developing generators of rectangular current
welding, laser heat hardening, laser piercing of holes, pulses of a regulated duration, partial discharge of capac-
capacitor discharge resistance welding, etc. Possibili- itive storage is most often used, when a fully-controlled
ty of adjustment of amplitude-time parameters of load power key connects the load to the storage for the time
current pulses in a broad range allows determination which is equal to pulse duration. A serious drawback of
and furtheron ensuring optimal parameters of pulsed such generators is the energy stored in the device signifi-
electrophysical units, in order to improve their quality cantly exceeding the energy evolving in the load during
and productivity. Generation of such currents requires the pulse time, as power key malfunction can result
special pulsed current generators (PCG) [1‒3]. in grace accidents, which can lead to load failure. The
Proceeding from the abovesaid, the proposed work is above disadvantage is overcome, when a fundamentally
devoted to creation and development of versatile pulsed new type — a homogeneous artificial long line (HALL)
current generators designed to construct welding current is used as storage and forming two-terminal network
formers with high characteristics of energy efficiency (FTTN). Here, not only the energy accumulated in the
and electromagnetic compatibility (EMC) [4]. storage is lowered, but also its weight and dimensional
The objective of this work is substantiation and characteristics are reduced.
creation of structures of powerful generators of regu- The proposed approach consists in a combined use
lated pulsed currents (GRPC) and their experimental of ALL-based generators and storages using superca-
investigation. Electrotechnological generators of reg- pacitors (SC) having high energy efficiency and elec-
ulated pulsed current are required first of all as a tool, tromagnetic compatibility characteristics. It allows
which allows experimental investigation of the main creating promising devices for pulsed-arc welding.
parameters of pulsed technological process, including METHODS OF INVESTIGATION
current shape, in order to improve both the process
The following solutions are mainly used when creat-
quality and its efficiency. In addition, pulsed techno-
ing pulsed current generators:
logical units with amplitude-time pulse parameters
● method of partial discharge of capacitive storage
regulated in a broad range are of great interest when
[5, 6];
working under the conditions of small-scale produc-
● ALL application as current formers [7, 8];
tion with frequent changes of product range and kinds
● multiphase step-down voltage converters (chop-
of structural materials. Recently, high-frequency con-
pers) with microprocessor control [9, 10].
verters operating in the pulse-width regulation mode
Figure 1 shows two variants of ALL application for
Copyright © The Author(s) welding equipment. The first variant (a) is a series con-

27
O.Ye. Korotynskyi et al.

Figure 1. Variants of ALL application (for a, b description see the text)

nection of three-phase rectifier, charge switch, ALL and where n is the number of forming line cells; Lc is the
switch of its discharge on the load. The second variant inductance of the forming line cell choke; Cc is the
(b) includes charging rectifier, capacitive storage and capacity of the forming line cell capacitor.
bridge inverter, with ALL connected into its diagonal. Thus, the duration of the pulse front is determined
Let us analyze ALL operation as part of pulsed by the following relationship:
welding current sources.
Based on the considered welding systems, a vari- tf ≈ 0.61 – (Lc – Cc)1/2 = 0.27t/n,
ant of the circuit of inverter-type pulsed current for- and pulse cutoff duration is given by the following
mer was proposed. It incorporates ALL shown in Fig- expression:
ure 2. A chain of three LC elements is used as current tend ≈ (0.075n + 2.3) (LcCc)1/2.
pulse former. The main calculations are used accord-
An operating mockup of pulsed welding current
ing to works [11, 12]:
source was developed and tested according to the dia-
Ip = (U + Uc – Ua)/2ρ; ρ = (Lc/Cc)1/2, gram in Figure 2 (Figure 3).
where Ip is the current pulse amplitude; U is the power Current pulses were generated using the known
source voltage; U is the charge voltage of the forming “skew bridge” circuit, with ALL connected into its di-
line; Ua is the voltage across the arc gap; ρ is the wave agonal in series with pulse transformer Tr2. Short-cir-
impedance. cuited ALL output is connected to Tr1 transformer,
Pulse duration is defined by the following expression: which is used in the circuit of the former of arc stand-
by current. Here, the standby current is formed in the
t = 2.2n(LcCc)1/2, pauses between working current pulses.

Figure 2. Explanatory diagram of ALL element calculation

28
 Welding current formers using artificial long lines

Figure 3. Block diagram of a mock-up of a welding current source based on ALL

As part of the work, a current sensor module was

developed on the basis of a bifilar shunt circuit with a
high level of in-phase interference suppression.
EXPERIMENTAL STUDIES
OF developed MOCKUP
The proposed device, the diagram of which is shown in
Figure 3, provides a stable arc burning in the dynamic
mode, which, in its turn, allows improving the quality
of the welded joint and the energy characteristics of the
device operation. This source was tested in the mode of
welding with pulse modulation of current at the frequen-
cies from 50 to 3500 Hz. Pulse amplitudes were varied
from 50 to 100 A. The lower limit of stable welding cur-
rents was observed at currents higher than 20 A.
Figure 3 shows the oscillograms of operation of a Figure 4. Oscillograms of voltage “a” and current “b” at the
welding device output
forward converter incorporating an ALL.
voltage converters is shown in Figure 5. Adjustment of
APPLICATION OF STEP-DOWN the duration and repetition frequency of current pulses
CONVERTERS AS PULSED WELDING is performed by electronic switch on transistor key T2.
CURRENT FORMERS The long line is charged from capacitive storage
Variants of ALL application in step-down voltage con- Csc, which is connected to charging device (CD).
verters (SDVC), which can be used in multistation weld- Monitoring and adjustment of CD charge parameters
ing complexes, were also studied as part of the performed are performed in keeping with feedback signals, re-
work. An example of ALL application in step-down ceived from current CScd and voltage VScd sensors.

Figure 5. Step-down voltage converter — pulsed current former: CD — charging device of super capacitor battery Csc, CSch.d and
VSch.d — current and voltage sensors; SCB — super capacitor battery; TI and T2 — solid state transistor switches; D1, D2 — normal-
izing diodes; D2, D5 — recovery diodes; Cb — battery capacitance; CSlc and VSlc — sensors of ALL charge current and voltage;
D3.1–D3.N, Lх1‒Lхn, С1‒Сn — ALL elements

29
O.Ye. Korotynskyi et al.

Figure 6. Multiphase SDVC — source of welding current pulses: CD — charging device; CSch.d and VSch — sensors of current and
voltage of charging device control; SCB — capacitive storage based on super capacitor battery; Cb1.1–Cb1.n — buffer capacitors; T1–
Tn — solid state transistor current switches; D1.1– D1.n — normalizing diodes; D2.1 – D2.n — recovery diodes; Т1–Тn, D1.Х– D1.n,
D2.х–D2.n, L1–Ln — elements of step-down converters — welding current pulse formers; CSw.1–CSw.n, VSw — sensors of control
current and voltage of control system of welding current pulse former
Step-down converter, consisting of TI, D1, D2, L1, 2. Good prospects for their application are shown
performs ALL charging in the constant power mode. when powering capacitive energy storages in the
Power monitoring and control are conducted by mode of dynamic burning of the arc.
signals, which are formed by current CScd and volt- 3. New circuits of combined power sources based
age VScd sensors. on artificial long lines were proposed, and their ex-
Pilot arc current source (PACS) operates continu- perimental study in the range of welding current fre-
ously, and its current is summed up on the load (weld- quencies of 50–3500 Hz was conducted. Features of
ing arc) with the main welding current pulses. artificial long line functioning as part of step-down
A variant with formation of welding current puls- converters are considered.
es without ALL application, but using a multiphase 4. As shown by experimental studies, these devic-
step-down voltage converter was also considered es are characterized by high values of energy efficien-
(Figure 6). Here regulation of the duration, frequency cy and electromagnetic compatibility.
and shape of welding current pulses is performed by
synchronous control of transistor switches T1‒Tn. References
The converter is powered from capacitive stor- 1. Anisimova, T.E., Akkuratov, E.V., Gromovenko, V.M. et al.
(1987) High-voltage pulse generator with alternating dura-
age Csc, which is charged from charging device CD.
tion. Pribory i Tekhnika, 4, 45–48 [in Russian].
Monitoring and regulation of SC charge parameters is 2. Opre, V. (2008) Generators of current rectangular pulses. Silo-
performed by feedback signals, coming from current vaya Elektronika, 1, 56–61 [in Russian].
CScd and voltage DVcd sensors. In this circuit the 3. Kazmierkowski, M.P., Krishnan, R., Blaabjerg, F. (2002)
mode of operation with interrupted currents in chokes Control power electronics. USA, Academic Press.
L1‒Ln is used to improve the dynamic parameters of 4. Tihanyi, L. (1995) EMC in Power Electronics. N.Y., IEEE Press.
current pulses. One of the converter channels is used 5. Allas, A.A., Korotkov, A.Yu., Opre, V.M., Fedorov, A.V.
(2001) Charging device. RF Pat. 18026. Publ. 10.05.2001 [in
to generate the arc standby current, and it operates in Russian].
the mode of continuous current of choke Lo. 6. Leonard, W. (1996) Control of Electrical Drives. Berlin,
Comparative analysis of operation of these devices Springer.
shows that preference should be given to multiphase 7. Korotynskyi, O.E., Skopyuk, M.I., Vertetska, I.V. (2021)
converter circuit, as it allows regulation of the time High-efficient sources for arc welding on the base of capaci-
parameters of the amplitude and shape of welding tive energy storage systems. The Paton Welding J., 3, 43‒48
current pulse in a broad range. DOI: https://doi.org/10.37434/tpwj2021.03.08 H
8. Gromovenko, V.M., Opre, V.M., Shchegoleva, N.A. (1997)
Conclusions Charging devices of split capacitive storages. Elektrotekhni-
ka, 3, 46–48 [in Russian].
1. Features of operation and application of welding 9. Korotynsky, A.E. (1999) Peculiarities of operation of high-fre-
current formers based on artificial long lines to create quency welding inverters on the basis of an artificial long line.
resource- and energy-efficient power sources for arc Avtomatich. Svarka, 1, 76–77 [in Russian].
welding were considered.

30
 Welding current formers using artificial long lines

10. Mohan, N., Undeland, T.M., Robins, W.P. (1995) Power elec- CORRESPONDING AUTHOR
tronics: Converters, application and design. USA, NYJohn O.Ye. Korotynskyi
Willey&Sons Inc.
11. Chebotaryov, V.I. (2008) Wave processes in long lines: Mo-
E.O. Paton Electric Welding Institute of the NASU
nographies. Kharkov, Izd-vo im. V.N. Karazin [in Russian]. 11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine.
12. Povh, D., Weinhold, M. (2000) Improvement of power quality by E-mail: epis@ukr.net
power electronic equipment. CIRGE. Paper 13/14/36-06. Paris.
Suggested Citation
ORCID O.Ye. Korotynskyi, M.P. Drachenko,
O.Ye. Korotynskyi: 0000-0002-6461-8980, A.M. Zhernosekov, I.V. Vertetska (2024) Welding
M.P. Drachenko: 0000-0002-4485-2403, current formers using artificial long lines. The Paton
A.M. Zhernosekov: 0000-0002-6404-2221, Welding J., 1, 27–31.
I.V. Vertetska: 0000-0003-4971-7929
Journal Home Page
Conflict of interest https://patonpublishinghouse.com/eng/journals/tpwj
The Authors declare no conflict of interest
Received: 09.10.2023
received in revised form: 23.12.2023
accepted: 15.01.2024

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31
THE PATON WELDING JOURNAL, ISSUE 01, January 2024 ISSN 0957-798X

DOI: https://doi.org/10.37434/tpwj2024.01.05

Nickel scrap recycling

by electron beam melting method
S.V. Akhonin1, V.O. Beresos1, O.G. Erokhin2, O.O. Kotenko2, M.I. Medvedev3, M.G. Lyashenko4
1
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
2
SC “SPC “Titan” of the E.O. Paton Electric Welding Institute of the NASU”
26 Raketna Str., 03028, Kyiv, Ukraine
3
Ukrainian State University of Science and Technologies
2 Lazaryan Str., 49010, Dnipro, Ukraine
4
“DZST” Ltd.
1 Horyanivska Str., Horyanivske Vil., 52035, Dnipropetrovsk Region, Ukraine
ABSTRACT
Experimental-production melting of low-grade nickel scrap was conducted to obtain nickel ingots of not lower than NP2 grade,
which are furtheron used to produce semi-finished products in the form of hot- and cold-formed seamless pipes, including cap-
illary, thin-walled and particularly thick-walled pipes, rings, squares, hexagons, etc. It is shown that during electron beam melt-
ing a significant removal of impurity elements from the nickel scrap occurred, and metal quality began to correspond to nickel
grade not lower than NP2. In order to further study the produced ingot quality, comprehensive research work was performed
on manufacturing semi-finished products in the form of elongated soft rods of 40 mm diameter and wire of 3 mm diameter. It
was determined that mechanical properties of semi-finished products from EBM nickel fully meet the standard requirements.
It is shown that electron beam melting is an efficient method of producing nickel ingots from secondary raw materials, as it
allows ensuring a high level of the produced material quality, and the semi-finished product quality fully meets the standard
requirements by chemical composition, structure and mechanical properties.

KEYWORDS: electron beam melting, electron beam unit, nickel, melting, ingot, scrap, refining

INTRODUCTION and own production. Involvement of secondary raw ma-

Nickel belongs to the group of heavy nonferrous met- terials into the metallurgical production cycle is of tre-
als used both in alloyed steel production and in man- mendous economic importance, as it allows rational use
ufacturing high-tech products in the sphere of aircraft of non-renewable natural resources, reducing the tech-
construction, medicine and electronics [1]. Wide ap- nogenic load on the environment, producing metal by
plication of nickel in different industries is due to its simpler and less costly methods.
unique properties. Nickel addition to alloys increases Therefore, under the conditions of unpredictable
their strength, wear resistance, corrosion resistance, changes in the world markets of metallurgical products,
heat- and electric conductivity, and improves their the manufacturers of structural elements from nick-
magnetic and catalytic properties. el and alloys on its base are faced with an acute issue
In Ukraine proven reserves of nickel ore deposits of improvement of production efficiency and ensuring
are small or depleted, new deposits are insufficiently output of high-quality competitive products. Further in-
explored, but the need in such kinds of raw materi- crease of competitiveness of local products from nickel
als is due to increased demand and industrial progress and alloys on its base due to an essential lowering of the
[2]. Thus, nickel production in Ukraine has limited material and energy costs for its production is a com-
development and it largely depends on price situation plex task. The urgency of solving it for Ukraine is de-
in the world market. termined both by the need of recycling low-grade local
Today, the proportion of secondary raw materials metal scrap from nickel, and by the wish to manufacture
during nonferrous metal manufacture is continuously final products meeting the world standards, as today
growing. Analysis of tendencies in production and con- strict requirements are put forward in the world mar-
sumption of products from nickel and alloys on its base kets for the quality of products made from recycled raw
shows that all the Ukrainian enterprises, operating in this materials. Therefore, one of the main stages of ensuring
market segment, more and more often use metal scrap the final product quality is producing a high-quality in-
as the starting raw material. In the long term it should got as the initial billet for further processing. Here, in
become the main source of producing many nonferrous order to ensure the required level of ingot quality, it is
metals, in particular, nickel, and its efficient recycling necessary to study in greater detail the influence of both
will cover the deficit of balance between consumption the secondary raw material properties and of the techno-
logical parameters of conducting the process. Both the
Copyright © The Author(s)

32
 Nickel scrap recycling by electron beam melting method

yield and the quality of the produced metal depend in

many respects on a rational organization of this process.
Increase of the quality and lowering of the ingot cost
can be achieved due to a detailed study of the processes
of crystallization and formation of casting defects. The
processes of secondary nickel refining at electron beam
melting (EBM), which are still insufficiently studied,
have an important role here. This is particularly relevant
in production of nickel ingots from low-quality scrap,
where impurity content can reach high values.
Therefore, a promising route for development of
Figure 1. Low-grade nickel wastes
enterprises using nickel in their manufacturing is cre-
ation and introduction of high-efficient technologies sumable billet; preparation of the equipment and tech-
of producing cleaned nickel, based on secondary raw nological fixtures for melting; melting process and
materials. control of the produced ingot quality.
A promising direction of modern metallurgy is ap- Low-grade nickel scrap was cleaned from surface
plication of electron beam heat sources for melting, contamination of different origin, compactly packed
refining, surface treatment and other technological into a nonconsumable box and loaded into electron
processes [3]. beam unit UE-208M (Figure 2).
Investigations of the processes of nickel refining The technology of cold-hearth EBM with por-
to remove impurities by electron beam melting were tioned liquid metal feeding into a water-cooled mould
considered in works [4‒7]. (Figure 3) was used to produce ingots of 150 mm di-
EBM is used to produce high-purity ingots of re- ameter (Figure 4).
fractory and highly reactive metals and alloys. As to The following technological parameters were mon-
their quality, EBM ingots are superior to the initial itored during melting: accelerating voltage of electron
material. At EBM of nickel effective removal of gases beam guns, beam currents, rate of initial charge feed-
and other impurities takes place. High vacuum, drop ing into the melting zone, speed of the ingot pulling
transfer and overheating of the metal pool surface in from the mould, and cooling water temperature.
electron beam melting create favourable conditions Numerical values of the technological parameters
for practically complete removal of such impurities as of melting, used at remelting the nickel scrap, are as
As, Zn, Se, Cl, Fe, P, Mg, etc. [4]. follows:
Ingots of EBM nickel have high ductility and are Melting speed, kg/h. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50
easily deformed at room temperatures [7]. Height of portions which are poured into the mould
With the purpose of vacuum refining, EBM is per- at a time, mm. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10
formed in a copper water-cooled mould with a cold Power, kW
hearth by horizontal feeding of the material being in the mould . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
in the cold hearth . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100
remelted. Here, charge materials can have the form
of ingots, lump charge, rods and various wastes, for
instance pressed chips [3].
INVESTIGATION PROCEDURE
Experimental-production melts of low-grade nickel
wastes were conducted in production facilities of SC
“SPC “Titan” of the E.O. Paton Electric Welding In-
stitute of the NAS of Ukraine” together with ”DZST”
Ltd. Company, in order to produce nickel ingots of not
lower than NP2 grade (not lower than 99.5 wt.% nick-
el content). These ingots are further on used to produce
semi-finished products in the form of hot- and cold-
rolled seamless pipes (capillary, thin-walled and particu-
larly thick-walled), rings, squares, hexagons, etc.
Used as the initial charge was low-grade scrap
with average nickel content of 98 wt.% (Figure 1),
which was subjected to double EBR in electron beam
unit UE-208M [8].
The technology of producing ingots in electron
beam unit included the following: forming the con- Figure 2. Initial charge from low-grade nickel scrap

33
S.V. Akhonin et al.

Figure 4. EBR nickel ingot of 150 mm diameter

and the metal quality began corresponding to nickel
grade not lower than NP2.
Comprehensive research work on manufacturing
semi-finished products was conducted at DZST Ltd.
for further study of the produced ingot quality. EBM
nickel ingots of 150 mm diameter were used to pro-
duce soft rods of 40 mm diameter and 3 mm diameter
Figure 3. Process of electron beam remelting of nickel wire (Figure 5).
Produced rods were subjected to heat treatment
At the end of melting, the shrinkage cavity was
(HT) by the following mode: heating up to 800 °C
removed by gradual lowering of the power of heating
temperature in vacuum; soaking for 0.5 h; cooling
the ingot upper end face in the mould.
with the furnace. After HT the rods were machined
The side surface of the produced ingots after cool-
to the required dimensions. Samples for macrostruc-
ing in vacuum to a temperature below 200 °C is clean,
tural studies were taken from the produced rods. No
and a higher concentration of impurity elements on
cracks, delamination, voids, metallic or nonmetallic
the surface in the form of an oxidized or alpha layer is
inclusions were detected in the macrostructure of the
absent (Figure 4). The depth of corrugation-type sur-
produced rods. The macrograin dimensions corre-
face defects is 1–3 mm, defects in the form of tears,
spond to 3‒4 grain size number to GOST 26492‒85.
cracks or discontinuities are absent.
Mechanical properties of rod metal were deter-
Metal of the produced ingots was studied to as-
mined at the temperature of 20 °C after the conducted
sess the depth of refining of low-quality nickel scrap
heat treatment. Table 2 gives the mechanical proper-
during EBM. Determination of chemical composi-
ties of drawn soft rods of 40 mm diameter. As one can
tion of samples taken by their length from the upper,
see from the Table, the mechanical property values
middle and lower parts was conducted by the method
of the produced samples fully meet the requirements
of inductively couple plasma optical emission spec-
of GOST 13083‒77 [9], which is indicative of the
trometry (ICP-OES) in ICP-stpectrometer ICAP 6500
high quality of the metal produced by the developed
DUO. Results of analysis of the produced ingot metal
technology. These data lead to the conclusion that the
showed that significant removal of impurity elements
mechanical properties of semi-finished products from
from nickel scrap occurred during EBM (Table 1),
EBM nickel fully meet the standard requirements.

Table 1. Element content in EBR nickel ingot of 150 mm diameter, wt.%

Not more than

Metal As Bi C Cd Cu Fe Mg Mn P

After double EBR 0.0001 0.0001 0.015 0.0001 0.01 0.04 0.0012 0.0023 0.0001
Norm by DSTU GOST 492:2007 0.002 0.002 0.1 0.002 0.1 0.1 0.1 0.05 0.002

Table 1 (Cont.)

Not less
Not more than
than

Metal Pb S Sb Si Sn Zn Co Ni+Co
After double EBR 0.0001 0.0043 0.0003 0.006 0.0016 0.0001 0.021 99.90
Norm by DSTU GOST 492:2007 0.002 0.005 0.002 0.15 0.002 0.007 0.2 99.5

34
 Nickel scrap recycling by electron beam melting method

Figure 6. Macrostructure of 40 mm dia rod from EBM nickel of

NP2 grade

Table 2. Mechanical characteristics of drawn soft rods of 40 mm

diameter Figure 5. Semi-finished products from EBM nickel of NP2 grade:
a — 3 mm wire; b — drawn soft rods of 40 mm diameter
Grade σt, MPa δ10, % δ5, % 7. Movchan, B.A., Tikhonovsky, A.L., Kurapov, Yu.A. (1973)
EBM nickel NP2 375‒390 28‒32 37‒40
Electron beam melting and refining of metals and alloys.
Kyiv, Naukova Dumka [in Russian].
GOST 13083‒77 >370 >26 >30 8. Akhonin, S.V., Pikulin, A.N., Berezos, V.A. et al. (2019)
Labo­ratory electron beam unit UE-208M. Sovrem. Elektro-
Conclusions metal., 3, 15–22 [in Russian].
9. GOST 13083–77: Rods from nickel and silicon nickel. Spec-
Thus, by the results of the conducted work, it was ifications.
shown that electron beam melting is an effective
method to produce nickel ingots from secondary raw ORCID
materials, as it provides a high level of the produced S.V. Akhonin: 0000-0002-7746-2946,
material quality, and the semi-finished product quality V.O. Berezos: 0000-0002-5026-7366,
fully meets the standard requirements by its chemical O.G. Erokhin: 0000-0003-2105-5783,
composition, structure and mechanical properties. O.O. Kotenko: 0000-0002-0930-9536,
M.I. Medvedev: 0000-0002-1230-420X
References
1. Grytsai, V.P., Bredykhin, V.M., Chervonyi, I.F. et al. (2010) Conflict of interest
Metallurgy of nonferrous metals: Manual. Pt 5. Book 2, Tech- The Authors declare no conflict of interest
nology of copper and nickel. Ed. by I.F. Chervonyi. Zapor-
izhzhiya, ZDIA [in Ukrainian]. CORRESPONDING AUTHOR
2. Pozhuev, V.I., Ivashchenko, V.I., Chervonyi, I.F., Grytsai, V.P. S.V. Akhonin
(2007) Metallurgy of nonferrous metals: Manual. Pt 1. Ed. by E.O. Paton Electric Welding Institute of the NASU
I.F. Chervonyi. Zaporizhzhiya, ZDIA [in Ukrainian]. 11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine.
3. Paton, B.E., Trigub, N.P., Akhonin, S.V., Zhuk, G.V. (2006)
Electron beam melting of titanium. Kyiv, Naukova Dumka E-mail: titan.paton@gmail.com
[in Russian]. Suggested Citation
4. Dmitrenko, A.E., Kozhevnikov, O.E., Pelykh, V.N. (2003)
Application of method of electron beam melting for refin-
S.V. Akhonin, V.O. Berezos, O.G. Erokhin,
ing of nickel. Voprosy Atomnoj Nauki i Tekhniki, 5, 162–166 O.O. Kotenko, M.I. Medvedev, M.G. Lyashenko
[in Russian]. (2024) Nickel scrap recycling by electron beam
5. Azhazha, V.M., Zejdlits, M.P., Shevchenko, S.V., Amonen- melting method. The Paton Welding J., 1, 32–35.
ko, V.M. (1973) Influence of chemically active elements on
properties of nickel of electron beam melting. Izv. AN SSSR. Journal Home Page
Metally, 4, 157–159 [in Russian]. https://patonpublishinghouse.com/eng/journals/tpwj
6. Trigub, N.P., Berezos, V.A., Kornijchuk, V.D., Mosunov, Yu.A.
(2011) Producing of high-quality nickel ingots-slabs by elec-
tron beam melting. Advances in Electrometallurgy, 2, 78–81.

Received: 31.10.2023
received in revised form: 07.12.2023
accepted: 16.01.2024

35
THE PATON WELDING JOURNAL, ISSUE 01, January 2024 ISSN 0957-798X

DOI: https://doi.org/10.37434/tpwj2024.01.06

Effect of the texture

of ferromagnetic Co‒Fe coatings
on their damping capacity
O.S. Kremenchutskyi1, S.S. Polishchuk2
1
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
2
G.V. Kurdyumov Institute for Metal Physics of the NASU
36 Academician Vernadsky Blvd, 03142, Kyiv, Ukraine
ABSTRACT
The effect of the crystallographic texture of Co-Fe coatings produced by the method of electron beam physical vapour deposi­
tion (EB PVD) on their damping capacity (DC) has been studied. It is found that the amplitude dependence of DC of a coating
with a fiber <111> texture exhibits a prominent maximum, while that of a coating with a multicomponent <100> + <111> +
<110> fiber texture shows the blurred maximum which has shifted to the higher amplitude deforma­tions. The effect of both
the fiber texture type and the level of internal (residual) stresses in Co‒Fe coatings on the amplitude dependence of the DC has
been analyzed within the framework of the Smith-Birchak model. It is shown that transition from a single-component to a mul-
ticomponent coating texture reduces the maximum value of DC. In contrast, an increase in the internal stresses in the coatings
leads to a shift and blurring of the DC maximum. On this basis, it is concluded that the maximum DC for Co‒Fe coatings can
be achieved provided that they have a fiber <111> texture and a minimum level of internal stresses.

KEYWORDS: EB PVD, coating, Co‒Fe alloy, damping capacity, texture, internal stresses

INTRODUCTION the type of coatings texture and residual stresses in

Suppression of resonance vibrations (RV) in products their volume.
that are exposed to intense vibrations, for example, in It is known that the microstructure of vacuum con-
blades of gas turbine engines is a prerequisite for pre- densates is largely affected by the temperature of their
venting their premature failure [1]. The use of highly deposition, in particular, at deposition temperatures Td,
damping coatings for such products is considered as for which the temperature Td/Tm = 0.3‒0.5, where Tm
one of the means of reducing the amplitude of RV in is the melting point of condensing metal correspond-
them. The works [2‒4] show the possibility of using ing to the second structural zone [7], which is charac-
ferromagnetic coatings that combine high damping terized by a columnar microstructure of the coating.
capacity (DC) with acceptable mechanical and corro- In this case, the thickness of columnar crystallites de-
sion properties. creases as the temperature of condensate deposition
In ferromagnets with BCC lattice, dissipation of drops. The example of vacuum condensates of Cu [8]
mechanical energy is mainly related with the irrevers- and Ni [9] shows that with a decrease in the thickness
ible shift of the boundaries of 90° magnetic domains of crystallites, their crystallographic texture changes.
in the action of dynamic stresses. In [5], a model dis- Moreover, this is accompanied by the transformation
playing magnetomechanical attenuation (MMA) of of a single-domain structure of the coatings in a mul-
oscillations is proposed, from which it follows that tidomain substructure, when columnar crystallites are
the maximum value of damping is proportional to the fragmented as a result of twins’ formation in them.
magnetostriction of material saturation and the vari- Such changes of microstructure in the coatings of
ation of amplitude dependence of DC characteristics ferromagnetic materials can significantly affect the
is determined by the level and dispersion of residual mobility of magnetic domain boundaries and, accord-
(internal) stresses, which interfere with the movement ingly, their DC.
of magnetic domain boundaries. The mobility of magnetic domain boundaries is
The work [6] shows that the magnitude of mag- also determined by the dislocation density in the vol-
netostriction in the massive textured material chang- ume of a coating. The work [10] shows that an in-
es depending on the direction of action of external crease of dislocation density in the Co‒20 wt.% Fe
stresses on it. Therefore, it can be expected that DC coating as a result of sequential plastic deformation
of coatings of ferromagnetic materials will depend on by a shock ultrasonic wave leads to shifting the am-
plitude maximum of a logarithmic decrement of vi-
Copyright © The Author(s) brations (LD) and reduction in its height.

36
 Effect of the texture of ferromagnetic Co–Fe coatings on their damping capacity

Based on the abovementioned, on the example of

coatings from Co‒Fe alloy, from MMA oscillations
obtained under different conditions of their depo-
sition, the effect of texture on the amplitude depen-
dence of DC characteristics of ferromagnetic alloys
with BCC lattice was studied.
Ferromagnetic Co‒Fe alloy as an object of study
was chosen taking into account its high DC in a wide
temperature range, which is important in terms of de-
veloping damping coatings based on materials of this
class and their practical application [11]. Figure 1. General appearance of substrate of a trapezoidal shape
with the Со‒20 % Fe coating for the study of DC characteristics
EXPERIMENT PROCEDURE of the substrate-coating system
Co‒20 % Fe coating of 90‒120 μm thickness was de- Examination of the coating microstructure was car-
posited by the method of EB PVD on 1.8 mm thick ried out on witness specimens produced in identical
substrates, produced in the form of an elongated conditions of deposition. For this purpose, plates of
trapezoid from the sheet of Ti-6-4 alloy (Figure 1). 5×10 mm of Ti-6-4 alloy were used, which were fixed
The coating was produced in the stationary and non- near the substrate on its holder. Further, these witness
stationary conditions at substrate temperatures of specimens were mechanically cut into two parts, from
350‒500 °C. In the first case, the substrate was fixed which specimens for electron microscopic and X-ray
over the evaporator, in the second case it rotated examinations were made. Trapezoidal substrates with
around its longitudinal axis at a speed of 80 rpm in coating were not subjected to any treatment and were
the process of coating formation. used further to determine the characteristics of their DC.
Characteristics of DC of coating material (am- Figure 2 presents the overall appearance of the mi-
plitude dependences of LD) were investigated in the crostructure and the distribution of chemical elements
laboratory installation described in [12]. The initial over the thickness of the witness specimen cross-sec-
values of LD of the substrate-coating system were tion. It is seen that the microstructure of the material
measured in the mode of freely attenuated bending is homogeneous, and there are no defects on the inter-
oscillations with a frequency of 130‒150 Hz. The in- face of the substrate with coating that contribute to the
trinsic of LD of the coating material were determined reduction of adhesion between them.
by the procedure described in [13], based on the ini- For X-ray structural analysis of Co‒Fe coatings,
tial data for the specimens with coatings and without the DRON-4M diffractometer in the radiation of the
them. Such approach allows eliminating nonuniform copper anode was used. Figure 3 presents the diffrac-
distribution of deformation on the surface of the spec- tion patterns of coatings, deposited under different
imen, i.e., it represents deformation of the coating ma- conditions. It is seen that on both diffraction patterns,
terial in the approximation of pure bending, and also only maxima of BCC structure are present. Such a
makes it possible to eliminate the effect of coating fact draws attention, that the ratio of intensities of
thickness on its DC. diffraction peaks changes during the transition from

Figure 2. General appearance of microstructure (a) and distribution of chemical elements (b) according to the thickness of witness
specimen cross-section with the Co‒20 % Fe coating

37
O.S. Kremenchutskyi et al.

Figure 4. Amplitude dependences of LD for substrates with

Co‒20 % Fe coatings deposited on a stationary substrate (1), sub-
Figure 3. Diffraction pattern of Co‒20 % Fe coatings deposited strate during rotation (2) and for a substrate without coating (3)
in the stationary conditions (a) and during substrate rotation (b)
Figure 5 presents the calculated amplitude depen-
the coatings produced in the stationary conditions and dences of intrinsic of DC for the coating material pro-
during substrate rotation.
duced in the stationary and nonstationary conditions
The analysis of the crystallographic texture of the
of deposition. It can be noted that the level of DC of
coatings was performed using an X-ray diffractometer
the coating produced in the stationary conditions is
DRON-3, equipped with a textured attached device, in
almost twice higher than that of the coating deposited
CuKα radiation. The measurements were carried out us-
in the nonstationary conditions. From the comparison
ing a parallel beam geometry at scanning angles from 0
to 80° and from 0 to 360° for α and β, respectively. The of the shape of the curves of amplitude dependences,
data obtained on a nontextual BaTiO3 specimen were it is seen that for the coating deposited in the station-
used to record the defocusing effect. The analysis of ary conditions, the rate of decrease of the curve for the
crystallographic texture was carried out by constructing descending part of the maximum is sharper.
straight and reverse pole figures (PF) by means of the The height and profile of the MMA maximum are
MTEX Matlab software package [14]. determined by saturation magnetostriction and residu-
al stresses [5]. Taking into account that the magnitude
DAMPING PROPERTIES OF COATINGS of the magnetostriction of the material depends on its
Figure 4 presents amplitude dependences of LD re- texture [6], it was assumed that the differences in the
flecting DC of substrate-coating systems, obtained in amplitude dependence curves of DC of Co‒Fe coat-
the conditions of stationary and nonstationary depo- ing materials produced under different conditions of
sition of coatings. It is seen that DC of both oscillat- their deposition (Figure 5) are predetermined by their
ing systems is several times higher than that of the different texture.
substrate without coating. Moreover, in the case of
coating deposited with a stationary substrate, the level MICROSTRUCTURE OF COATINGS
of DC is significantly higher compared to the coating The characteristic microstructure of the Co‒20 % Fe
deposited on a rotating substrate. coating is shown in Figure 6. It is seen that the coating

Figure 5. Amplitude dependences of energy loss coefficient (ψ = 2δ, where δ is the intrinsic of LD) of intrinsic (a) and values normal-
ized to the maximum (ψmax) (b) for Co‒20 % Fe coatings deposited in the stationary (1) and nonstationary (2) conditions

38
 Effect of the texture of ferromagnetic Co–Fe coatings on their damping capacity

consists of columnar grains oriented perpendicular to

the surface of the substrate.
It turned out that such coatings are characterized
not only by the elongation of grains in the direction
of their growth, but also by the presence of a certain
predominant crystallographic orientation. In Figure 7,
a, PF (110), (100) and (211) are presented, built for
the Co‒20 % Fe coating, deposited in the stationary
conditions. It is seen that the distribution of density of
the poles (110) and (100) has a circumferential char-
acter. Taking into account the angular distance of the
circumferential distributions, it was concluded that
this type of pole density distribution can be obtained
in the case of fiber texture with a predominant orien- Figure 6. Cross-sectional microstructure of Co‒20 % Fe coating
tation of crystallites in the <111> direction. Figure 7, etched to reveal grain boundaries
b presents PF of the Co‒20 % Fe coating deposited in
the nonstationary conditions. It is seen that also in this and in a coating produced in the nonstationary condi-
case, a fiber texture is formed. However, the grains tions, it is a multicomponent <100> + <111> + <110>.
are mainly oriented along the <100> axis, where the Moreover, the volume fractions of the components
maximum pole density is observed in the center of PF. differ. The largest volume fraction is characteristic of
To evaluate the volume fraction of crystallites the component of the fiber texture of <100> type and
characterized by different orientations based on the the smallest is typical of <110> component (Table 1).
obtained results on the distribution of pole density, in- It is seen that, despite the presence of <111> and
verse PF were built. Figure 8 presents inverse PF built <110> components, the <100> texture component is
for the Co‒20 % Fe coatings deposited in the station- dominant in the multicomponent texture of a coating
ary and nonstationary conditions. It is seen that in the deposited in the nonstationary conditions. A coating
case of coating deposited in the stationary conditions, deposited in the stationary conditions has a one-com-
the fiber texture is a single-component of <111> type, ponent fiber texture <111>.

Figure 7. Distribution of pole density for Co‒20 % Fe coatings deposited on the surface of a titanium plate in the stationary (a) and
nonstationary (b) conditions: 1 — (110); 2 — (100); 3 — (211)

39
O.S. Kremenchutskyi et al.

Figure 8. Inverse PF built in the direction perpendicular to the surface of Co‒20 % Fe coating deposited in the stationary (a) and non-
stationary (b) conditions
EFFECT OF TEXTURE loss coefficient ψ for the coatings with different fiber
ON THE DC AMPLITUDE DEPENDENCE textures. According to the Smith–Birchak model, this
OF COATINGS dependence is determined by the expression:
According to the Smith‒Birchak model [5], the height
ψ = (2KEλs/σi){[1 – exp(–2x)×
of the MMA maximum in the materials with BCC lat- (2)
×(1 + 2x + 2x2)]/x2},
tice is determined by the dependence:
3KE λs where x = σ/σi; σ is the amplitude of alternating stress-
 ∆U 
 U=  2∆σi
× es of the oscillating specimen.
 max
In [2] it is shown that a satisfactory correspon-
2/3  (1)
  ∆σ   ∆σi    dence between the experimentally measured values
× 1 − 1 − i  1 + σ   , of DC of the specimen for different oscillation ampli-
  σi   i  

tudes and values calculated by the formula (2) can be
where K is the constant that depends on the shape of obtained in the condition that the value of the internal
the hysteresis loop; E is the modulus of elasticity; λs is stresses is σi = 17.5 MPa.
the saturation magnetostriction; ∆σi is the dispersion In Figure 9, a, the amplitude dependences of the
of value of internal stresses; σi is the average value of energy loss coefficient of the coating material with dif-
internal stresses. ferent types of texture are given. It is seen that when
In [6] it was shown that the value of λs is deter- the type of fiber texture changes, the height of the DC
mined by the type of fiber texture of the material and maximum changes: the largest value is observed in the
the direction of application of alternating deforma- case of the fiber texture of <111> type, and the smallest
tions. Based on the obtained data on the texture of the value is <100>. However, the shape of the amplitude
coatings and taking into account the direction of their dependence of DC for the coatings with different tex-
deformation during oscillations of flat specimens, the ture remains unchanged (Figure 9, b).
values of λs for the coatings with different texture (Ta- Comparing these dependencies with the experi-
ble 2) were calculated, using the procedure [6] and mental results obtained for the coatings with different
experimental values of magnetostriction for Co‒Fe types of fiber texture (Figure 5), it can be assumed
alloy along the crystallographic <100> and <111> di- that a change in the type of fiber texture of the coat-
rections [15]. ing can only lead to a decrease in the level of DC.
To evaluate the influence of the texture type, let At the same time, the experimental amplitude depen-
us calculate the amplitude dependence of the energy dences of the energy loss coefficient for the coatings
produced under different conditions show not only a
Table 1. Characteristics of fiber textures of specimens produced
decrease in the value of the DC maximum, but also
in the process of deposition of Co‒20 % Fe alloy on titanium sub-
strates Table 2. Magnitude of magnetostriction of Co‒20 % Fe coating
saturation with different types of fiber texture under its tension/
Conditions of coating Volume fraction of texture components compression
deposition <110> <100> <111>
Type of fiber texture of coating <100> <110> <111>
Stationary 0.0 0.0 1.0
Magnitude of saturation
93.9 119.1 127.5
Nonstationary 0.16 0.6 0.24 magnetostriction, 10‒6

40
 Effect of the texture of ferromagnetic Co–Fe coatings on their damping capacity

Figure 9. Amplitude dependences of energy loss coefficient of coatings with fiber textures of <100> (1), <110> (2) and <111> (3) type
at the same level of internal stresses (a) and normalized to the maximum value ψmax (b)

Figure 10. Amplitude dependences of the energy loss coefficient of coatings with a fiber texture of <111> type at different levels of
internal stress σi, MPa: 1 — 17.5; 2 — 25; 3 — 30; 4 — 40 (a) and normalized to the maximum value ψmax (b)
its shape (the maximum is blurred towards larger de- of the maximum on the amplitude dependence of DC
formation amplitudes). Therefore, it was assumed that is observed, but also its blurring, it can be assumed
such a phenomenon may be associated with a change that such changes are predetermined by an increase
in internal stresses in the coatings produced under in the level of internal stresses in the coatings with a
different conditions. To find out this possibility, the multicomponent texture.
amplitude dependences of the energy loss coefficient
Conclusions
of the coating material with different levels of internal
stresses were calculated. 1. DC of titanium plates with the coatings of ferro-
From the calculated amplitude dependences of the magnetic Co‒20 % Fe alloy changes depending on
energy loss coefficient of the coating material with the the conditions of coating deposition. In the station-
same type of fiber texture, but with different levels of ary conditions of their deposition, the characteristics
internal stresses presented in Figure 10, it is seen that of DC of the substrate-coating system are described
when the internal stresses grow, the height of the peak by a curve with a maximum, and in the nonstationary
decreases, shifts and expands towards larger stress conditions, they are described by a curve with the sat-
amplitudes during a alternating deformation. uration on the side of large amplitudes of oscillations.
According to the obtained modeling results, it can 2. Intrinsic of DC of the Co‒20 % Fe coatings
be assumed that when the conditions for coating pro- formed in the stationary conditions are approximately
duction change, variation in their amplitude depen- twice as large as those of the coatings produced in the
dence of DC is mainly predetermined by the change nonstationary conditions.
in the fiber texture of the coatings from a single-com- 3. It was determined that the conditions of deposi-
ponent one of <111> type, which is formed in the sta- tion of the Co‒20 % Fe coatings affect the character-
tionary conditions of deposition, to a multicomponent istics of the coating material texture. In the stationary
<100> + <111> + <110>, formed in the nonstationary conditions of deposition, a fiber texture of <111> type
conditions of deposition. However, since, as is seen is formed, and in the nonstationary conditions, a mul-
from Figure 5, for the coatings deposited in the non- ticomponent texture of <100> + <111> + <110> type
stationary conditions, not only a decrease in the height is formed.

41
O.S. Kremenchutskyi et al.

4. The level of DC of the Co‒20 % Fe coatings 10. Ustinov, A.I., Movchan, B.A., Skorodzievskii, V.S. et al.
with a fiber texture of <111> type is predetermined (2004) Effect of thermomechanical treatment onto damping
capacity Co–20 % Fe coatings. Vibr. Tekh. Tekhnol., 3, 104–
by a high value of the magnetostriction magnitude, 106 [in Russian].
which is consistent with the Smith-Birchak model for 11. Herman Shen, M.-H. (2008) Free layer blade damper by
magnetomechanical damping. magneto-mechanical materials. United States. Pat. WO
5. A decrease in the level of damping in the 2008/127375 A1.
12. Ustinov, A.I., Nekrasov, А.А., Perederiy, V.A. et al. (2012)
Co‒20 % Fe coatings with a multicomponent fiber
Device for dissipative properties research of metallic flat sam-
texture of <100> + <111> + <110> type can be a con- ples and coatings. Zavod. Laboratoriya, 10, 41–44 [in Rus-
sequence of both a decrease in the average value of sian].
the magnetostriction magnitude as well as an increase 13. Ustinov, A.I., Skorodzievskii, V.S., Kosenko, N.S. (2007) A
in the level of internal stresses. study of the dissipative properties of homogeneous materials
deposited as coatings. Pt 1. Method for the determination of
References the amplitude dependence of the true vibration decrement of
1. Matveev, V.V. (1985) Vibration damping of deformed bodies. the coating material. Strength Mater., 39(6), 663–670. DOI:
Kyiv, Naukova Dumka [in Russian]. https://doi.org/10.1007/s11223-007-0076-3
2. Yen, H.-Y., Herman Shen, M.-H. (2001) Passive vibration 14. Hielscher, R., Schaeben, H. (2008) A novel pole figure in-
suppression of beams and blades using magnetomechanical version method: Specification of the MTEX algorithm.
coating. J. of Sound and Vibration, 245(4), 701–714. DOI: J. Appl. Cryst., 41, 1024–1037. DOI: https://doi.org/10.1107/
https://doi.org/10.1006/jsvi.2001.3561 s0021889808030112
3. Torvik, J., Langley, B. (2015) Material properties of hard coat- 15. Noro, S., Ohtake, M., Kawai, T. et al. (2022) Magnetostric-
ings developed for high damping. In: Proc. of 51st AIAA/SAE/ tive properties of Co–Fe alloy epitaxial thin films with Co-
ASEE Joint Propulsion Conf. (Orlando, Florida, USA, July rich composition. AIP Advances, 12, 035144. DOI: https://
29, 2015), 4195. DOI: https://doi.org/10.2514/6.2015-4195 doi.org/10.1063/9.0000352
4. Ustinov, A.I., Movchan, B.A., Lemke, F., Skorodzievskii, V.S. ORCID
(2001) Damping capacity of Co‒Ni and Co‒Fe coatings pro-
duced by electron-beam deposition. Vibr. Tekh. Tekhnol., 4, O.S. Kremenchutskyi: 0000-0001-7650-0122,
123–126 [in Russian]. S.S. Polishchuk: 0000-0002-8403-5360
5. Smith, G.W., Birchak, J.R. (1969) Internal stress distribution
theory of magnetomechanical hysteresis-an extension to in-
Conflict of interest
clude effects of magnetic field and applied stress. J. Appl. Phys., The Authors declare no conflict of interest
40, 5174–5178. DOI: https://doi.org/10.1063/1.1657370
6. Frank, R.C., Johnson, B.G., Schroeder, C.W. (1969) Crystal
CORRESPONDING AUTHOR
orientation and magnetomechanical damping of torsional O.S. Kremenchutskyi
vibrations. J. Appl. Phys., 40, 3189–3192. DOI: https://doi. E.O. Paton Electric Welding Institute of the NASU
org/10.1063/1.1658164 11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine.
7. Movchan, B.A., Demchishin, A.V. (1969) A study of the
E-mail: kremens44@gmail.com
structure and properties of thick vacuum condensates of nick-
el, titanium, tungsten, aluminum oxide and zirconium diox- Suggested Citation
ide. Fiz. Metall. Metalloved., 28(4), 653–660 [in Russian]. O.S. Kremenchutskyi, S.S. Polishchuk (2024) Effect
8. Ustinov, A.I., Fesyun, E.V., Melnichenko, T.V., Romanenko,
S.M. (2007) Effect of substrate temperature on micro- and of the texture of ferromagnetic Co‒Fe coatings on
substructure of copper condensates deposited from a vapor their damping capacity. The Paton Welding J., 1,
phase. Advances in Electrometallurgy, 4, 18–24. 36–42.
9. Ustinov, A.I., Skorodzievskii, V.S., Fesiun, E.V., Taranenko,
V.N. (2012) Structure and mechanical properties of nano- Journal Home Page
structured vacuum nickel condensates. Nanosystemy, Nano- https://patonpublishinghouse.com/eng/journals/tpwj
materialy, Nanotekhnologii, 10(1), 11–18 [in Russian].

Received: 14.11.2023
received in revised form: 07.12.2023
accepted: 22.01.2024

42
ISSN 0957-798X THE PATON WELDING JOURNAL, ISSUE 01, January 2024

DOI: https://doi.org/10.37434/tpwj2024.01.07

Nitrogen absorption by 04Cr18Ni10 steel

in plasma-arc melting under slag
of CaO‒Al2O3 system
V.O. Shapovalov1, V.G. Mogylatenko1,2, R.V. Lyutyi2, R.V. Kozin1
1
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
2
National Technical University of Ukraine “Igor Sikorsky Kyiv Polytechnic Institute”
37 Prospect Beresteiskyi (former Peremohy), 03056, Kyiv, Ukraine
ABSTRACT
Nitrogen as an alloying element of steel is a strong austenitizer, and it has an essential influence on mechanical properties of
steels of different grades. It is rational to use gas phases for nitrogen alloying, and the process can be intensified with the ap-
plication of highly-concentrated energy sources, for instance, plasma. One of the determining factors of metal alloying is the
partial pressure of nitrogen and the process temperature. It is difficult to find works, dealing with nitriding of metal melts in
plasma-slag process. The paper gives experimental data on dissolution kinetics and nitrogen solubility in 04Cr18Ni10 steel.
The derived mathematical models of nitrogen dissolution in plasma melting are highly significant, which is indicated by the
respective determination coefficients. At less than 0.1 atm partial pressure of nitrogen above the melt, the temperature in the
range of 1823‒2323 K practically does not influence the content of nitrogen in steel, and at a higher partial pressure, the content
of nitrogen in steel decreases with the temperature rise. The melt temperature under the experimental conditions was assessed
as 2385 K. It was determined that the coefficient of nitrogen distribution between the metal and slag changes only slightly at up
to 1 atm partial pressure of nitrogen and is equal to 1.1‒1.2.

KEYWORDS: nitrogen-containing steels, plasma-slag melting, nitrogen, absorption kinetics, solubility, distribution coeffi-
cient

INTRODUCTION times stronger than the effect of nickel. It is known

04Cr18Ni10 steel, along with other chromium-nickel that 0.15 % of nitrogen in chromium-nickel steels is
steels, is one of the most common structural austenitic equivalent to 2‒4 % Ni, and 0.25 % of nitrogen is
steels used in industry, as its products can operate in equivalent to 2.5‒6.0 % Ni. Considering the cost of
various aggressive environments (solutions of salts, nickel, substitution of its part with nitrogen signifi-
alkalis, acids) and has an operating temperature of up cantly increases the cost-effectiveness of stainless
to 900 K under the normal conditions and up to 600 K steel production [3].
in aggressive environments. Nitrogen-containing steels with an over-equilibri-
Any high-temperature metallurgical process of um nitrogen content should be melted under the ex-
steel production is a process of interaction of several cess nitrogen pressure or by the plasma-arc remelting
phases. As a rule, it is liquid metal, liquid slag and (PAR) method with alloying of metal with nitrogen
gas phase or vacuum. The main gases interacting with directly from the gas phase. Arc-slag remelting (ASR)
the slag and metal are oxygen, hydrogen and nitrogen. [4] allows performing alloying of metal with nitrogen
The presence of the first two gases in the gas atmo- from the gas phase and treatment of metal with slag.
sphere in the absence of melt protection inevitably Thus, the formation of ingots of various cross-sections
leads to the formation of defects in ingots and cast- is provided, that have satisfactory surface, chemical
ings: nonmetallic inclusions, gas and gas shrinkage and structural homogeneity. The nitrogen content in
porosity, flakes. And only nitrogen can perform both a steel is determined by the composition of the slag and,
negative and a positive function. The first one consists depending on the method of melting, changes both in
in the formation of strengthening inclusions in steel the metal and in the slag.
under certain conditions, and the second one is in the The process of alloying steels with nitrogen from
fact that nitrogen is a substitute for nickel. plasma is characterized by a high rate of gas absorp-
Nitrogen, known for a long time [1, 2] as an alloy- tion by the liquid metal, which is an order of magni-
ing element of chromium-nickel and chromium-man- tude higher than in conventional metallurgical units.
ganese steels, is a strong austenizer along with nickel. Therefore, plasma-slag remelting (PSR), which com-
The effect of nitrogen in the γ-region of iron is several bines plasma as a source of metal heating and acti-
vation of nitrogen molecules, with slag treatment to
Copyright © The Author(s) refine the melt from undesired impurities and nonme-

43
V.O. Shapovalov et al.

tallic inclusions, is more promising from the point of [ N ]Me = KN PN ,

view of obtaining an superequilibrium nitrogen con- 2

tent in the metal. where [N]Me is the equilibrium solubility of nitrogen,

Taking into account that the absorption of nitrogen %; KN is the Sieverts’ constant, %∙atm‒1/2; PN is the
2
by liquid steel in the ternary gas–slag–metal system equilibrium partial pressure of nitrogen in the gas
depends on the transfer capacity of the slag, the prob- phase, atm.
lem of studying the kinetics of the process of nitrogen During plasma heating, the square root law for ni-
absorption by the slag and through the slag by metal trogen absorption is also fulfilled, but here the pro-
during plasma-slag melting on the example of steel portionality factor is much higher than the Sieverts’
04Cr18Ni10 arises. constant. Alloying of metal with nitrogen will occur if
STATE OF THE PROBLEM the inequality is observed [10]:

( )
2
Nitrogen significantly affects the mechanical proper- [ n ]me / Kn
pn > Pn = ,
ties, phase stability, corrosion and heat-resistant prop- 2 2

erties, crack resistance and fatigue strength of ledebu- where pN is the partial pressure of nitrogen in the gas
2
rite tool and high-speed steels, stainless steels of the phase.
austenitic, ferritic-austenitic, ferritic-martensitic and The temperature dependence of the nitrogen dis-
martensitic grades [5–8]. The effect of nitrogen on the solution constant in a liquid iron has been studied by
strength of austenitic steels is stronger than the effect many researchers. The most probable results are satis-
of carbon (Table 1). factorily described by the equation [11]:
Melting in a nitrogen-containing gas environment, ( −293 / T ) − 1.16.
lgKN =
especially at elevated pressure, allows alloying with 2

nitrogen directly from the gas phase. Plasma-arc re- According to the specified dependence at 1600 °С,
melting (PAR) provides a higher efficiency of alloying KN = 0.048 %. Thus, for this temperature we can
2
steel with nitrogen at a low gas pressure compared to write:
pressure melting [9]. For chromium-manganese steels
scarcely alloyed with nickel, such as Cr21Mn10NNi4,
[ n ]me = 0.048 pn .
2

plasma-arc melting at nitrogen partial pressures of

Alloying elements and impurities in iron change
60‒120 kPa allows obtaining a nitrogen content that
the ability of iron to dissolve nitrogen [12].
exceeds its standard solubility. The content of nitro-
In real steelmaking processes, nitrogen dissolution
gen in ingots during PAR is regulated by the change in
occurs in parallel with oxidation processes. A flow of
the partial pressure of nitrogen in the plasma-forming
oxygen causes a counter flow of nitrogen. It was estab-
gas, the total pressure in the melting chamber of the
lished [13] that the behavior of nitrogen in Fe‒C alloy
furnace, and the rate of ingot extraction, i.e., the melt
and killed steel (0.3 % C) and 15MnAlTi steel as well
temperature [9].
as pure iron is determined by the rate of metal saturation
Thus, the partial pressure of nitrogen in the plas-
with oxygen and when the maximum concentration of
ma-forming gas and the process temperature are one
oxygen in metal is reached, the absorption of nitrogen
of the determining factors of alloying metal in an su-
by steels stops. The higher the rate of oxygen absorption,
perequilibrium amount with nitrogen during PAR.
the sooner nitrogen absorption stops and the lower its
SOLUBILITY OF NITROGEN IN METAL MELT concentration is achieved in steels.
The equilibrium solubility of nitrogen in the metal The solubility of nitrogen in steels can be calcu-
melt is subjected to the Sieverts’ law or the square lated theoretically having known values of interaction
root law: parameters [10]. In [14], a thermodynamic model for
predicting the solubility of nitrogen in liquid stainless
Table 1. Influence of alloying elements (1 %) on yield strength of steels depending on the concentrations of alloying
04Cr18Ni10 steel [2] elements, temperature and pressure is given, which
takes into account a new factor of the pressure effect
Yield strength of
Alloying element Type of solution
steel, kg/mm2
on the coefficient of nitrogen activity. The results of
the calculations, which agree well with the experi-
N Penetration 70
ment, are subjected to the Sieverts’ law. However, at a
C ‒»‒ 40 high pressure (> 1 atm) and especially at a high con-
Cr Substitution 0.4 centration of alloying elements, a deviation from the
Sieverts’ law occurs and the nitrogen pressure begins
Mo ‒»‒ 1.5
to negatively affect the solubility.

44
 Nitrogen absorption by 04Cr18Ni10 steel in plasma-arc melting under slag

SOLUBILITY OF NITROGEN IN SLAGS ments due to the formation of nitrides, the maximum
As for the sorption capacity of slag in relation to im- solubility of nitrogen in the slag usually does not ex-
purities, it depends on the oxidation-reduction poten- ceed 2 % [15].
tial of the medium, which can be determined by the Unfortunately, in the scientific literature, absorp-
equilibrium partial pressure of oxygen (PO ). Unlike tion of nitrogen by slags is most often associated with
2
the Sieverts’ law for metals, in the case of slag, the the presence of carbon in the slag or in the atmosphere
following dependence is fulfilled [15]: above the slag. Pure experiments without the influ-
ence of carbon are not available. According to [16,
( G ) = K(G ) p1/2 VG /4
p ,
( 2 ) ( O2 )
G 17], the simultaneous increase in the concentrations
of carbon and nitrogen, as well as reaching the satura-
where VG is the degree of oxidation or valence of G
tion, indicate that nitrogen and carbon in the slag are
element in the slag.
combined into one compound, for example, cyanide
Since the valence of nitrogen in the slag VN = –3,
or cyanamide. Nitrogen at different content of CaO
then
dissolves in the form of free nitrogen N3‒ and binds
( n ) = K( n ) p1/2n p−3/4 . to Ca2+ (Ca3N2) or Al3+ (AlN) ions. The mechanism of
( 2 ) ( O2 )
nitrogen dissolution is very complex, because, as is
If the slag contains carbon, the following reaction known, it reacts and substitutes all three types of slag
occurs: oxygen: free (O2‒), final type (O‒) and bridging or one
C + ½{O2} = {CO}. which combines (O°).
At the same time, from 1 mole of oxygen, 2 moles When carbon and nitrogen are dissolved simul-
of CO are formed, which reduce the partial pressure taneously in molten slags [17] or there is carbon in
of nitrogen above the slag, which should be taken into the slag, they can be dissolved in the form of cyanide
account during calculations of the solubility. In real CN‒. In [17], analysis of literature data on the rela-
conditions, in the presence of nitride-forming ele- tionship between the solubility of nitrogen, cyanide,

Figure 1. Dependence of total nitrogen content in the slag (a, b) and CN‒ (c, d) on the partial pressure of nitrogen and CO above the
slag 50 % CaO‒50 % Al2O3 [17]: a, c — volume image; b, d — surface topography

45
V.O. Shapovalov et al.

and carbon in CaO‒Al2O3 slag with the partial pres- vection and again diffusion transfer of nitrogen in the
sures of nitrogen, CO and argon in the system was slag to the slag-metal interface; transition through the
performed, and model calculations of the solubility slag-metal interface; distribution of nitrogen over the
of nitrogen in various forms of existence in the slag volume of metal until equalizing the chemical poten-
were performed. Figure 1 shows the graphical depen- tial of nitrogen in the metal phase (dissolution).
dences built based on the results of this work for 50 % Depending on the speed of elementary links, one
CaO‒50 % Al2O3 slag. or the other link can determine the overall speed of
The main conclusions from these graphical depen- the process.
dencies are the following: The largest amount of information in the scientif-
● both an increase in the partial pressure of nitro- ic literature concerns the interaction of nitrogen with
gen and a decrease in the partial pressure of CO above melts based on iron, and at the second place — in-
the slag lead to an increase in the content of nitrogen teraction with slags. A little attention is paid to the
and cyanide CN– in it; processes of saturation of melts with nitrogen from
● the partial pressures of nitrogen and CO have a the gas phase through the slag. It was established [19]
stronger effect on the total nitrogen content than on that nitriding from the gas phase is possible when the
the cyanide content; slag is deoxidized by metallic calcium and alumini-
● the higher partial pressure of CO, the higher the um, but a clear dependence of the nitrogen content in
oxygen content and lower nitrogen content in the slag. Cr6WV steel on the content of oxygen in metal was
It is known that aluminium-oxygen anions of not detected by the authors (Figure 2). Moreover, it
(AlO2‒, AlO33‒, AlO45‒) slag as well as silicon-oxygen can be said that the oxygen content in steel does not
ones are able to associate with each other and form affect the content of nitrogen in it.
complex anions of large sizes [10]. During dissolution The content of nitrogen in Cr6WV steel, which is
of nitrogen, it can be embedded in the complex anions melted under CaO‒Al2O3‒15 % TiO2 slag, is 0.035 %,
instead of oxygen [18]. and under Al-295 slag, it is 0.026 %. These values
The solubility of nitrogen in the main slags two are lower than the equilibrium content (0.17 %) calcu-
to three times exceeds the solubility of nitrogen in lated for the conditions of interaction of liquid metal
iron. In acidic slags with an increase in temperature, with nitrogen [19].
nitrogen dissolves almost three times faster [16] than The amount of nitrogen dissolved in metal is di-
in iron, which may be associated with the formation
rectly proportional to the amount of slag nitrogen
of complex anions. Studying the dependence between
and inversely proportional to the distribution factor
nitrogen and carbon content in ANF-7 (80 % CaF2 + (N)
[ N ] . Taking into account the practical permanency
20 % CaO) flux showed that the concentration of ni- L=
trogen in the initial flux is much lower than the equi-
of the dependencies shown in Figure 2, in the con-
librium one and this slag not only cannot be a source
ditions of the experimentat, the coefficient of nitro-
for enrichment of metal with nitrogen, but can even
gen distribution between the slag and metal does not
facilitate the removal of this gas from it.
change.
NITROGEN IN THE GAS–SLAG–METAL SYSTEM Nitriding of metal under the slag occurs at a
The interaction in the nitrogen–slag–metal system lower rate, than in the case of contact of the metal
consists of the following links: convection and diffu- melt directly with the gas phase [20]. The highest
sion transfer of nitrogen to the slag surface in the gas rate of nitriding is fixed in arc melting conditions
phase; adsorption, dissociation, transition of nitrogen (2.6∙10‒5‒1.55∙10‒4 m/s), and the lowest is in melting
atoms through the gas–slag interface; diffusion, con- in a resistance furnace ((1.5‒3.0)∙10‒6 m/s).

Figure 2. Effect of oxygen content in Cr6WV steel, melted under the deoxidated slag, on the content of nitrogen in it: a — ANF-1P +
7 % Ca slag; b — ANF-1P + 7 % Al [19]

46
 Nitrogen absorption by 04Cr18Ni10 steel in plasma-arc melting under slag

Today, it is very difficult to find works that would

combine kinetic dependences of the slag and steel sat-
uration with nitrogen in a one process. Therefore, it
was the aim of this work.
RESEARCH PROCEDURE
As a slag composition, binary slag of the composition
of 50 % Al2O3 and 50 % CaO was selected. In the area
of this concentration, there are low-melting slags with
eutectic at a temperature of 1658‒1668 К. To prepare
slag, calcium oxide CaO powders and aluminium ox-
ide Al2O3 of fine grade (clean for analysis) were used.
The slags were preliminary melted in a graphite cru-
cible in an induction furnace with a protective atmo-
Figure 3. Content of nitrogen in 04Cr18Ni10 steel depending on
sphere. As a protective atmosphere, argon of the first the partial pressure of nitrogen above the metal (1) and the square
grade was used: 0.002 % O2; 0.01 % N2; 0.03 g/m3 of root from this pressure (2)
pair H2O. The carbon content in the slag was 0.29 %,
CaO — 49.3, Al2O3 — 50. [% N] = 0.0604545 + 0.0436667PN – 0.257576P2N .
2 2
As a metal specimen, 04Cr18Ni10 steel of the fol-
lowing composition was used, %: 0.04 C, 18.1 Cr, The coefficients of determination (R ) are equal to
2

10.65 Ni, 0.8 Si, 1 Mn. This steel does not contain ni- 0.845973 and 0.982467 for degrees 0, 1 and 2, respec-
tride-forming elements, that allows detecting nitrogen tively. Zero value means that the first coefficient is a
in the dissolved state. constant.
The determination of nitrogen content in the slag The dependence 2 in Figure 3 indicates that the
and steel specimens was carried out in the equipment content of nitrogen in the melt of 04Cr18Ni10 steel is
by the Kjeldahl procedure, which was updated in re- subjected to the Sieverts’ law. The Siverts constant for
lation to the determination of nitrogen in slags, and its the conditions of the experiment amounts to 0.277674,
detailed description is given in [21]. and the obtained dependence is the following:
The study of the kinetics of nitrogen absorption
[% N] = 0.27764 Pn2 . (1)
was carried out in the UPI installation [22], which al-
lows studying the absorption of gas from plasma by
the melt in the conditions, where its entire surface is The coefficient of determination (R2) is 0.979081,
covered with a plasma plume and active gas contacts and the mean square deviation is 0.00091. The prox-
with liquid pool. imity of the coefficient of determination to 1 indicates
After melting a weighted specimen of steel with that the model has a high significance.
the slag and its necessary holding in nitrogen-argon The theoretical calculations were carried out to
plasma, the plasmatron was switched off with simul- determine the solubility of nitrogen in the investigat-
taneous opening of the wedge mould for quenching ed steel. To do this, the Chipman‒Corrigan equation
the melt from liquid state and fixing the amount of was used [23]. The solubility of nitrogen in iron is
nitrogen in it. described by the equation

EXPERIMENTAL RESULTS 850 1 (2)

lgKN =
− − 0.905 + lg pN .
AND THEIR DISCUSSION T 2 2

The initial experiments were carried out with steel in

The interaction parameters [14, 23] of the first and
contact with the gas atmosphere in the absence of a
second order required for the calculation are given in
slag. The results are shown in Figure 3.
Table 2.
The main results of the statistical processing of the
The equilibrium constant of nitrogen dissolution in
obtained results are as follows. Polynomial regression
liquid iron at a temperature (T) amounts to:
(dependence 1) has the following appearance:
Table 2. Parameters of interaction of elements in iron at a temperature of 1873 K

Alloying element (j) Cr Ni Mn Si C Сr–Ni

1 order interaction parameter (e )

st j
N
–0.047 0.0063 –0.02 0.047 0.118 –

2nd order interaction parameter (rjN) 0.00032 0.00007 0.000032 – – –0.00008

47
V.O. Shapovalov et al.

Figure 4. Effect of temperature and pressure of nitrogen on its content in 04Cr18Ni10 steel (a) and surface topography (b)
f [% N ] After the data is substituted, we obtain the equa-
N (T )
KN =
pN 1/2
. (3) tion:
2
 1328 
Tacking of the logarithm we have:  T −1.4029 
[% n ]04Kh18n10 = pn 10 .
(7)
2
1
lg [ % n=
] 2 lg pn + lgKn − lgfn(T ) . (4)
2 During plasma melting, the distribution of tem-
The activity coefficients at a certain temperature peratures over the surface of the melt is very non-
can be determined by the known data at a temperature uniform. The region of the highest temperatures is
of 1873 K and by the equation concentrated inside the plasma plume and can reach
18000‒19000 K [24]. On the axis of the nozzle sec-
lgf
 3280
= 

− 0,75  ×
tion, depending on the electrical mode of melting,
N (T )
 T  pressure and composition of the gas atmosphere, the
(5)

( ) ( ) temperature usually exceeds 12000 K, and the surface

× ∑ e j
N (1873)
[ j ] + ∑ rNj(1873) [ j ]2 + ∑∑ rNj ,k [ j ][ k ] . ( ) of the metal pool is not lower than 2000 K in the areas
 j j j k 
adjacent to the anode spot.
Then after substitution of the equations (2) and (5) The obtained dependence (7) made it possible to
into the equation (4) we obtain: evaluate the melt temperature at PSR. The experimen-
tal dependence of the nitrogen content in 04Cr18Ni10
1 850
lg [ % n=
] lgp − − 0.905 − steel is described by the dependence (1), and the de-
2 n2 T
pendence (7) shows the calculated change in the solu-
( )
 3280  
−  − 0.75  ∑ e j [ j] + (6) bility of nitrogen in steel from the temperature.
 T   j n(1873)
It can be written that experimental and calculated

( ) + ∑∑ ( r values with the consideration of temperature are equal


+∑ r j
n (1873)
[ j ]2 n
j ,k
[ j ][ k ]) . to each other:
j j k 

[% n ] = [% n ]04Kh18n10 ,
 1328 
 T −1.4029 
0.277674 Pn = pn 10 ,
2 2

from which the temperature of the steel melt is about

2385 K.
The results of the calculation by the formula (7) in
graphical form are shown in Figure 4.
The obtained data showed that in the conditions
of melting without the use of plasma at small partial
pressures of nitrogen, of approximately up to 0.1 atm,
the temperature within 1823‒2323 K has practically
Figure 5. Kinetic dependence of nitrogen absorption at PSR no effect on the content of nitrogen in steel; at high-
(pN = 0.7 atm): 1 — for slag; 2 — for steel er partial pressures with an increase in temperature,
2

48
 Nitrogen absorption by 04Cr18Ni10 steel in plasma-arc melting under slag

Figure 6. Effect of nitrogen pressure on the nitrogen content in the slag and 04Cr18Ni10 steel in PSR: 1 — slag; 2 — metal
the nitrogen content in the melt of 04Cr18Ni10 steel =[% N ] 0.286549 pN + 0.0129405, (9)
decreases; the maximum nitrogen content that can be 2

achieved at a nitrogen pressure of 1 atm and a tem- the coefficient of determination is 0.964987, and the
perature of 2323 K is 0.145 %, and at a temperature of mean square deviation is 3.01468∙10‒4. The compari-
1823 K and the same pressure it is 0.21 %. son of the Sieverts’ constant of steel in PSR (1) with
The comparison of calculated and experimental data the constant in PAR (9) shows the proximity of their
shows that the obtained calculated data are lower than values. The difference amounts to 0.008875, which
the experimental, which may be associated with an in- corresponds to 3.2 % of the error.
crease in the energy of molecules in the plasma plume The coefficient of the nitrogen distribution be-
and the appearance of active nitrogen atoms and ions. It  (N) 
tween the slag and metal L = determined accord-
affects the intensity of gas absorption by the melt.  [ N ] 
An example of a kinetic dependence of simultane- ing to the obtained data at a partial nitrogen pressure
ous nitrogen absorption by a liquid 04Cr18Ni10 steel of not more than 1 atm changes slightly and is equal
and a slag at a partial nitrogen pressure of 0.7 atm in to 1.1‒1.2.
the furnace atmosphere is shown in Figure 5.
Conclusions
On all kinetic dependencies at pressures lower than
1 atm, the content of nitrogen in steel is lower than in 1. The kinetic dependences of nitrogen absorption
the slag and not higher than in the steel produced at a by a liquid 04Cr18Ni10 steel from nitrogen-argon
direct contact of liquid metal with plasma. plasma were obtained in the range of nitrogen partial
The derived dependencies of changing the nitrogen pressure from 0.05 to 1.0 atm. In all cases, the nitro-
content in the molten PSR steel on the partial pressure gen content in the slag exceeded its content in steel.
of nitrogen and the square root from this pressure are It was determined that in plasma-slag melting, the
shown in Figure 6. absorption of nitrogen by the slag and 04Cr18Ni10
In the conditions of plasma-slag melting, the content steel occurs in accordance with the Sieverts’ law. The
of nitrogen in steel grows with an increase in the nitro- obtained mathematical models of nitrogen dissolution
gen pressure in the atmosphere above the melt (Figure 6, in plasma melting have a high significance and their
a) and is subjected to the Sieverts’ law (Figure 6, b). coefficients of determination are close to 1. The value
The linear dependence of the nitrogen content in of the Sieverts’ constant in the conditions of both plas-
the slag is described by the equation ma-slag melting, as well as by the calculation way in
the equilibrium conditions for 04Cr18Ni10 steel was
=( % n ) 0.360025 pn + 0.00336233 (8) experimentally determined.
2
2. It was determined that at a partial nitrogen
with the coefficient of determination R2 = 0.995207 pressures of up to 0.1 atm and a temperature within
and the mean square deviation of 6.31644∙10‒5, which 1823‒2323 K, the content of nitrogen in steel is prac-
indicates a high significance of the model. The free tically independent of the temperature unlike higher
member of the last equation indicates the initial (re- partial pressures. With an increase in temperature, the
sidual) content of nitrogen in the slag. amount of equilibrium nitrogen dissolved in steel de-
The same can be said about the dependence of the creases.
nitrogen content in steel under the slag. The linear de- 3. It was established that in the conditions of the
pendence is as follows: experiments, i.e., when the plasma plume complete-

49
V.O. Shapovalov et al.

ly covers the molten specimen, the melt temperature 16. Grigorenko, G.M., Kozin, R.V. (2018) Nitrogen solubility
reaches 2385 K. In the conditions of PSR, the coef- in fluxes for electroslag technologies. Suchasna Elektromet-
al., 2, 37–40 [in Russian]. DOI: http://dx.doi.org/10.15407/
ficient of nitrogen distribution between the slag and sem2018.02.04
metal is 1.1‒1.2. 17. In-Ho Jung (2006) Thermodynamic modeling of gas solubili-
ty in molten slags (I)–carbon and nitrogen. ISIJ International,
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[in Russian]. (1993) The solubility of nitrogen in the CaO–CaF–Al2O3
2. Lakomsky, V.I., Lakomsky, V.V. (2012) Nitrogen in liquid system and its relationship with basicity. ISIJ International,
steels and slags. Kyiv, Naukova Dumka [in Russian]. 33(1), 48–52.
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steel. Moscow, Metallurgizdat [in Russian]. renko, G.M. (2006) Metal nitriding from gas phase in ESR.
4. Paton, B.E., Saenko, V.Ya., Pomarin, Yu.M. et al. (2004) Arc Sovrem. Elektrometall., 4, 3–5 [in Russian].
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ter. Sci., 39, 7269–7274. DOI: https://doi.org/10.1023/B:JM- phase nitrogen interaction with metal covered with molten
SC.0000048741.47509.b3 slag. Sovrem. Elektrometall., 3, 35–37 [in Russian].
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gen content on the structure and properties of tool steels. sokovska, I.A. (2019) Determination of nitrogen content in
J. of Achievements in Materials and Manufacturing Engi- СаО–Al2O3 systems by Kjeldahl method. Metrologiya ta Pry-
neering, 29(1), July. https://www.researchgate.net/publica- lady, 3, 55–60 [in Ukrainian].
tion/26872202 22. Zabarilo, O.S., Lakomsky, V.I. (1968) Carbon behavior in
6. Cieśla, M., Ducki, K.J. (2008) Influence of increased nitro- plasma-arc remelting of alloy 50N and Armco iron. Spec.
gen content on tool steels structure and selected properties. Elektrometallurgiya, 4, 78–85 [in Russian].
J. of Achievements in Materials and Manufacturing Engi- 23. Grigoryan, V.A., Stomakhin, A.Ya., Utochkin, Yu.I. et al.
neering, 27(2), July. https://www.researchgate.net/publica- (2007) Physical-chemical calculations of electric steelmaking
tion/26872252 processes: Coll. of Problems with Solutions. 2nd Ed. Moscow,
7. Saeed Nabil Ghali, Mamdouh Eissa, Hoda El-Faramawy MISiS [in Russian].
et al. (2013) Production and application of advanced high 24. Grigorenko, G.M., Pomarin, Yu.M. (1989) Hydrogen and ni-
nitrogen steel. In: Proc. of Int. Conf. on Science and Tech- trogen in metals during plasma melting. Kyiv, Naukova Dum-
nology of Ironmaking and Steelmaking (November 2013, ka [in Russian].
Jamshedpur, India). 1. https://www.researchgate.net/pub-
lication/262698868_Production_and_Application_of_Ad- ORCID
vanced_High_Nitrogen_steel V.O. Shapovalov: 0000-0003-1339-3088,
8. Grigorenko, G.M., Pomarin, Yu.M., Lakomsky, V.V. (2010) V.G. Mogylatenko: 0000-0002-6550-2058,
Properties of steels of X13 type, alloyed with nitrogen.
Sovrem. Elektrometal., 4, 26–29 [in Russian]. R.V. Lyutyi: 0000-0001-6655-6499,
9. Burnashev, V.P., Nikitenko, Yu.O., Yakusha, V.V. et al. (2020) R.V. Kozin: 0000-0002-8501-0827
Some aspects of melting high-nitrogen steel Kh21G17AN2
in plasma arc furnace. Suchasna Elektrometall., 2, 23–26 [in
Conflict of interest
Ukrainian]. DOI: https://doi.org/10.37434/sem2020.02.04 The Authors declare no conflict of interest
10. Shurkhal, V.Ya., Larin, V.K., Chernega, D.F. et al. (2000)
Physical chemistry of metallurgical systems and processes:
CORRESPONDING AUTHOR
Manual. Kyiv, Vyshcha Shkola [in Ukrainian]. V.G. Mogylatenko
11. Pomarin, Yu.M., Byalik, O.M., Grygorenko, G.M. (2007) E.O. Paton Electric Welding Institute of the NASU
Influence of gases on structure and properties of metals and 11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine.
alloys. Kyiv, NTUU KPI [in Ukrainian].
E-mail: vmogilatenko@gmail.com
12. Shapovalov, V.O., Biktagirov, F.K., Mogylatenko, V.G. (2023)
Out-of-furnace processing of steel: methods, processes, tech- Suggested Citation
nologies: Manual. Ed. by I.V. Krivtsun. Kyiv, Khimjest [in V.O. Shapovalov, V.G. Mogylatenko,
Ukrainian].
13. Enders, V.V. (2002) Nitrogen in steelmaking processes. Lityo R.V. Lyutyi, R.V. Kozin (2024) Nitrogen absorption
i Metallurgiya, 1, 95–100 [in Russian]. by 04Cr18Ni10 steel in plasma-arc melting under
14. Zhouhua Jiang, Huabing Li, Zhaoping Chen et al. (2005). The slag of CaO‒Al2O3 system. The Paton Welding J., 1,
nitrogen solubility in molten stainless steel. Steel Research 43–50.
Int., 76(10), 730–735. https://www.researchgate.net/publica-
tion/281642445 Journal Home Page
15. Grigoryan, V.A., Belyanchikov, L.N., Stomakhin, A.Ya. https://patonpublishinghouse.com/eng/journals/tpwj
(1987) Theoretical fundamentals of electric steelmaking pro-
cesses. Moscow, Metallurgiya [in Russian].

Received: 31.08.2023
received in revised form: 18.10.2023
accepted: 15.01.2024

50
ISSN 0957-798X THE PATON WELDING JOURNAL, ISSUE 01, January 2024

DOI: https://doi.org/10.37434/tpwj2024.01.08

Nondestructive method
of residual stress determination
in welded joints based on application
of high-density current pulses
and speckle-interferometry
L.M. Lobanov, V.V. Savitsky, O.P. Shutkevych, K.V. Shyian, I.V. Kyianets
E.O. Paton Electric Welding Institute of the NASU
11 Kazymyr Malevych Str., 03150, Kyiv, Ukraine
ABSTRACT
A procedure was developed for nondestructive evaluation of residual stresses in welded joints based on application of high-den-
sity current pulses and laser speckle-interferometry. Comparison of the results of residual stress measurement in welded joints,
obtained by the developed method and by hole-drilling method, was performed.

KEYWORDS: residual stresses, welded joints, high-density current pulse, speckle-interferometry, electroplastic effect,
hole-drilling method

INTRODUCTION relaxation methods. For this aim, it is proposed to

One of the important tasks in manufacturing, design- replace the process of drilling holes for stress relax-
ing and operation of welded structures is testing their ation, which violates the structural surface integrity,
stressed state, since residual stresses significantly af- with a nondestructive method of local stress relax-
fect the life of structures. To determine residual stress- ation based on the use of high density current pulses
es, such nondestructive methods as X-ray, ultrason- of 107‒1010 A/m2 (Figure 1) [14‒16]. It is believed that
ic, magnetic, etc., and mechanical ones based on the upon introduction of a current pulse, a zone with a
principle of stress relaxation are widespread [1‒3]. hemispherical appearance is formed, in which an elec-
During determination of residual stresses in elements troplastic effect arises and stress relaxation around the
and assemblies of structures for their elastic unloading, place of pulse introduction occurs. At the same time,
different methods are used, such as thermal effect [4], the integrity of the tested area of the material is not
plastic ball indentation [5], hole drilling, etc. The meth- violated. Since the size of the area, in which the stress
od of drilling small non-through holes (with a diame- relaxation occurs is unknown in advance and depends
ter of 1.0‒3.0 mm and a depth of 0.5‒3.0 mm) became on the electrode system parameters, it is necessary to
the most widespread for elastic unloading of residual evaluate the effectiveness of residual stress relaxation
stresses. Measurements of deformations and displace- upon the introduction of a current pulse.
ments around the zone of elastic unloading (around the PULSE CURRENT SOURCE
drilled hole) are performed using different methods of
experimental mechanics, as for example, electric strain A pulse current source (PCS) was created at the Insti-
gauging, mechanical strain gauges and optical meth- tute of Electrodynamics of NASU, and two types of
ods, including laser interferometry methods [6‒13]. shock-pulse and pulse effect electrode systems were
However, despite the widespread use of the hole drill- developed at the PWI of NASU, which are used to in-
ing method, it is still destructive. This imposes a num- troduce a pulse current into the studied area of the ob-
ber of restrictions on its application. For example, the
method has limited use in the diagnostics of real struc-
tures in the process of their operation. In this regard, the
determination of residual stresses is often performed on
individual mock-ups of elements and assemblies of real
structures. Therefore, the development and creation of
nondestructive methods for elastic unloading of residual
stresses of full-scale structures is a relevant task.
The aim of this study is to create a nondestructive
method for residual stress determination, which has
high accuracy and reliability inherent in destructive
Figure 1. Scheme of residual stress relaxation after introduction
Copyright © The Author(s) of a current pulse

51
L.M. Lobanov et al.

The created electrode system provides introduction

of a pulsed current into the studied material, the effect of
which leads to arising displacements around the point of
introduction. The values of displacements depend on the
stress state at the point of pulse introduction, as well as
on such parameters of the electrode system as charging
voltage on the capacitor storage U, inductance L, etc. In
order to determine the effective parameters of the elec-
trode system, it is necessary to assess the degree of stress
relaxation upon introduction of a current pulse into the
tested area of the studied element.
INFLUENCE OF THE ELECTRODE SYSTEM
PARAMETERS ON STRESS RELAXATION
To determine the influence of the electrode system
parameters on stress relaxation, a mechanical device
for loading test specimens was designed and manu-
Figure 2. Scheme of the electrode system with a shock-pulse factured (Figures 3, 4). A beam of equal bending re-
type of action: L — inductance coil; C — capacitor battery with sistance is used as a test specimen, as far as stresses
a charging voltage U; e — electrode; D — disc; O — object into in all cross-sections will not exceed the preset ones.
which a current pulse is introduced; d — dielectric gasket; PCS —
pulse current source
The speckle-interferometry device designed at
the PWI [11–13] allows registering displacements of
ject. The power source has wide capabilities for regu- surface points in the range of 0.03‒3 μm. Figure 5
lation of the basic electrical parameters of the system, presents patterns of interference fringes, that contain
which provide the required shape of current pulses. information about the displacements ux obtained after
The electrode system of a pulse action includes the introduction of a local current pulse into the test
electrode e, load P, which ensures the necessary con- specimen under loading (Figure 4). The interference
tact between the electrode and the point of current in- patterns show that with an increase in the values of
troduction and the pulse current source (Figure 2). In stresses σхх, the disturbance area around the point of
the electrode system of pulse action, the inductance L pulse introduction also grows.
of the coil varies in the range of 4700‒1900 μH, the A current pulse was introduced into the area of the
capacity C of the capacitor battery is 3400‒17000 μF, specimen being tested at the points ni with the level of
the charging voltage is 50‒186 V. stresses σхх from ‒100 to 100 MPa (Figure 4). As a re-
sult of the local stress relaxation in the vicinity of the
point of pulse introduction, displacements appeared,
which were registered by the noncontact method of
electron speckle-interferometry.
Figure 6 shows a diagram of dependence of the dis-
placements ux on the specified stresses σхх. This diagram
shows that the measured displacements ux at the points
A, located at a distance of 1.25 mm from the place of the
current pulse introduction (Figure 1), depend linearly on
the stress state at the place of measurements.
However, the use of data on the displacements
only at a point A is insufficient for calculating stress-
es σхх in real structures, since displacements will also
arise at these points due to the action of stresses σхх.

Figure 3. Device for loading test specimens: 1 — round plate, on

which a base plate is placed for positioning specimen on the de-
vice; 2 — polished plate, on which test specimen is located; 3 —
plate for placing speckle-interferometer; 4 — clamping element
for fixing test specimen; 5, 6 — units for loading test specimen;
7 — device for testing beam (test specimen) displacements; 8 — Figure 4. Scheme of specimen with equal bending resistance:
beam of equal bending resistance (test specimen) ni — points where a current pulse was introduced

52
 Nondestructive method of residual stress determination in welded joints based

Figure 6. Dependence of displacements ux0 measured by the meth-

od of electron speckle-interferometry at a distance of 1.25 mm
from the point of introduction of a high-density current pulse,
on the level of specified stresses σxx (value of the approximation
probability R2 = 0.99)

ur (r , θ) = Aσ xx + Bσ xx cos 2θ; (1)

uθ (r , θ) = C σ xx sin 2θ, (2)

where the coefficients A, B and C depend on the bound-

ary conditions, sizes of stress relaxation area, etc.
Considering that upon the introduction of a current
pulse, complete stress relaxation does not occur, and
also an initial effect (displacements arise after the intro-
duction of a current pulse in a material without stresses)
takes place, we introduce the following notations:
σim
xx = bxx + k σ xx ; (3)

σim
yy = byy + k yy σ yy ; (4)

τim
= k xy + τ xy ,
xy (5)

where bxx, byy і kxx, kyy, kxy are the constants that char-
acterize, respectively, the initial effect and the degree
Figure 5. Interference patterns obtained after the introduction of
a current pulse in the areas with a residual stress level. Electrode
of stress relaxation compared to the drilled hole with
system parameters: L = 3.26 μH, U = 150 V, s = 0.5 mm. The lines a diameter and depth of 1 mm.
indicate an increase in the zone of stress relaxation effect on the For a plane stressed state, using the principle of
displacements with an increase in the stress state stress superposition and axisymmetric problem, after
Therefore, it is necessary to introduce a new parame- transforming the equations (1)‒(2), we obtain:
ter, that depends linearly on residual stresses and dis-
ur (r , θ)= A(σim im
xx + σ yy ) +
placements and does not depend on the location of the
(6)
main coordinate axes and the type of stress state. + B (σim im im

xx − σ yy ) cos 2θ + 2τ xy sin 2θ  ;

NUMERICAL CALCULATIONS
We assume that upon the current pulse introduction, θ) C (σim
uθ ( r , = im im

xx − σ yy )sin 2θ − 2τ xy cos 2θ  .
(7)
stress relaxation occurs in the region with the axis of An important advantage of the electron speckle-in-
OZ symmetry, as when drilling a hole. In this case, the terferometry method in registering displacements is the
dependence of the displacements ur and uθ, arising as possibility of simultaneous determination of displace-
a result of unloading the stresses σхх, σyy and σхy on the ments in a large number of points. This makes it possible
angle θ at some distance from the center of the hole to obtain data on the displacements ux at points located
r in polar coordinates, is expressed by the following around a circle with a radius r with the center at the place
formulas [11–13, 17]: of the current pulse introduction (Figure 7).

53
L.M. Lobanov et al.

u x (r=
, θ) ur (r , θ) cos θ − uθ (r , θ)sin θ; (8)

u x (θ) =τ 1.25 mm= F (θ)σim im im

xx + G (θ)σ yy + H (θ) τ xy . (9)

Measuring their displacements ux at the points of

the circle (more than three points) allows calculating
the values σim im im
xx , σ yy and τ xy by the method of least
squares using a system of equations:

 F (θ1 ) G (θ1 ) H (θ1 )  σ xx 

im
u x (θ1 ) 
   im 
  
 F ( θ 2 ) G ( θ 2 ) H ( θ 2 )  σ yy  = u x (θ2 )  . (10)
 F (θ3 ) G (θ3 ) H (θ3 )  τim   
u x (θ3 ) 
 xy 
In order to evaluate the possibility of applying the
proposed algorithm and using σim im im
xx , σ yy and τ xy in
determination of residual stresses, let us build the di-
Figure 7. Scheme of displacement measurement in the vicinity of
agram of dependence of σim xx on the specified stress
the point (place) of introduction of a high-density current pulse
state σхх on the basis of the data on their displacements
As a result of preliminary studies, it was established uх presented in Figure 6. As in the case of the data on
that with the introduction of a current pulse when using the displacements uх (Figure 6), the dependence be-
different parameters of the electrode system, an elec- tween σim xx and σхх has a linear character (value of the
trode imprint with a diameter of 0.4‒1.0 mm appears approximation probability R2 = 0.99). This diagram
on the object. Therefore, the minimum radius of the is characterized by the angle of inclination relative to
circle is selected, in which the ratio “speckle pattern the OX axis, as well as by the value at the point with
quality ‒ displacement value” is optimal. It is known zero stresses. From the diagram, it is possible to de-
that when measuring displacements in the vicinity of a termine the values kxx and bxx, which correspond to the
drilled hole by the method of electron speckle interfer- set parameters of the electrode system, and calculate
ometry, it is optimal to use the distance r, which is equal the value of residual stresses at the point of the current
to 2.5 of the hole radius r0. Therefore, in the further pulse introduction by the formula (3).
studies, data on displacements at points located on a Thus, in order to determine residual stresses in
circle with a radius of 1.25 mm were used. full-scale objects, it is necessary to perform prelim-
To simplify the calculations from the equations inary calibration of the electrode system on the test
(6) and (7), let us separate the multiplier components specimen in order to obtain the dependence of σim xx

F(θ), G(θ) and H(θ) before σim im im on σхх (3). To do this, stresses σхх are specified on a
xx , σ yy and τ xy , which
depend on the angle θ and the coefficients A, B and C, beam of equal bending resistance and a current pulse
is introduced in the observation area of the speckle-in-
we obtain the following dependence:
terferometer. Based on the data obtained by the speck-
le-interferometry method on displacements of the sur-
face points after the introduction of a current pulse
into the metal being under the effect of mechanical
stresses, the value σim xx is calculated from the system

Figure 8. Scheme of welded specimen from AMg5 alloy: S1‒ Figure 9. Results of residual stress determination by the method
S6 — sections, in which residual stresses were determined; R — of electron speckle-interferometry based on the use of high-densi-
area, in which the weld reinforcement was removed ty current pulses for their relaxation

54
 Nondestructive method of residual stress determination in welded joints based

Figure 10. Distribution of stresses σxx passing through the weld reinforcement in the sections S1‒S3 (a) and after its removal in the
sections S4‒S6 (b)
of linear equations (10). Based on the data of certain therefore, to increase the accuracy of stress determi-
values σim xx at five points, a diagram of dependence
nation, it is suggested to use stress values averaged
of σimxx on σхх is constructed, according to which the over three points.
values kxx and bxx corresponding to the parameters of Stresses were determined by two methods in the
this electrode system are determined. sections passing both through the weld reinforcement
Therefore, using the value σim (S1, S2 and S3) as well as after its removal (S4, S5 and
xx calculated on the
basis of the data on the displacements registered by the S6) (Figure 8). The results of residual stress determi-
nation obtained based on the use of the pulse method
method of electron speckle-interferometry, and the val-
(sections S3 and S4) were compared with the data ob-
ues kxx and bxx determined from the diagram and by the
tained by the method of drilling holes with a diameter
formula (3), it is possible to determine the stress state at and depth of 1.0 mm, respectively (sections S1 and
the point of the current pulse introduction. S6) and 0.5 mm (sections S2 and S5) (Figure 10).
The use of functions σim im im
xx , σ yy and τ xy is better It is important that the results of stress determina-
compared to the data on the displacements uх at the tion in the area of the weld and the near-weld zone us-
points A and B (Figure 1), because in this case the ing the created nondestructive method are significantly
use of information about the displacements uх at the closer to the data obtained by the hole drilling method
points along the entire length of the circle reduces the (drilling holes with a diameter and depth of 0.5 mm),
relative error and increases the reliability of determi- and differ from the data obtained when using holes
nation of residual stresses. with a diameter and depth of 1.0 mm (Figure 10.) This
indicates that in the welded specimen there is a stress
EXPERIMENTAL STUDIES gradient over the thickness of the plates.
With the help of the developed technology, the residual The stress curves show the coincidence of the
stresses in butt-welded joints of AMg5 alloy were de- results of residual stress determination obtained by
termined based on the use of high-density current puls- the proposed approach and the hole drilling method,
es. The results of stress measurements obtained using which indicates that the developed nondestructive
current pulses were compared with the data obtained method for residual stress determination based on the
with the use of the equipment of the PWI based on the use of high-density current pulses for their relaxation
use of the hole drilling method in combination with allows obtaining reliable data on the distribution of
speckle-interferometry. The effectiveness and accuracy stresses along the selected section, at the same time
of residual stress measurement with this device is con- not violating the integrity of the studied material. The
firmed by the results of the Round Robin test, which maximum deviation of the results of stress determi-
was conducted by the International Institute of Welding nation by the nondestructive method compared to the
[18, 19]. Figure 8 shows a diagram of the welded spec- hole drilling method (drilled holes with a diameter of
imen with the specified cross-sections, along which 0.5 mm) does not exceed 20 MPa.
measurements were made.
Figure 9 shows the results of residual stress deter- Conclusions
mination using high-density current pulses along the A new nondestructive method for residual stress deter-
section S3. In the experiment, data on the displace- mination in welded joints has been developed, which
ments were registered after the introduction of current is based on the use of high-density current pulses for
pulses at the points located at the same distance from local stress relaxation and subsequent registration of
the center of the weld. The diagram shows that there displacements by the laser speckle-interferometry
is a scatter of the obtained experimental data and, method. It is shown that as a result of the introduc-

55
L.M. Lobanov et al.

tion of current pulses, a partial relaxation of residual tures using methods of electron shearography and speckle-in-
stresses occurs, and the value of measured displace- terferometry. The Paton Welding J., 8, 35–40.
12. Lobanov, L., Pivtorak, V., Savitsky, V., Tkachuk, G. (2013)
ments depends on the level of stresses at the place of
Technology and equipment for determination of residu-
pulse introduction. It was experimentally confirmed al stresses in welded structures based on the application of
that the developed method allows determining residu- electron speckle-interferometry. Mat. Sci. Forum, 768–769,
al stresses with an error of up to 20 MPa compared to 166–173. DOI: https://doi.org/10.4028/www.scientific. net/
the destructive method of drilling holes. MSF.768-769.166
The proposed approach allows determination of re- 13. Lobanov, L.M., Pivtorak, V.A., Savitsky, V.V., Tkachuk, G.I.
sidual stresses directly on real structures during their (2006) Procedure for determination of residual stresses in
welded joints and structural elements using electron speck-
operation, without violating the material integrity. This
le-interferometry. The Paton Welding J., 1, 24–29.
opens up wide prospects to use the method for monitor- 14. Lobanov, L.M., Pashchin, N.A., Loginov, V.P., Loginova
ing the stress-strain state of critical welded assemblies Yu.V. (2005) Application of electric pulse treatment of struc-
and structures in the power, aircraft and space rocket tural elements to extend their service life. The Paton Welding
industries, shipbuilding, etc. The use of the proposed J., 11, 19–23.
approach will increase the reliability and safety of the 15. Stepanov, G.V., Babutsky, A.I., Mameev, I.A. (2004) Nonstation-
operation of technical objects due to the timely detection ary stress-strain state in long rod produced by pulses of high-den-
sity electric current. Problemy Prochnosti, 4, 60–67 [in Russian].
and elimination of dangerous stresses. 16. Lobanov, L.M., Pashchin, N.A., Mikhodui, O.L. (2012) Effi-
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Express control of quality and stressed state of welded struc- Journal Home Page
https://patonpublishinghouse.com/eng/journals/tpwj
Received: 04.10.2023
received in revised form: 11.12.2023
accepted: 25.01.2024

56