Biuletyn Instytutu Spawalnictwa No. 2/2013

Transkrypt

BIULETYN
ISSN 2300-1674
INSTYTUTU SPAWALNICTWA
No. 2/2013
INSTITUTE OF WELDING BULLETIN
BIULETYN
INSTYTUTU SPAWALNICTWA
No. 2
BIMONTHLY
Volume 57
CONTENTS
•M. St. WĘGLOWSKI – Electrolytic etching in welding metallography.....................5
•J. Hilkes, V. Gross – Welding of CrMo steels for power generation and petrochemical applications - past, present & future............................................................11
•A. Kurc-Lisiecka – Forming of the texture, structure and mechanical properties
of cold-rolled AISI 304 steel........................................................................................... 23
•O. K. Makowieckaja – Technological innovations – a basis for the increase of
competitiveness of welding industry in the USA............................................................ 30
•A. Sawicki – Damping Factor Function in AC Electrical Arc Models. Part 1: Heat
Process Relaxation Phenomena, their Approximations and Measurement..................... 37
This work is licenced under
Creative Commons Attribution-NonCommercial 3.0 License
INSTITUTE OF WELDING
The International Institute of Welding
and The European Federation for Welding,
Joining and Cutting member
Summaries of the articles
M. St. Węglowski – Electrolytic etch- chanical properties of AISI 304 steel. The
ing in welding metallography
texture analysis was carried out basing oneIt has been discussed the process of electrolytic etching of metals and possibilities of
its application to welding technology. The
attention has been drawn to such process
parameters as: potential difference, current
density, electrolyte temperature, electrolyte stirring and polishing time as well as to
their influence on the metallographic specimens quality. The procedure of preparation
of metallographic specimens to the electrolytic etching process has been described.
It has been indicated also the necessity of
application of appropriate equipment which
will provide stability of etching process
parameters. Selected results of electrolytic
etching of welded joints and parent materials have been presented.
self on pole figures and three-dimensional
function of orientation pattern. Diffraction
phase analysis and magnetic examination
have shown the presence of α’ martensite
in the steel structure after deformation. The
volume fraction of martensite grew larger together with the increase of the degree
of deformation. The plastic strain induced
γ → α’ martensitic transformation in the
whole deformation range. It was observed
the development of both austenite and martensite texture. On the basis of mechanical
tests it has been found that together with
the increase of the plastic strain amount in
AISI 304 steel its mechanical properties and
hardness grow while its plastic properties
decrease.
J. Hilkes, V. Gross – Welding of
CrMo steels for power generation
and petrochemical applications - past,
present & future
O. K. Makowieckaja – Technological
innovations – a basis for the increase
of competitiveness of welding industry
in the USA
The paper provides an overview of the
development and applications of the classic
CrMo, the new CrMoV steels all the way up
to 12Cr1Mo. Also, the corresponding welding consumables for the power generation
and the petrochemical industry have been
discussed. Reference has been made to the
various international material standards that
are applicable and the specific properties
and requirements as set by today’s industries.
It has been presented the tasks and issues
in the field of materials joining. The model
of development and implementation to production of technological innovations proposed by the Edison Welding Institute in the
USA has been analysed.
A. Sawicki – Damping Factor Function in AC Electrical Arc Models.
Part 1: Heat Process Relaxation Phenomena, their Approximations and
Measurement
A. Kurc-Lisiecka – Forming of the texThe article is dedicated to technical evalture, structure and mechanical prop- uation of knowledge about arc damping
erties of cold-rolled AISI 304 steel
factor function. Special attention is paid to
It has been determined the influence of its specific value - the time constant, which
the plastic strain in the cold rolling process decides about the functioning quality of
on forming of the texture, structure and me- electrotechnological devices and electrical
No. 2/2013
BIULETYN INSTYTUTU SPAWALNICTWA
3
apparatus. Factors affecting heat process
relaxation phenomena in electrical arc have
been described. Approximation possibilities
of energetic process damping factor func-
tions in electrical arc have been examined.
Experimental methods of determining dynamic parameters of AC arc have been described.
Biuletyn Instytutu Spawalnictwa
ISSN 2300-1674
Publisher:
Instytut Spawalnictwa (The Institute of Welding)
Editor-in-chief: Prof. Jan Pilarczyk
Managing editor: Alojzy Kajzerek
Address:
ul. Bł. Czesława 16-18, 44-100 Gliwice, Poland
tel: +48 32 335 82 01(02); fax: +48 32 231 46 52
E-mail: [email protected];
[email protected];
[email protected]
www.bis.is.gliwice.pl
Prof. Jacek Senkara - Warsaw University of Technology,
Biuletyn Scientific Council:
Akademik Borys E. Paton - Institut Elektrosvarki im. E.O.
Patona, Kiev, Ukraine; Nacionalnaia Akademiia Nauk
Ukrainy (Chairman)
Prof. Luisa Countinho - European Federation for Welding,
Joining and Cutting, Lisbon, Portugal
Dr Mike J. Russel - The Welding Institute (TWI),
Cambridge, England
Prof. Andrzej Klimpel - Silesian University of Technology,
Welding Department, Gliwice, Poland
Prof. Jan Pilarczyk - Instytut Spawalnictwa, Gliwice, Poland
Biuletyn Program Council:
External members:
Prof. Andrzej Ambroziak - Wrocław University
of Technology,
Prof. Andrzej Gruszczyk - Silesian University of Technology,
Prof. Andrzej Kolasa - Warsaw University of Technology,
Prof. Jerzy Łabanowski - Gdańsk University of Technology,
Prof. Zbigniew Mirski - Wrocław University of Technology,
Prof. Jerzy Nowacki - The West Pomeranian University
of Technology,
Dr inż. Jan Plewniak - Częstochowa University
of Technology,
4
Prof. Edmund Tasak - AGH University of Science
and Technology,
International members:
Prof. Peter Bernasovsky - Výskumný ústav zváračský Priemyselný institút SR, Bratislava, Slovakia
Prof. Alan Cocks - University of Oxford, England
Dr Luca Costa - Istituto Italiano della Saldatura,
Genoa, Italy
Prof. Petar Darjanow - Technical University of Sofia,
Bulgaria
Prof. Dorin Dehelean - Romanian Welding Society,
Timisoara, Romania
Prof. Hongbiao Dong - University of Leicester, England
Dr Lars Johansson - Swedish Welding Commission,
Stockholm, Sweden
Prof. Steffen Keitel - Gesellschaft für Schweißtechnik
International mbH, Duisburg, Halle, Germany
Ing. Peter Klamo - Výskumný ústav zváračský Priemyselný institút SR, Bratislava, Slovakia
Prof. Slobodan Kralj - Faculty of Mechanical Engineering
and Naval Architecture, University of Zagreb, Croatia
Akademik Leonid M. Łobanow - Institut Elektrosvarki
im. E.O. Patona, Kiev, Ukraine;
Dr Cécile Mayer - International Institute of Welding,
Paris, France
Prof. Dr.-Ing. Hardy Mohrbacher - NiobelCon bvba,
Belgium
Prof. Ian Richardson - Delft University of Technology,
Netherlands
Mr Michel Rousseau - Institut de Soudure, Paris, France
Prof. dr Aleksander Zhelew - Schweisstechnische Lehrund Versuchsanstalt SLV-München Bulgarien GmbH, Sofia
Instytut Spawalnictwa members:
dr inż. Bogusław Czwórnóg;
dr hab. inż. Mirosław Łomozik prof. I.S.;
dr inż. Adam Pietras; dr inż. Piotr Sędek prof. I.S.;
dr hab. inż. Jacek Słania prof. I.S.;
dr hab. inż. Eugeniusz Turyk prof. I.S.
No. 2/2013
Research
Marek St. Węglowski
Electrolytic etching in welding metallography
Introduction
An increase in the types of applied metals and the related growth of various requirements for structural materials used
in welded structures or cast elements create a number of new metallographic issues
such as, for instance, testing alloy steels,
nickel alloys or titanium alloys. Proper
preparation of metallographic specimens of
such materials using classical etching methods poses numerous difficulties and sometimes proves impossible. In addition, when
it becomes necessary to test a great number
of specimens, e.g. in batch production, etching metallographic specimens is too time-consuming. Therefore, it is essential to implement new, more efficient methods. One
of the ways of allowing a significant reduction of the time needed for etching metallographic specimens without deteriorating their quality is electrolytic etching. If
materials have a complex chemical or, more
importantly, structural composition, it is of
great importance to develop proper etching
procedures making it possible to reveal only
selected microstructure components. And
also for this purpose, one can use electrolytic etching.
trochemical phenomena which cause the surface of a metal, i.e. the anode, to dissolve.
Cathodes play here the role of an element
which enables closing the current circuit by
the electrolyte and are responsible for proper
distribution of the current density. Cathodes
used in electrolytic etching do not wear even
after a longer period of time [1].
This method can be applied for most
metals and their alloys making it possible
to obtain a properly etched surface. Figure
1 presents the dependence between the density of current flowing through a specimen
and voltage. The process of etching takes
place within a voltage range strictly specified for given conditions. Applying higher
voltage is followed by a polishing process,
whilst applying too high a voltage triggers
excessive gas (oxygen) emission, causing
non-uniform removal of specimen material
and the formation of pits on the surface [2].
Principle of electrolytic etching
The process of electrolytic etching of metals is a complex electrochemical phenomenon. As an electric current flows through an
electrolyte, strong polarisation and proper
distribution of current density trigger elec-
Fig. 1. Dependence between the density of current flowing through specimen and voltage [2]
Marek St. Węglowski PhD Eng. - Instytut Spawalnictwa, Zakład Badań Spawalności i Konstrukcji
Spawanych / Testing of Materials Weldability and Welded Constructions Department
No. 2/2013
5
One of the most popular theories explaining the phenomenon of electrolytic etching/
electrolytic polishing is the theory developed by Jacquet [3], according to which
etching/polishing is caused by the anodal
action of electrolyte on, first of all, the
peaks of surface roughness of a metal subjected to etching/polishing and, to a lesser
degree, on cavities. Such diversified dissolving can be explained by the formation
of an anode film known also as a viscose
film. This film covers the metal surface,
with projecting surface roughness peaks being covered by a thinner film layer than the
cavities (Fig. 2). The resistance of the thinner film layer (on the peaks) is lower than
that of the thicker film layer (in the
cavities).
Fig. 2. The formation of anode film and distribution of
ions in the inter-electrode space [3]
For this reason the peaks of surface
roughness dissolve to a greater extent than
the cavities. In this manner, the surface irregularities are gradually made even until
polishing or etching. This theory clarifies
the basic course of electrolytic etching or
electrolytic polishing, yet it does not explain individual cases. The Jacquet theory
is right only for a certain group of phenomena; it properly explains the processes of
electrolytic etching/electrolytic polishing
of copper and steel in an electrolyte composed of phosphorus acid and organic additions as well as electrolytic etching/ electrolytic polishing of steel in a chloric-acetic
electrolyte. In these cases one can observe
the formation of a dense colloidal or liquid
film from the reaction products. Many theo6
ries have been developed in order to explain
this and other phenomena. Yet they all share
the same disadvantage, i.e. they explain
phenomena in a temporary manner and are
usable only in individual cases [3].
Factors
affecting
etching conditions
electrolytic
The most important factors affecting
etching conditions include the potential difference (PD) and the density difference of
current, electrolyte temperature, electrolyte
stirring, surface pre-treatment, heat treatment of the specimens, treatment time, dimensions of the electrolyser used for etching and electrolyte consumption degree.
Difference of current potentials and
density
Usually during electrolytic etching it is
possible to obtain better results if during the
process the difference of potentials is permanently monitored rather than the density of current. Independent of the difficulty
in maintaining constant current density, a
longer time of electrolysis combined with
heightened current density cause the emission of a gas which may trigger the formation of pits on the surface of a specimen.
Moreover, at the initial stage of anodal dissolving, prior to the stabilisation of etching
conditions, current density should significantly exceed the normal value for a period sufficient for the stabilisation of etching
conditions. It is important that at the initial
stage of etching the difference of potentials
on the electrolyser should be within the
boundary values used for etching.
Temperature of electrolyte
The resistance of the electrolyte decreases along with an increase in temperature.
Consequently, the voltage necessary for obtaining the same current density decreases.
The voltage corresponding to a given cur-
No. 2/2013
rent density is defined by an experimentally face of objects subjected and those unsubjected to heat treatment. The heat treatment
determined equation:
has a positive effect on the quality of the surK
face intended for etching, provided it has led
U=
a ×T + b
to the formation of a uniform structure. If a
heat treatment has triggered the formation of
where:
K, a, b – constants dependent on electrolytic carbides, as is the case with alloy steels durconductance, electrolyser dimensions and ing electrolytic etching for instance, the socalled point corrosion may occur as a result
current flowing through the electrolyser,
of the intense dissolving of areas around the
T – temperature.
The equation reveals that the power nec- carbides.
essary for maintaining given current density
decreases along with an increase in temper- Etching time
ature [1].
The time of electrolysis necessary for obtaining the desired condition of the surface
Electrolyte stirring
changes depending on a metal and electroDuring the process of electrolysis in stabi- lyte used. According to a general principle,
lised conditions, the products of the reaction the time of etching is inversely proportional
accumulate around the electrode. In some to current density. In this manner, in the case
cases the inflow of fresh electrolyte is insuf- of solutions containing orthophosphorous
ficient , and stirring is necessary to remove acid, for which low current density is used,
some of reaction products. However, one the treatment requires a longer time than in
should avoid excessive stirring as it could the case of the solutions of tetraoxochloric
destroy the film and prevent the stabilisation acid, for which high values of current densiof optimum etching conditions. Moving and ty are usually used. In general, etching lasts
stirring also prevents excessive local heat- between several seconds and several minutes
ing caused by the flow of current through [1].
the high-resistant layers on the anode and
favours maintaining a more uniform electro- D imensions of elec t rolyser for
lyte temperature. In many cases the best re- e tching
sults are achieved by rotating and swinging
An important factor associated with electhe anode rather than by stirring the solution. trolytic etching is the dimensions of an electrolyser, as their changes may significantly
Impact of heat treatment on the out- affect the process conditions.
come of electrolytic polishing
The structure of metal significantly affects
its properties and consequently its behaviour during treatment. This impact is visible
during mechanical working and in particular
during electrolytic etching. The change of the
structure and of the mutual relations of individual components causes changes both in
the potential and dissolving degree, affecting the surface appearance. The difference is
particularly evident while comparing the surNo. 2/2013
Impact of electrolyte consumption
During etching iron, for example, the content of Fe ions in the electrolyte increases
gradually as the amount of metal deposited on the cathode is much smaller than the
amount obtained from the dissolved anode.
Depending on the shape and dimensions of
the cathode, this amount makes up 5% - 15%
of the dissolved anode metal. An increase
in electrolyte density is accompanied by
7
a worsening quality of the surface subjected to etching. In consequence, the solution
becomes useless. Although a slight etching
effect can be observed, the phenomena accompanying the process render the microscopic observation of the etched specimen
difficult. The etching ability of the solution
can be slightly extended by using a slightly
higher current density or by adding between
5 ml and 10 ml of distilled water per one
litre of solution. A greater amount of water
also worsens the etching effect.
While using chemical or electrochemical
etching one should bear in mind that used
chemicals ought to be disposed of in accordance with environmental protection regulations.
Preparation of specimens for electrolytic etching
The quality of an electrolytically etched
surface is much more dependent on the
stage of mechanical polishing and grinding
than that obtained through chemical etching.
Electrolytic etching is a process which can
be initiated only at a certain specified surface
roughness. If the roughness is too high even
properly selected process conditions will
not result in a properly etched surface (the
process of electrolytic etching ”highlights”
scratches formed as a result of mechanical
polishing). This is probably because the film,
instead of filling the surface cavities, covers
both the peaks and the cavities with the layer of the same thickness. While considering
the applied gradation of abrasive paper, prior to electrolytic etching surfaces should be
precisely ground using abrasive paper, starting with the paper of the greatest granularity (e.g. 80 or 100), and next using paper of
lower granularity (280, 500, 800 etc.) Each
grade of abrasive paper should be used until
all marks (especially scratches) coming from
previously conducted grinding with the paper of a greater granularity are removed from
8
the surface of a metallographic specimen.
This can easily be observed if upon changing
the abrasive paper grade one also changes
the grinding direction by an angle of 10°20°. During grinding with abrasive papers it
is necessary to neither deform nor overheat a
specimen being ground. For this reason it is
advisable to wet grind specimens using water-resistant abrasive paper and devices ensuring permanent wetting of a grinding area
[4]. The process of precise grinding finishes
with abrasive paper of a granularity dependent on the type of a material out of which a
given specimen is made. It is assumed that
such granularity should amount to 600 for
steels and 1000 for non-ferrous metals.
The process of mechanical polishing is
carried out with a rotating disk covered with
a special cloth, onto which one applies diamond-based polishing materials of various
gradation (e.g. 6÷0.25 μm). Polishing usually
finishes with liquid slurry of aluminium oxide Al2O3 (grade e.g. 0.25 μm). The procedure of polishing is strictly dependent on the
type of a material being treated. The selection of optimum conditions requires preliminary tests. It may happen that due to improperly selected mechanical treatment, scratches
after electrolytic etching are not removed
but, on the contrary, become more visible. In
such a case, the scratches that are revealed
are those which during polishing were covered up by metal, which during electrolytic
etching are dissolved more intensively than
the remaining part of the specimen [3].
A better effect from electrolytic etching is
achieved sooner if the surface to be etched
is better prepared in the process of mechanical polishing. Objects intended for electrolytic etching should be carefully cleaned and
degreased. If impurities such as aluminium
oxide Al2O3 remain, the films floating on the
surface may disturb treatment by causing
the formation of stains on the surface being
polished.
No. 2/2013
A simple easy-to-use design guarantees
failure-free operation also for less experienced personnel. This is a unique selling
point and a competitive edge over devices
offered by other companies. A person carrying out tests can precisely adjust etching
time and voltage.
The device for electrolytic etching ElekExemplary results of electrolytic troMat ET1 can be used in the following inetching
stitutions:
Table 1 presents examples of electrolytic • metallographic laboratories of universities
and research institutes,
etching results (device ElektroMat ET1).
An indication of properly conducted degreasing is uniform water wetting of the
whole specimen surface after its rinsing. The
specimen should be subjected to electrolytic
etching immediately after the last rinsing as
it prevents accidental soiling and corroding
of the surface [3].
Table 1. Exemplary results of electrolytic etching
Alloy steel; X5CrNi18-10, electrolytic
etching, HCL + methanol, time 4 s, voltage 4 V
Austenitic weld,
electrolytic etching,
HCL + methanol,
time 4 s, voltage 4 V
Hasteloy C-2000,
5ml H2SO4, 95 ml H2O,
Hasteloy X,
5ml H2SO4, 95 ml H2O,
Steel 18-8,
HCL + methanol,
Transition zone of steel
grade 18-8, electrolytic
etching, HCL+ methanol, time 6 s, voltage 3 V
Steel 1.4404,
HCl + methanol,
Steel 1.4404,
HCl + methanol,
Device for electrolytic etching
In order to satisfy customers’ requirements, Instytut Spawalnictwa has developed a device for electrolytic etching ElektroMat ET1 (Fig. 3) [5]. The device is
of complex portable modern design ensuring the full repeatability of electrolytic
etching process parameters. In addition,
ElektroMat ET1 is resistant to operating
conditions usually present in the industry.
No. 2/2013
• industrial quality control laboratories,
• industrial DT laboratories.
The device can be used for most metals
and their alloys. It ensures obtaining a properly etched surface and allows electrolytic
etching of the following materials:
• iron alloys i.e. alloy and unalloyed steels,
• aluminium alloys,
• nickel alloys,
• titanium alloys,
• copper alloys.
9
Fig. 3. Device for electrolytic etching ElektroMat ET1
The ElektroMat ET1 is composed of an
adjustable laboratory power supply unit and
a system of electrodes with a vessel for conducting the electrolytic etching. The device
does not require a complicated installation
procedure, and its design features high-power transistors. The electronic control system
ensures operation with optimum efficiency
and output parameters. The output voltage
can be adjusted in an infinitely variable manner within a 0 V ÷ 30 V range. The device
operation (etching time) control is set up by
means of an electronic time relay, ensuring
the repeatability of an electrolytic etching
process. The device operation status is signalled by means of LEDs placed on the front
panel. In addition, the operator can read out
current and voltage values on a digital display.
Summary
Using highly-alloyed materials, developing technologies for welding structures
of critical importance, and manufacturing
products meeting more and more demanding customers’ needs has caused that the
recent years have seen an increasing role
of microscopic metallographic testing. For
this reason it is necessary to develop proper testing procedures for parent metals and
10
welded joints in order to eliminate errors
already at the stage of specimen preparation as such errors could adversely affect
the interpretation of obtained test results.
To this end, while carrying out microscopic metallographic tests, it is essential to
pay particular attention to the manner of
metallographic specimen preparation. It is
especially important to use a proper etching technique so that on the basis of microscopic metallographic observations one
can reveal the presence (or absence) of a
given microstructural component. In many
cases chemical etching is insufficient or so
problematic that the only solution is to use
electrolytic etching. Yet, also in this case,
the use of a proper metallographic reagent
is insufficient and should be supported by
a proper device for conducting electrolytic
etching.
References:
1. Tegard W. J.: Elektrolityczne i chemiczne
polerowanie metali. WNT, Warsaw, 1961
2. Cebula D., Widermann J.: Badania metalograficzne. Wyd. Biuro Gamma, Warsaw, 1999
3. Dobrowolski J.: Polerowanie elektrolityczne. Państwowe Wydawnictwa Techniczne, Warsaw, 1952
4. Łomozik M. et al.: Makroskopowe
i mikroskopowe badania metalograficzne materiałów konstrukcyjnych i ich
połączeń spajanych. Instytut Spawalnictwa, Gliwice, 2009
5. Węglowski M. St., Czylok K.: Wniosek
o udzielnie prawa ochronnego na wzór
użytkowy: „Obudowa urządzenia do
ujawniania mikrostruktury metali metodą
trawienia elektrolitycznego”. Instytut
Spawalnictwa, W.120687
No. 2/2013
Jan Hilkes, Volker Gross
Welding CrMo steels for power generation
and petrochemical applications - past, present and future
Introduction
Creep and high temperature resistant
CrMo steels have been around for a very long
time and have found use with great success
for applications in the petrochemical, and
respectively in the power generation industry. Typical products for these industries are
boilers, heaters, heat exchangers, reactors,
and hydrocrackers, usually built as heavy
wall pressure vessels.
In a continuous strive for optimizing the
economics in the various process installations in these industries, the service pressures and/or temperatures have increased.
This implied that the respective base materials either had to be made available in heavier thicknesses or they had to be developed
to meet higher strength and impact toughness requirements. Increased mechanical
properties will reduce or at least restrict the
necessary wall thickness which generates
an additional economical advantage in production, handling and installation of heavy
process equipment. An example of a heavy
all pressure vessel is a part of a Hydro
Conversion Unit as shown in Figure 1.
The basic and classic CrMo steels are
alloyed with 0,5%Mo – 1%Cr/0,5%Mo
– 2,25%Cr/1%Mo – 5%Cr/1%Mo –
9%Cr/1%Mo and 12%Cr/1%Mo. From
these steels further development has taken
place by adding elements such as V, W, Ni, Ti,
Nb, B and/or N to arrive at the new grades of
today such as the T/P22V, T/P23, T/P24, T/
P91, T/P92 and VM 12-SHC. Many of these
new grades have been applied successfully
in industry but the development continues.
Obviously, development of the welding
consumables had to and still must follow the
direction of the base materials with the assurance of meeting the same stringent requirements for the process equipment as the base
materials, even more so since the HAZ is
usually also considered part of the weld. Extensive research and development has taken
place at Böhler Schweisstechnik in Germany
to arrive at a full consumable range for the
new generation of CrMo(V) steels for which
also creep data up to 60 000 hours have been
collected. With increasing alloy level the specific welding procedures have to be adjusted
and will call for more precise and strict control of welding parameters and heat treatment.
Technical Details of a Hydro Conversion Unit
Base material: 2,25%Cr-1%Mo
Sizes:
thickness:
358 mm
length:
21 m
diameter:
5.3 m
total weight: 706 t
Fig. 1: Part of Hydro Conversion Unit by ATB, Italy
Service
conditions:
215.5 bar pressure and max. 454°C
Welding
consumables:
SAW: Union S1CrMo2/UV 420TTR
SMAW: Phoenix SH Chromo 2 KS
Jan Hilkes, Volker Gross - Böhler Schweisstechnik Deutschland GmbH, Hamm, Germany
No. 2/2013
11
Creep resistant CrMo steels
Basic metallurgy for base material
and weldmetal design
Creep resistant steels are steels that can
resist a certain stress at a specific service
temperature without exceeding a specified
amount of elongation. The maximum stress
to rupture at a specific temperature after a
specific time, e.g. 600°C and 105 h, is referred to as Creep Rupture Stress. For example, an engineering design criterion for a
power plant could require a minimal stress
of 100MPa for 105 h at service temperature.
The basic idea is that the vessel remains its
original sizes and shape while in service for
up to 20 to 30 years. Due to the fact that in the
processes used within the Power Generation
and Petrochemical Industry and the many
different service conditions such as pres-
sure, temperature and environment, a wide
variety of CrMo creep resistant steels with
additions of V, W, Ti, Nb, B and/or N have
been developed, while new types are also
still under development. Due to increased
pressures and temperatures, up to 370 bar
and 650°C, as for example in components
for Ultra-Super-Critical (USC) steam power generation plants, CrMo creep resistant
base materials with increased strength are
required to allow wall thicknesses that are
within the range of what fabricators can
handle in their facilities. Also for petrochemical applications (P22V), sizes now
up to 350mm are no longer an exception.
An overview of the international standards,
chemical compositions and maximum service temperatures of the actual and most
popular CrMo creep resistant steels is given
in Table 1.
Table 1: Overview of the international standards, chemical composition and maximum service temperature
of the actual and most popular CrMo creep resistant steels
CrMo type
INTERNATIONAL STANDARDS
ASTM & ASME
DIN/VdTÜV
EN
2.25Cr-1Mo
2.25Cr-1MoV
2.25Cr-MoVW
T/P 1
T/P 11
T/P 12
T/P 36
T/P 22
T/P 22V
T/P 23
16 Mo 3
10 CoMo 5-5
13 CrMo 4-5
15 CrMoV 5-10
15 NiCuMoNb 5 (WB 36)
10 CrMo 9-10
HCM 2S
8MoB 5-4
10 CrMo 5-5
13 CoMo 4-5
15 NiCuNb 5
10 CrMo 9-10
7CrWVMoNb 9-6
2.25Cr-1MoVTiB
T/P 24
7CrMoVTiB 10-10
7CrMoVTiB 10-10
5Cr-0.5Mo
9Cr-1Mo
9Cr-1Mo mod.
9Cr-0.5MoWV
T/P 502
T/P 9
T/P 91
T/P 911
12 CrMo 19-5
X12 CrMo 9-1
X10 CrMoVNb 9-1
X11 CrMoWVNb 9-1-1
X12 CrMo 9-1
X10 CrMoVNb 9-1
X11 CrMoWVNb 9-1-1
9Cr-0.5MoWV
T/P 92
X10 CrWMoNb 9-2
-
12Cr-0.25Mo
+1.4W1.3Co0.2V
-
X12 CrCoWVNb 11-2-2
(VM 12-SHC) t<10mm
-
12Cr-1MoNiV
-
X20 CrMoV 12-1
X20 CrMoV 11-1
0.5Mo
1.25Cr-0.5Mo
1,00Cr-0.5Mo
1.25Cr-1MoV
12
No. 2/2013
The creep resistance of a CrMo steel is
based on the formation of stable precipitations such as alloy carbides in a ferritic,
bainitic and/or martensitic microstructure
in the normalised condition. Due to a subsequent tempering treatment, a stable microstructure with precipitations is generated
that remains stable at the service temperature for which the steel has been developed. The precipitations formed will block
the grain-boundaries and prevent sliding of
the slip-planes to give the desired creep resistance properties. They should therefore
have the correct shape, be present in the
right amount and be evenly distributed to
obtain a homogeneous structure with homogeneous properties. Depending on the alloy level and the heat treatment(s), specific
types of precipitations will be formed in a
Table 2: Precipitations that can be found
in creep resistant CrMo steels /1, 2/
Precipitations and possible phases in CrMo steels
Graphite
Epsilon
= Fe2.4C
Cementite
= Fe3C
Chi
= Fe2C
M 2X
M 6C
M23C6
M 7C 3
Laves
M 5C 2
Z-phase
Mo2C
Cr3C
NbC
NbN
VN
specific amount. The governing parameters
for the heat-treatment are temperature and
time. The variety of precipitations that can
be expected and that are mainly used in the
design of classic and modern creep resistant
CrMo steels are listed in Table 2.
Table 1: Overview of the international standards, chemical composition and maximum service temperature
of the actual and most popular CrMo creep resistant steels (continued)
TYPICAL CHEMICAL COMPOSITION (wt%)
C%
Si %
Mn %
Cr %
Mo %
0,16
0,10
0,13
0,15
0,15
0,10
0,12
0,08
0,30
0,32
0,70
0,30
0,35
0,36
0,08
0,34
0,82
0,68
0,60
0,75
0,95
0,69
0,50
0,42
0,07
0,28
0,60
2,25
1,04
0,12
0,12
0,10
0,11
0,35
0,60
0,36
0,28
0,65
0,40
0,52
0,54
5,10
9,00
8,82
8,80
0,10
< 0,50
0,55
8,80
SERVICE
Ni %
V%
W%
Nb %
other %
Temp. °C
< 0,30 0,32 < 0,30
1,25
0,50
1,00
0,50
1,25
1,05
0,45
1,12
2,20
1,02
2,25
1,00
2,32 < 0,30
-
0,26
0,30
0,02
1,55
0,22
0,06
Cu: 0,62
N < 0,010
< 460
< 545
< 545
< 545
< 545
< 545
< 545
< 550
-
0,24
-
-
0,54
1,00
1,02
1,02
< 0,40
0,25
0,22
0,22
1,05
0,08
0,08
0,52
< 0,40
0,23
1,55
0,08
0,11
0,45
0,20
11,50
0,23
0,28
0,24
1,40
0,07
0,20
<0,50
<1,00
12,10
1,05
0,65
0,28
-
-
No. 2/2013
N < 0,010
B: 15-70 ppm
Ti: 0,05-0,10
N: 30-70 ppm
N: 0,05-0,09
N: 0,03-0,07
B: 0,001-0,006
Co: 1,30
N: 0,055
B: 0,003
-
< 550
< 550
< 585
< 585
< 625
< 625
< 650
<585
13
Heat treatments for CrMo steels and
welded joints
The heat treatments for the base materials are reasonably complex but are required
to obtain the optimal mechanical properties.
Depending on the alloy content a Normalising, Tempering and Annealing treatment at
various temperatures for several hours with
a controlled cooling rate have to be executed according strict procedures. The same is
valid for the weldmetal, with increasing alloy content the Post Weld Heat Treatment
(PWHT) for welded joints gets more complicated as illustrated in Figure 2.
When in subsequent PWHT, Intermediate
Stress Relieving (ISR) or in service, the ultimate heat treatment temperature of the base
material is exceeded too much and too long,
the precipitations can dissolve again which
causes reduction of the mechanical properties of the base material. This implies that,
for example, for this reason the maximum
temperature of 760°C for P91 in Figure 2
shall not be exceeded. For T/P23 in Figure
2, an Intermediate Stress Relieving is indicated for constructions with different material thicknesses. For each application the optimum PWHT shall be determined. Further
elaboration will follow in the welding chapter of this paper (Table 4).
Temper Embrittlement
When CrMo base material and the weld
metal is exposed to a temperature range of
400-500°C for a very long time there is a
risk of Temper Embrittlement. This type
of embrittlement is caused by trace elements as P, Sb, Sn and As that migrate to the
Fig. 2: Temperature cycle and heat control during
welding and PWHT of martensitic steel P91, E911 and
P92 (above) and ferritic/bainitic steel T/P23 (below).
For GTAW joints in <10mm wall thickness T/P 23,
no PWHT is required
To establish the sensitivity of a material to
temper embrittlement, a Step Cooling (STC)
heat treatment is carried out in the range of
593-316°C for a duration of 240 hours. The
difference in transition temperature (impact
properties) from before and after the heat
treatment is a measure for the sensitiveness
to temper embrittlement. A maximum allowable shift in transition temperature after step
cooling can be specified as a requirement for
base material and weldmetal. In order to reduce the risk of temper embrittlement, the responsible trace elements need to be restricted. Bruscato and Watanabe have developed
formulas to express the tendency of temper
embrittlement /3, 4/.
Watanabe:
J = (Mn + Si) x (P + Sn) x 104
Bruscato:
X = (10 P + 5 Sb + 4 Sn + As) / 100
elements in wt%
element in wt% and result in ppm
The formula of Watanabe is only valid for the
grain boundaries and can reduce the ductility in both base material and weldmetal. base material and is usually restricted to a value
To which extent this phenomena will occur of J < 160 but also requirements for J < 120 or
depends merely on temperature and time. 80 are being specified by the industry today.
14
No. 2/2013
The Bruscato formula, also referred to as
the X-factor, is valid for both weld metal and
base material. For weldmetal the specifications are becoming more and more stringent
with increasing wall thickness and desire
for additional assurance of the mechanical
properties. Initially, the required value of the
X-factor was X < 15, but present specifications already ask for X < 10. An additional
requirement for the Mn and Si content can be
set accordingly: Mn + Si < 1.1%.
Specifically for SAW where the trace elements can be picked up from both wire and
flux, the combination should be tested to
comply with the requirements. This means
one source for both wire and flux would be
recommended /5/.
Corrosion: Resistance to Oxidation,
Sulphidation and Hydrogen attack
In addition to the creep resistance and
resistance to embrittlement, CrMo steels
also show increased high temperature oxidation resistance with increasing alloy content. Comparing the scaling loss for plain
carbon steel and 1%Cr0.5%Mo with that
of 5%Cr0.5%Mo steel at 675°C, the scaling loss is reduced from >2.5 mm/y for the
first two to about 0.1 mm/y for the latter.
This makes these steels also very suitable
for gas-fired furnaces in the petrochemical
industry /6/.
Also sulphidation corrosion resistance increases with increasing alloy content. Comparing the corrosion rate of carbon steel with
that of 9%Cr1%Mo steel at 700°C, the corrosion rate is reduced from 1.0 to 0.2 mm/y.
Sulphur combines with Chromium to form
Chromium-Sulphides, and hence reduces the
amount of Cr-carbides required for creep resistance. Since most crude oils and other gaseous fuels contain either certain amounts of
Sulphur or H2S, sufficient sulphidation corrosion resistance is required for petrochemical installations.
No. 2/2013
Another important phenomena is High
Temperature Hydrogen Attack (HTHA), a
formation of Methane from Cementite (Fe3C
+ 2H2 → CH4 + 3Fe) in the base material under high Hydrogen pressures at high temperatures, as for example in heavy wall pressure
vessels for high-temperature, high hydrogen
services in oil refineries. The 2.25%Cr1Mo
and 3%Cr1Mo steels are typical base materials with good resistance to HTHA in this
application.
Welding and welding consumables
for CrMo steels
In general, creep resistant CrMo-steels are
welded with matching consumables in order
to have a homogeneous welded joint with
about equal mechanical properties. Matching compositions also have the same coefficient of thermal expansion, which prevents
or at least reduces the risk of thermal fatigue
in service. In this respect, the heat affected
zone (HAZ) is a vulnerable area.
In principle, all arc welding processes can
be applied as SMAW. GTAW, GMAW, SAW
and FCAW. For manual processes it is important to take sufficient measures to protect
the welders from heat, and then it is of utmost importance that the preheat as well as
the interpass temperatures are respected and
not reduced to accommodate the welders, as
well as while tacking. With the gas-shielded
processes it is vital to assure proper shielding of the weld. Due to the high preheat, the
gas-shield can be distorted and provide less
protection as required. Special nozzles and
gas cups are available to reduce the problem.
Over the last decades, Böhler Schweisstechnik Germany has developed a wide range
of welding consumables for welding CrMo
steels for the processes: SMAW, GTAW,
SAW, GMAW and FCAW. A selection table
for the respective welding consumables and
welding processes for creep resistant CrMo
steels can be found in listed in Table 3.
15
Table 3: Selection table for the respective welding consumables and welding processes for creep resistant CrMo steels
BASE MATERIAL
ASTM
CrMo type
&
EN
ASME
SMAW
WELDING CONSUMABLES FOR CrMo STEELS
SAW
GTAW
GMAW
wire
flux
Phoenix SH
Union I
Schwarz 3 K
Mo
1.25Cr10 CrMo
Phoenix
Union I
T/P 11
-0.5Mo
5-5
Chromo 1
CrMo
1.00Cr13 CoMo
Phoenix
Union I
T/P 12
-0.5Mo
4-5
Chromo 1
CrMo
1.25Cr15 CrMoV
Phoenix SH
-1MoV
5-10
Kupfer 3 K
15 NiCuNb
Phoenix SH
Union I
T/P 36
5 (WB 36) Schwarz 3 K Ni
Mo
20 MnMoPhoenix SH
Union I
Ni 5-5
Schwarz 3 K Ni
MoMn
2.25Cr10 CrMo
Phoenix SH
Union I
T/P 22
-1Mo
9-10
Chromo 2 KS
CrMo 910
2.25CrT/P
Phoenix SH
-1MoV
22V
Chromo 2 V
7CrMo2.25CrThermanit
Union I
T/P 23 WVMoNb
-MoVW
P23
P23
9-6
7CrMo2.25CrThermanit
Union I
T/P 24
VTiB
-1MoV
P24
P24
10-10
T/P
12CrMo
Phoenix
Union I
5Cr-0.5Mo
502
19-5
Chromo 5
CrMo 5
X12 CrMo
Thermanit
Thermanit
9Cr-1Mo T/P 9
9-1
Chromo 9 V
MTS 3
Thermanit
9Cr-1Mo
X10 CrMo- Chromo 9 V;
Thermanit
T/P 91
mod.
VNb 9-1
Thermanit
MTS 3
Chromo T91
9CrX11 CrT/P
Thermanit
Thermanit
-0.5MoMoWVNb
911
MTS 911
MTS 911
WV
9-1-1
9CrX10
Thermanit
Thermanit
-0.5Mo- T/P 92 CrWMoNb
MTS 616
MTS 616
WV
9-2
X12 Cr12CrCoWVNb
-0.25Mo
11-2-2
Thermanit
Thermanit
+1.4W1.
(VM12MTS 5 CoT
MTS 5 CoT
3Co0.2V
-SHC)
t<10mm
12Cr-1MoX20 CrThermanit
Thermanit
NiV
MoV 11-1
MTS 4
MTS 4 Si
0.5Mo
T/P 1
8MoB 5-4
Depending on the alloy level, from only
0.5%Cr to 12%Cr-1%Mo the welding condition regarding preheat (Tp) and interpass
(Ti) temperature as well as the subsequent
temperature cycles during SR, ISR, STC
16
FCAW
Union I
Mo
Union I
CrMo
Union I
CrMo
Union S
2 Mo
Union S
2 CrMo
Union S
2 CrMo
UV 420
TT
UV 420
TT
UV 420
TT
Union TG
Mo R
Union TG
CrMo R
Union TG
CrMo R
-
-
-
-
Union I
Mo
Union I
MoMn
Union I
CrMo 910
Union S
3 NiMo 1
Union S
3 NiMo 1
Union S
1 CrMo 2
Union S
1 CrMo 2V
UV 420
TT(R)
UV 420
TT(R)
UV 420
TTR
UV 430
TTR-W
Union TG
Mo R
Union TG
CrMo 9 10 R
Union I
P23
Union S
P23
UV 430
TTR-W
→UV P23
Union I
P24
Union S
P24
UV 430
TTR-W
→UV P24
-
-
-
Union I
Union S1 Marathon
CrMo 5
CrMo 5
543
Thermanit Thermanit Marathon
MTS 3
MTS 3
543
Thermanit
MTS 3 PW
MTS 3
MTS 3
543
Thermanit
MTS 3 PW
MTS 911 MTS 911
543
-
MTS 616 MTS 616
543
-
-
-
-
MTS 4 Si
MTS 4
543
-
-
-
and PWHT´s change drastically. An overview with typical guidelines in this regard
for is provided in Table 4. Also see Figure
2 above for examples of complicated heat
treatments. The required heat treatment
No. 2/2013
Table 4: Overview of typical guidelines for Preheat & Interpass temperatures and PWHT as SR, ISR and STC
for CrMo steels. Also see Figure 2.
STANDARDS
CrMo type
0.5Mo
1.25Cr-0.5Mo
1,00Cr-0.5Mo
1.25Cr-1MoV
ASTM &
ASME
EN
Tp °C
Ti °C
T/P 1
8MoB 5-4
RT
RT
T/P 11
10 CrMo 5-5
T/P 12
13 CrMo 4-5
15 CrMoV 5-10
T/P 36
15 NiCuNb 5
(WB 36)
21 MnMoNi 5-5
2.25Cr-1Mo
2.25Cr-1MoV
2.25Cr-MoVW
2.25Cr-1MoV
T/P 22
10 CrMo 9-10
T/P 22V
T/P 23
7CrWVMoNb 9-6
T/P 24
7CrMoVTiB 1010
5Cr-0.5Mo
T/P 502
12 CrMo 19-5
9Cr-1Mo
T/P 9
X12 CrMo 9-1
9Cr-1Mo
mod.
9Cr-0.5MoWV
9Cr-0.5MoWV
12Cr-0.25Mo
+1.4W1.
3Co0.2V
12Cr-1MoNiV
PREHEAT & INTERPASS TEMPERATURE, PWHT as SR, ISR
and STC GUIDELINES for CrMo STEELS
T/P 91
T/P 911
T/P 92
X10 CrMoVNb
9-1
X11 CrMoWVNb
9-1-1
X10 CrWMoNb
9-2
X12 CrCoWVNb
11-2-2
(VM 12-SHC)
t<10mm
X20 CrMoV 12-1
200250°C
200250°C
200250°C
200250°C
200250°C
200300°C
200300°C
200300°C
200280°C
225300°C
200300°C
200300°C
200300°C
200300°C
200280°C
200280°C
> 200°C
> 200°C
> 200°C
> 200°C
> 200°C
200300°C
200250°C
200300°C
200280°C
SR h, °C
2-4h @
580-630°C
2-4h @
660-700°C
2-4h @
660-700°C
2-4h @
660-700°C
2-4h @
580-620°C
2-4h @
580-620°C
2-4h @
670-720°C
No. 2/2013
60h @ 550°C +
40h @ 620°C
8h @ 705°C + STC
+32h @ 705°C
0.5-4h @ 740°C**
0.5-4h @ 740°C**
200300°C
200300°C
200300°C
200300°C
200280°C
200280°C
slow cool
after welding
* with great differences in wall thickness
** no PWHT required for GTAW up to wall thickness of 10mm
depends also on the thickness of the construction and has to be determined by the
fabricator as part of the welding procedure
development. The main factor is to have a
controlled, slow and even heating up and
PWHT/STC h, °C
STC depending on
application
1h @
680°C
1h @ 540560°C*
2-4h @
730-760°C
slow cool
after welding
slow cool
after welding
slow cool
after welding
slow cool
after welding
slow cool
after welding
> 225°C
ISR h, °C
xh@ 750°C
xh @ 730-780°C
xh @ 730-780°C
xh @ 770°C
xh @ 760°C
x depends
on thickness
cooling down to prevent additional stresses
in the welded joint. For heavy thicknesses
this means heating up from as many sides
as possible to get the required heat distribution in the material. These precautions have
17
Figure 3 shows a very heavy wall examto be taken to safeguard the base material,
the weldmetal and the heat affected zone ple of a pipe connection of a live-steam pipe
(HAZ).Recent developments in P22V, P23, of P91 base material in a Power Station.
P24, P92 and VM 12-SHC have
governed more detailed and precise welding and production procedures to retain control over the
outcome of the final product. Although these materials are not as
forgiving as the basic CrMo steel,
the weldability is excellent when
the correct procedures are followed. Depending on the application, there can be requirements
for STC and Bruscato´s X- factor.
Fig. 3: Weld preparation and final weld in a pipe connection
of a live-steam pipe of P91, welded with SMAW
For very heavy wall-thickness in
using Thermanit Chromo 9 V
P22V it could be necessary to apAs indicated the heat treatments includply intermediate stress relieving treatments
as to reduce the overall stress level before ing preheat and interpass temperature have
the final heat treatments applied. With the to be under strict control to successfully
experience that Böhler Schweisstechnik complete these types of welded joints. The
Germany has built up over the last decades, temperature ranges for the preheat and inthe support that can be provided to the cus- terpass temperatures given in Table 4 are
tomers has become a vital link in the supply to be respected throughout completion of
the joint. For this application, SMAW is
chain in today’s business.
As already mentioned, the thicknesses very suitable due to its flexibility and low
for a welded part in the power generation investments regarding equipment. In orand petrochemical industry keeps increas- der to increase efficiency, higher weldmeting and higher tensile strength materials, al deposition per unit of time, developwith more stringent mechanical properties ment is ongoing for FCAW consumables
and chemical composition, are used to keep for CrMo steels. As listed in Table 3, a
fabrication feasible. This means that the number is already available but the range
welding consumables have to be adapted to will be extended upon the demand of the
industry.
follow this trend.
Technical Details of the Reactor:
Base
2,25%Cr-1%Mo
material:
Fig. 4: Heavy wall Reactor in 2.25%Cr-1%Mo
by GODREJ, INDIA
18
thickness:
124, 132 and 153 mm
total weight:
about 500 t
Service
conditions:
120 bar pressure and 437°C
SAW: Union S1CrMo2/UV 420TTR
Welding
consumables: SMAW: Phoenix SH Chromo 2 KS
No. 2/2013
Applicable manufacturing parameters,
which include the welding parameters as well
as the quality of the welding equipment and
the skill-level of the welders, become more
important with an increasing initial strength.
The “operating window” will become smaller. Therefore suitable control mechanisms
and procedures have to be set up to ensure
the proper application of the required parameters. In particular the control of the following items shall not be neglected for achieving successful welds:
• Selection of the suitable SAW wire & flux
combination
• Proper rebaking of fluxes and electrodes
• Verification of preheating & interpass
temperatures
• Setting of the electrical welding parameters
• Weld build-up and beadsequence
• Verification of the heat treatment temperParameter control and suggestions for
ature.
“Best Practice”
Almost all issues encountered in CrMo
CrMo(V) weld metal typically shows a welds could be related to the non-observance
bainitic/martensitic micro structure that respond very sensitively to any kind of heat
put in by means of welding and heat treatment. Furthermore, the high strength in the
as welded condition requires accurate handling in terms of Hydrogen and ISR in order
to avoid cracking due to Hydrogen and/or the
restrained condition of welds in heavy wall
nozzles for example.
Fig. 5. Ferrite precipitations in P11 SA welds
To elaborate on some of the influences,
typical observations in welds made in CrMo(V) creep resistant steels are illustrated in
the next paragraph.
Figure 5 shows ferrite precipitations in
P11 due to excessive PWHT temperature.
The micrograph in figure 6 shows Hydrogen
damage due to a improperly applied soaking
treatment, leaving too much residual Hydrogen in the weldmetal. Figure 7 shows the effect of bead-thickness in SMA welds, a shift
of the impact properties to higher temperatures, due to a much courser grain-structure. Fig. 6. Crack surface due to Hydrogen in P22V SA welds
The SAW consumables range covers all
the CrMo steels available today. GTAW is
mainly used for root welding or automated welding in demanding industries. The
GMAW range is available but not popular in
the Power Generation industry.
Another practical example is that of a Reactor build in 2.25%Cr-1%Mo steel. Figure 4
shows one of a number of these types of heavy
wall pressure vessels produced by Godrej in
India. They have built up excellent and practical experience to be able to build such units.
When dealing with heavy wall thicknesses,
modern CrMo creep resistant steels and very
stringent specifications, it is absolutely necessary to build up sufficient experience to
be able to satisfy the demanding engineering
companies as well as the Oil and Power companies, who are the ultimate client.
No. 2/2013
19
Table 5: Overview of typical applications of CrMo steels in the Power Generation & Petrochemical Industry
CrMo type
BASE MATERIAL
ASTM &
EN
ASME
INDUSTRIAL APPLICATIONS
POWER GENERATION
PETROCHEMICAL
Pressure vessels
T/P1
8MoB5-4
Pressure vessels; Rp0.2 >
290 MPa, Rm > 500 MPa
T/P11
10CrMo5-5
Steam headers
T/P12
13CoMo4-5
1.25Cr-1MoV
-
15CrMoV5-10
-
T/P36
-
-
15NiCuNb5
(WB 36)
20MnMoNi5-5
2.25Cr-1Mo
T/P22
10CrMo9-10
0.5Mo
1.25Cr-0.5Mo
1.00Cr-0.5Mo
2.25Cr-1MoV
2.25Cr-MoVW
2.25Cr-1MoV
T/P22V
T/P23
T/P24
Water walls;
parts of evaporater
Main steam pipe; reheater
steam pipe; Rp0.2 > 440
MPa, Rm 590-780 MPa
High pressure steam drums
< 545
Reactor vessels (nuclear)
Parts of superheaters;
Rp0.2 > 310 MPa,
Rm 515-690 MPa
-
Parts of superheater;
membrane walls
Parts of superheater;
membrane walls
5Cr-0.5Mo
T/P502
12CrMo19-5
-
9Cr-1Mo
T/P9
X12CrMo9-1
-
9Cr-0.5MoWV
T/P911
X11CrMoWVNb9-1-1
9Cr-0.5MoWV
T/P92
X10CrWMo
Nb9-2
-
X12CrCoWVNb11-2-2
(VM12-SHC)
t<10mm
12Cr-0.25Mo
+1.4W
1.3Co0.2V
12Cr-1MoNiV
20
-
X20
CrMoV11-1
< 545
Feed water pipe
7CrMoWVMoNb9-6
7CrMoVTiB10-10
X10CrMoVNb9-1
< 535
< 545
-
T/P91
< 460
Heat exchangers
Rp0.2 > 415 MPa,
Rm 585-760 MPa -->
9Cr-1Mo
mod.
Heavy Wall Pressure
Vessels, Coke Drums,
Hydrofiner Reactors, Catalytic
Reformer Reactors
Service
max. T
in °C
Steam headers,
superheaters for ultra
super critical boilers;
Rp0.2 > 450 MPa,
Rm 630-790 MPa
Steam headers,
superheaters
Steam headers,
superheaters for Ultra
Super Critical boilers
Reactors, coke drums,
furnaces, piping
< 535
Hydrocrackers, Heavy Wall
Pressure Vessels
for Hydrogen Service
< 482
-
< 550
-
< 550
Pressure vessels in high temperature sulfur corrosion, resistance reactor furnaces and reactors
Reactors, High Temperature
Sulphur corrosion resistance,
furnaces and piping
< 585
High pressure steam headers
& piping
< 585
-
< 625
-
< 625
-
< 650
Tubing in H2S environments
< 585
Superheater tubes with
thickness < 10mm
Steam headers,
superheaters;
Rp0.2 > 500 MPa,
Rm 700-850 MPa
High Pressure &
High Temperature
< 550
High Pressure, High
Temperature & Corrosion
No. 2/2013
Fig. 7. Influence of weld build-up on impact toughness
Fig. 8: QA to be included to verify required parameters
of the above mentioned items. Consequently suitable control mechanisms have to be
developed to ensure proper welds. Quality
assurance becomes a major factor and must
be included in the CrMo welding fabrication.
QA has to be considered as an essential variable, as illustrated with Figure 8.
In conclusion we can state that CrMo
creep resistant steels are widely and successfully applied in the Power Generation and
Petrochemical Industries. The development
towards higher service temperatures ask for
new materials, both for base material as for
welding consumables.
To illustrate typical examples of where
the various CrMo materials are applied,
an overview of typical applications of CrMo
steels in the Power Generation and the Petrochemical Industry is given in Table 5.
With this paper we intended to provide
an overview of the available materials, the
standards, the consequences and the implications with regard to welding, heat treatments
and fabrication. When the correct procedures are developed up front and adhered
to throughout the production, projects can be
and have been successfully completed.
2. Bhadeshia H.K.D.H: Design of Creep-Resistant Steels. Proceedings of Ultra-Steel
2000. National Research Institute for
Metals, Tsukuba, Japan 2000, pp. 89-108
3. Bruscato R.: Temper Embrittlement and
Creep Embrittlement of 2.25%Cr - 1%Mo
shielded metal arc weld deposits. Welding Journal 49 (4), 1973, pp. 148-156
4. Watanabe J. et. al.: Temper Embrittlement of 2.25%Cr - 1%Mo Pressure Vessel
Steel. ASME 29th Petroleum Mechanical Engineering Conference, Dallas,
USA, 1974
5. Gross V., Heuser H., Jochum C.:
Schweisstechnische Herausforderung bei
der Verarbeitung von CrMo(V)-Stählen
für Hydrocracker. Publication of Böhler Thyssen Schweisstechnik, Germany,
2007
6. Handel Geert van den: Chroom-Molybdeen staalsoorten. Lastechniek, Nederlands
Instituut voor Lastechniek (NIL), No. 5,
May 2008, pp.10-14
7. Gross V.: Improved toughness in
2.25%Cr - 1%Mo(V) weldmetals for joining heavy walled reactors. Publication
of Böhler Thyssen Schweisstechnik,
Germany, 2006.
8. Fuchs R., Gross V., Heuser H., Jochum
C.: Properties of matching filler metals
for the advanced martensitic steels P911,
P92 and VM12. Proceedings of 5th International EPRI RRAC Conference, Alabama, USA, June 26-28, 2002
References
1. Cole D., Bhadeshia H.K.D.H.: Design
of Creep-Resistant Steel Welds. Research work. University of Cambridge,
Department of Materials Science and
Metallurgy, 1998
No. 2/2013
21
neración de Energía y Petroquímica“9. Heuser,H, Jochum C.: Neue Schweiss- Pasado, Presente & Futuro. CESOL
zusatzwerkstoffe für neue KraftwerksConf. Proc. 1er Congreso Internacional de
stähle. Publication of Böhler Thyssen
Schweisstechnik, Germany, 2004
Soldadura y Technologías de Unión (17as
10.Gross V, Heuser H., Jochum C.: Neuartige
Journadas Téchnicas), Madrid, Spain,
Schweisszusätze für bainitische und mar7-9 October 2008., pp119-124.
tensitische. Publication of Böhler Thys- 13.Hilkes J., Gross V.: Het lassen van CrMo
sen Schweisstechnik, Germany, 2005
stalen voor de Energieopwekking en de
11.Valaurec, Mannesmann: Seamless boiler
Petrochemische Industrie - Verleden,
tubes and pipes. Publication of Valaurec
Heden en Toekomst. Dutsch & Belgium
& Mannesmann Tubes, V&M 507-7e
Welding Institute, NIL/BIL Lassymposi12.Hilkes J., Gross V.: Soldadura de los aceum, Eindhoven, The Netherlands, 26/26
ros CrMo para aplicaciones en la GeNovember 2008
22
No. 2/2013
Agnieszka Kurc-Lisiecka
Forming of the texture, structure and mechanical properties
of cold-rolled AISI 304 steel
Introduction
Stainless austenitic steels are commonly
used for their possible unique mechanical
and plastic properties combined with corrosion resistance. From a structural point of
view these steels can be divided into steels
with a stable austenite structure, steels with
an unstable austenite structure and steels
with an austenitic-ferritic structure [1-4].
Steels 18-8, i.e. those with metastable austenite, may undergo transformations induced
both by plastic strain and by quenching. Depending on the steel’s chemical composition,
stacking fault energy, the size and shape of
grains as well as plastic working conditions
(degree, rate and temperature of strain),
phase change in such steels may proceed as
follows: γ → ε, γ → ε → α’ or γ → α’ [5-6].
The obtained volume fraction of individual
phases affects the mechanical properties and
corrosion resistance of these steels [7]. During plastic strain, metastable austenitic steels
undergo the development of an austenite texture as well as the development of a martensite texture, formed as a result of the transformation [8]. The texture plays a significant
role in the process of product formation and
in finished products. The texture obtained in
the post-transformation material is closely
connected with the texture of the material at
the initial state. Metals and their alloys with
a face-centred cubic lattice (A1) after plastic strain may have one of the two types of
texture deformation, namely, a copper-type
texture (high value of stacking fault energy)
or an alloy-type texture (low value of stack-
ing fault energy). The crystallographic dependences between the austenite texture (γ)
and the martensite texture (α’) are described
by a range of relationships such as the Bain
relationship, the Kurdjumov-Sachs (K-S) relationship and the Nishiyama-Wassermann
(N-W) relationship [9]. The purpose of this
research work was to determine the impact
of cold plastic strain during rolling, on shaping the texture, structure and mechanical
properties of steel AISI 304.
Materials used and research methodology
The research involved austenitic steel
AISI 304 with the chemical composition presented in Table 1. The starting material in the
form of a sheet (2 mm × 40 mm × 700 mm)
underwent hyperquenching at 1100°C for
1 hour and cold rolling until it reached 70%
strain. The rolling was carried out at room
temperature, maintaining the same direction
and side of the band being rolled.
Table 1: The chemical composition of steel AISI 304
[% by weight]
C
Cr
Ni
Mn
Si
0,033
18,08
9,03
1,32
0,41
Mo
P
S
N
Fe
0,23
0,026
0,002
0,026
70,84
Based on empirical formulas [10] the following parameters were calculated for the
austenitic steel tested – stacking fault energy (SFE), the temperature at the beginning
of a martensite transformation (Ms) and
mgr inż. (MSc Eng.) Agnieszka Kurc-Lisiecka - Instytut Materiałów Inżynierskich i Biomedycznych, Politechnika Śląska, Gliwice /Institute of Engineering Materials and Biomaterials, Silesian
University of Technology, Gliwice/
No. 2/2013
23
the temperature of a martensite transformation induced by a plastic strain (Md30),
which were SFE = 32.1 mJ/m2, Ms = -63.11°C
and Md30 = 22.7°C respectively.
Metallographic tests were carried out on
mechanically ground and polished longitudinal metallographic specimens. In order to
reveal their structure, the specimens were
etched in the so-called “aqua regia” heated
up to approximately 40°C. The observations
of the steel were carried out with a light microscope GX71 produced by OLYMPUS, using magnification from 100 to 1000x.
X-ray tests included the phase analysis
of the surface and of the middle layer of the
bands as well as the measurements of material textures at the initial state well as after
hyperquenching and plastic strain. The x-ray
phase analysis was conducted with a diffractometer D500, using a lamp with a copper
anode CuKα (λKα = 0.154 nm). Diffraction
lines were registered in the range of angle
2Θ from 40° to 92°, by means of a stepping
method, with a step of angle 2Θ equalling
0.02° and a pulse-counting time of 5 seconds
in one measurement position.
for martensite, the orientation distribution
function (ODF) and orientation fibres were
calculated. The calculations of the quantitative fraction of martensite α’ in the structure
of the steel involved the use of the magnetic
method.
The tests of mechanical properties were
carried out with a universal testing machine
ZWICK 100N5A, using a static tensile test according to standard PN-EN ISO 6892-1:2010
[11]. The samples for tests were cut out
of a sheet in parallel to the direction of rolling.
The hardness measurements of steel
AISI 304 were carried out by means of the
Vickers hardness tests, on metallographic
specimens under a load of 50g, using a hardness testing machine PMT-3.
Test results
Based on the metallographic tests, it was
possible to establish that steel AISI 304 at
the initial state is characterised by a structure composed of equiaxial austenite grains
with an average diameter of approximately 22 μm, containing annealing twins and
few spot non-metallic inclusions (Fig. 1a).
Fig. 1. Structure of steel AISI 304 at the initial state (a) and after 30% (b) and 70% (c) plastic strain, respectively;
etchant – aqua regia
The tests of the textures were carried out
using a diffractometer D8 Advance manufactured by the Bruker company and equipped
with the Euler’s wheel. The source of radiation was a lamp with a cobalt anode CoKα
(λKα = 0.179 nm). On the basis of three incomplete polar figures of planes {111}, {200},
{220} for austenite and {110}, {200}, {211}
24
The steel was characterized by a similar
structure after hyperquenching.
After cold plastic strain within a 10% ÷
20% strain range, the steel revealed a structure composed of elongated grains γ with slip
bands, deformed twins and non-metallic inclusions. The elongated character of austenite
grains corresponds to the crushed condition of
No. 2/2013
The diffraction patterns prepared for the
the steel, where austenite grains undergo elongation in the direction of rolling. The structure surface and middle layers of the steel at the
of the steel after over 30% strain also revealed, hyperquenched state did not reveal any sigin addition to elongated austenite grains with nificant changes as to the intensity of the indeformed twins and non-metallic inclusions, dividual diffraction lines originating from
few areas of parallel lamellas characteristic of the phase γ when compared with the diffraction patterns of the steel at the initial state
martensite α’ (Fig. 1b).
During cold rolling of steel AISI 304, (Fig. 2a,b). The phase α’ at the hyperquenched
an increase in strain is accompanied by the state was not revealed.
The presence of the diffraction lines origformation of new phase α’ dividing elongated austenite grains, which results in the inating from the martensite phase on the
so-called “refinement” of the steel structure diffraction patterns of steel AISI 304 at the
and hardening of the steel (Fig.1c). The met- initial state reveals the process of the phase
allographic observations revealed that the change γ → α’, where the martensite reamount of the phase α’ in the structure of vealed in the steel at the initial state might
the steel tested increases along with the steel have been formed during the initial treatment
of the material.
strain degree.
The diffraction image of the surface of
The results of the diffraction phase qualitative analysis of steel AISI 304 at the initial steel AISI 304 after cold plastic strain withstate (SD), hyperquenched (PP), and plas- in a 10% ÷ 40% range contains peaks (111)
tic-strained within a 10% ÷ 70% range are γ, (200)γ, (220)γ and (311)γ originating from
austenite (γ) as well as peaks (110)α’, (200)
presented in Figure 2.
The diffraction phase analysis of steel AISI α’ and (211)α’ originating from martensite α’
304 at the as-delivered state revealed diffrac- (Fig. 2a). After further straining of the steel
tion lines originating both from the phase γ (over 50%), in the diffraction pattern one
and α’ (Fig. 2a, b). The diffraction patterns can observe the disappearance of diffraction
prepared for the surface of the steel at the in- lines (200)γ and (311)γ originating from ausitial state contain four diffraction lines origi- tenite and the intensification of diffraction
nating from the phase γ, corresponding to the lines (200)α’ and (211)α’ originating from
planes {111}γ, {200}γ,
{220}γ and {311}γ. There
is also one peak (110)α’
originating from the martensite phase (Fig. 2a).
Identical diffraction lines
can be observed in the
diffraction patterns prepared for the middle layer
(Fig. 2b). In addition, the
middle layer of steel AISI
304 at the initial state revealed weak peaks (200)
α’ and (211)α’, originating
from the martensite phase Fig. 2. Diffraction patterns of steel AISI 304 at the initial state (SD), hyperquenched (PP), and plastic-strained within 10% ÷ 70% range: a) surface, b) middle
(Fig. 2b).
No. 2/2013
25
martensite. In the whole range of strains, the
strongest peak originating from austenite was
(220)γ. In turn, the intensity of peak (111)
γ underwent changes. After the maximum,
i.e. 70% degree of strain, the strongest peak
originating from martensite was peak (211)
α’. It was also possible to observe a widening of the diffraction lines originating both
from phases γ and α’, which was associated
with an increase in structural defects formed
during the plastic strain (Fig. 2a). In the diffraction patterns prepared for the middle layers of steel AISI 304, no significant changes
in the intensity of individual diffraction lines
were observed (Fig. 2b).
The conducted diffraction tests revealed
that diffraction lines (111)γ, (110)α’; (200)α’,
(220)γ, (211)α’ of (311)γ of the tested phases
of steel AISI 304 after cold rolling with 40%
strain reveal distinct steel texturing (Fig. 2a,
b). The diffraction phase analysis of the steel
deformed within a 10% ÷ 70% range did not
reveal diffraction lines originating from the
phase ε, which is consistent with information found in reference publications [1-10].
A phase change proceeds directly according
to the sequence γ → α’.
The texture of the austenite of steel AISI
304 both at the initial and hyperquenched
state was relatively weak (Fig. 3a, b). Yet, it
should be mentioned that the austenite of the
tested steel is a metastable phase and that the
development of the steel texture is complex.
During the plastic strain of austenite the
following processes take place at the same
time: austenite texturing, phase change γ →
α’ and the change in orientation of the martensite formed during the strain. The texture
of the steel strain is therefore described by
the texture constituents as both of austenite
and martensite (Fig. 4 and 5).
The main constituent of the austenite
texture of the steel at the initial and hyperquenched state was a confined fibre α
(<110>║ND (ND – normal direction), in
26
which the strongest orientation was close
to the orientation {110}<112> of the alloy
type. The maximum value of the orientation
distribution function for this orientation was
ODF = 3.9 for the initial state and ODF = 3.3
for the hyperquenched austenite respectively
(Fig. 3a, b).
a)
b)
Fig. 3. Austenite texture at the initial state (a),
at the hyperquenched state (b) on ODF cross sections
φ2=0°, φ2 =45°
In the texture of the austenite after the
strain within a 10% ÷ 70% draft, it was possible to observe orientations described by
the fibre α =<110>║ND, τ =<110>║TD (TD
– transverse direction), β ={110}<112> by
{123}<634> to {112}<111>. The strongest
constituent of the austenite texture of steel
AISI 304 was the orientation {110}<113> of
the fibre α =<110>║ND, which is close to the
constituent of the alloy type {110}<112>. It
is also possible to observe the Goss orientation {110}<001> of the fibre α =<110>║ND
(Fig. 4a and 5a). The increase in the strain
was accompanied by the reinforcement of the
austenite texture. During the conducted tests
it was possible to observe that an increase in
the strain degree is accompanied first by the
No. 2/2013
elongation, and next by the contraction of the composing the confined fibre α1 and the oriaustenite fibre. The texture of the austenite entation {111}<112> being the main constitdeformation is typical of materials with low uent of the homogenous fibre γ. Rolling the
and medium value of SFE (Stacking Fault steel with a 70% deformation degree causEnergy).
es that in the martensite texture the fibre
During the plasa) Austenite
b) Martensite
tic strain martensite is
formed in the structure
of the steel tested. The
presence of the martensite and an increase in its
content is the result of
the phase change γ →
α’ induced by the strain.
An increase in the strain
is accompanied by a
change in the martensite texture, caused by the
texturing of the initial
phase of the austenite,
from which the phase α’
is formed. The texture
of martensite after the
strain within a 10% ÷
70% range is described
by the fibres of the orientation α1 =<110>║RD
(RD – rolling direction),
γ ={111}║ND and ε
=<001>║ND. The dominant orientation of the
martensite texture of
steel AISI 304 was the
orientation {111}<112>
of the fibre γ ={111}║ND
(Fig. 4b and 5b). After
30% strain, in the martensite texture, one can
observe a rotated cubic
orientation {001}<110>
(Fig. 4b). In turn, after
70% strain, the texture
of strained martensite Fig. 4. Orientation distribution function for steel AISI 304 after various degreis dominated by the ori- es of restraint presented in cross sections φ2 =0°, φ2 =45° for austenite (a) and
φ1 =0°, φ2=45° for martensite (b)
entation
{112}<110>
No. 2/2013
27
γ ={111}║ND is stronger than the fibre
α1 =<110>║RD (Fig. 4b and 5b). The martensite texture remained weak within the
whole range of restraint.
The crystallographic relationships between
the texture of austenite and of martensite
formed during the strain are best described
by the Kurdjumow-Sachs (K-S) and Nishiyama-Wassermann (N-W) relationships.
50% range in the steel is accompanied by a
change in its tensile strength from approximately 784 MPa to approximately 1257
MPa, conventional yield point from approximately 586 MPa to approximately 960 MPa,
and elongation from approximately 32% to
approximately 2%. The maximum, i.e. 70%
strain of the steel, causes a significant increase in its values of Rm, Rp0.2 and HV0.05.
Tensile strength increases to approximately 1496 MPa, yield point to approximately
1161 MPa, and hardness to approximately
400 HV0.05. It is also possible to observe a
significant decrease of elongation i.e. to approximately 1% (Fig. 6).
Fig. 6. Changes in mechanical properties of steel AISI
304 in the function of plastic strain
Fig. 5. Values of orientation distribution function f(g)
along fibres α, τ, β for austenite (a) and fibres α1, γ, ε for
martensite (b) of steel AISI 304 after 70% strain
The conducted tests of mechanical properties revealed that the values of hardness
HV0.05, tensile strength Rm and conventional yield point Rp0.2 of steel AISI 304
increase along with an increase in the strain
degree, whereas the value of elongation A
decreases (Fig. 6).
At the initial state, the conventional yield
point for steel AISI 304 is approximately
330 MPa, tensile strength approximately 647
MPa, hardness approximately 162 HV0.05,
and elongation approximately 52%. An increase in the strain degree within a 10% ÷
28
The tests of mechanical properties confirmed the analytical dependence of the yield
point of austenitic steel AISI 304 on the strain
degree in the process of rolling.
Based on the analysis of the magnetic tests,
it was established that the amount of martensite phase in the structure of steel AISI 304
increases along with the strain degree in the
process of rolling. After 70% strain, the steel
contains approximately 28% of martensite α’.
Conclusions
The analysis of the test results for austenitic steel AISI 304 leads to the following
conclusions:
1. The plastic strain induces the martensite
transformation γ → α’ in the whole range of
applied strains.
No. 2/2013
2. At the initial state, the steel structure
is composed of equiaxial grains γ with an
average diameter of approximately 22 µm
with annealing twins and non-metallic inclusions, whereas after the plastic strain of the
steel with a draft of approximately 30% - the
structure of elongated austenite grains with
areas of parallel lamellas characteristic of the
martensite α’.
3. The texture of the strained austenite is described by orientation fibres α
=<110>║ND, τ =<110>║TD, β ({110}<112>
by {123}<634> to {112}<111>); this texture
is typical of materials with a low value of
SFE.
4. An increase in the strain degree is accompanied by the development of the martensite texture; its main constituents are the
orientations of the fibre α1 =<110>║RD, γ
={111}║ND and ε =<001>║ND.
5. The fraction of the martensite phase α’
in the steel structure increases along with an
increase in the steel strain degree. After the
maximum, i.e. 70% strain, the steel contains
approximately 28% of the phase α’.
6. The changes in the volume fractions of
the phases γ and α’ during the cold strain of
steel AISI 304 and the texture development
in these phases, affect the mechanical properties.
The research was conducted within
the confines of research project
no. 2632/B/T02/2011/40 funded by
the National Science Centre
(Narodowe Centrum Nauki).
References
1. Donadille C., Valle R., Penelle R.: Development of texture and microstructure during cold-rolling and annealing
of fcc alloys: Examples of an austenitic
stainless steel. Acta Metallurgica, 1989,
vol. 37, s. 1547.
No. 2/2013
2. Abreu H., Carvalho S., Neto P., Santos
R.: Deformation induced martensite in
an AISI 301LN stainless steels. Materials
Research, 2007, Vol. 10, s.359.
3. Angel T.: Formation of martensite in austenitic stainless steels: Effects of deformation, temperature and composition.
Journal of the Iron and Steel Institute,
1954, s.165.
4. Ozgowicz W., Kurc A.: The effect of the
cold rolling on the structure and mechanical properties in austenitic stainless
steels type 18-8. Archives of Materials
Science and Engineering, 2009, vol. 38,
s.26.
5. Reed R.: The spontaneous martensitic
transformations in 18%Cr, 8%Ni steels.
Acta Metallurgica, 1962, vol. 10, s.865.
6. Kurc-Lisiecka A., Kalinowska-Ozgowicz
E.: Structure and mechanical properties
of austenitic steel after cold rolling. Archives of Materials Science and Engineering, 2011, vol. 44, s.148.
7. Kowalska J., Ratuszek W., Chruściel
K.: Crystallographic relations between
deformation and annealing texture in
austenitic steels. Archives of Metallurgy
and Materials, 2008, vol. 53, s.131.
8. Singh C. D., Ramaswamy V.: Development of rolling texture in an austenitic
stainless steel. Textures and Microstructures, 1964, vol. 12, s.145.
9. Łuksza J., Rumiński M., Ratuszek W.,
Blicharski M.: Texture evolution and
variations of α’-phase volume fraction
in cold rolled AISI 301 steel strip. Journal of Materials Processing Technology,
2006, vol. 177, s.555.
10.Padilha A.F., Pault R.L., Rios P.R.:
Annealing of cold-worked austenitic
stainless steels. ISIJ International, 2003,
vol. 43, s.135.
11.Norma PN-EN ISO 6892-1:2010 Metale.
Próba rozciągania. Część 1: Metoda
badania w temperaturze pokojowej.
29
Olga K. Makowieckaja
Technological innovations – a basis for the increase
of competitiveness of welding industry in the USA
The threat of losing its position as leader
in the global economy causes an increasing
concern to the US government and business
community. In recent years the USA has lost
its leadership in competitiveness, moving
from first place in 2009 to fifth in 2011, and
in 2010 the USA lost its position as number
one in industrial production to China [1].
Industrial production, being the bedrock
of the US economy, makes up 11% of the
country’s GDP, with the export of industrial
products constituting over 60% of total export. Industry employs approximately 13.4
m people, i.e. almost 9% of the total number
of people employed. Remunerations in the
industrial sector are over 20% higher than in
other extra-industrial sectors of the economy.
Since 2008 the economic crisis has remained the major issue of the US economy.
Yet negative tendencies could be observed as
early as 2001 when 2.5 m jobs were cut in
the industrial production sector in just over
a year. Experts indicate the following worrying trends in the US industrial production:
• decreasing industrial production; share
of industrial production in GDP in 20002010 fell from 17 to 11%.
• employment reductions; in 2000-2010 the
industry saw a decrease in employment by
37% (6.5 m),
• decrease in foreign trade; the US global
market share went down from 19 to 11%
(2000-2010), which has caused a foreign
trade deficit,
• growing prices of industrial products; increasing outlays related to industrial safety, environmental protection, taxes, remu-
nerations, complaints etc. affected prices
of finished products, which has decreased
US competitiveness in the global market,
• shortage of skilled labour [2].
Technologies for joining materials are an
indispensable part of the economy’s industrial sector. Welding engineering and related
technologies are closely integrated with production processes in basic sectors of the industry. They are of key importance and cannot be replaced by other alternative solutions.
Having in mind the significance of joining
technologies for the economy, in 2010 Edison Welding Institute (EWI) in conjunction
with the American Welding Society (AWS)
initiated extensive investigation into the condition and possible ways of improving the
competitiveness of industrial production using materials joining as an example. Within
the project “Future trends in materials joining in the USA”, manufacturers representing
six core industrial sectors were surveyed in
order to uncover the major problems of these
sectors and their needs related to materials
joining. The results of the investigation were
published in February 2011 at the sum-up
conference “Strengthening Manufacturing
Competitiveness: the Future of Materials
Joining in North America, attended by the
representatives of scientific, governmental
and social institutions as well as of the welding engineering market leaders – Lincoln
Electric, Trumpf, Miller Electric and others.
The final conference document outlines the
main problems in the field of materials joining and the tasks for the next five year.
In today’s globalised economy there is
K.e.n. Olga K. Makowieckaja – the E.O. Paton Electric Welding Institute of the National Academy
of Science, Kiev, Ukraine
30
No. 2/2013
only one possibility of increasing industrial
competitiveness, namely, through the development of innovativeness, i.e. increasing the
level of solutions, shortening the time for
industrial implementation of technical innovations and improving personnel skills. The
innovative development of the economy also
entails intense relations between science, education and personnel training. According to
the aforesaid survey carried out by EWI, the
problems appearing in the area of materials
joining are closely connected with solving
the tasks mentioned above. The results of the
survey are presented in Tables 1 and 2 [3].
The results presented in the tables reveal
that in all industries, new structural mate-
rials and their composites are becoming
increasingly popular (Table 1). This is a
major task in the automotive industry and
power engineering. It also tops the tasks
lists of other industries. Research workers,
designers, constructors and manufacturers
are more and more interested in applying
new materials improving technical characteristics of elements and reducing their production costs. For instance, the necessity
of reducing a car body mass increased the
expansion of high-strength steels, aluminium, magnesium alloys and composites. An
increase in the use of new structural materials requires the development of new joining technologies (Table 2). This problem is
Table 1. Problems and the most important tasks in US materials joining engineering in the next five years
(the first 4 issues vs. industries)
Automotive
industry
Problems and tasks
Shortage of highly qualified
engineers and specialist in joint
quality inspection
Shortage of skilled welders and
workers of other professions
Greater competition from
countries with lower labour costs
Greater outlays on developing
and implementing new processes,
products and methods
Increase in time needed for
assessing the quality of joints
Increase in use of new materials
and their composites
Implementation of new
technological processes
Shortened time between the
development of a solution and its
industrial implementation
Development of on-line systems
reporting the latest technologies
and methods, and providing
access to them
Increased requirements related to
the quality of joints
No. 2/2013
Order of importance in industries
PetroHeavy
Military
Space
Power
chemical
machinery
industry industry
engineering
industry
industry
1
4
3
1
3
2
2
3
1
4
3
4
2
1
1
3
4
4
2
4
1
3
2
2
31
Table 2. Materials joining technologies and other works essential for industry (the first 4 issues vs. industries)
Order of importance in industries
AutoPetroHeavy
Military Space
Power
motive chemical
machinery
industry industry
engineering
industry industry
industry
Required technologies/works
Development of technologies for joining wend advanced materials
Increasing the number and improving
qualifications of engineers and
constructors dealing with in joining
technologies
Development of arc welding
processes (efficiency, quality etc.)
Development of new methods for
joining dissimilar materials
Providing on-line access to databases
with materials joining technologies
Development of more sensitive,
accurate and failure-free NDT methods
Development highly efficient
technologies for welding materials
of great thickness
Improvement (updating, channelling
and making cheaper) methods for
training welders
Development of policy for
development of new joining processes
Modernisation of resistance welding
technologies (quality, reliability etc.)
Development of additive (supporting)
industrial technologies
1
1
2
2
1
3
2
1
3
3
1
1
2
2
4
3
4
4
2
4
4
4
mentioned by the representatives of all the
surveyed industries, and in the military and
space industries it is particularly visible.
The respondents also maintain that it is necessary to shorten the time for implementing
new solutions in industrial production, find
ways to reduce the costs of developing and
implementing innovations, and develop online databases with new solutions related to
materials joining. In short, it is necessary to
develop policies allowing the development
of joining technologies (Table 2).
The second task in the order of importance for all the representatives is to ensure
qualified personnel specialised in joining
technologies. According to the Office of Statistics, in the USA between 2002 and 2009
32
3
3
the number of workers of all welding-related jobs dropped from 1 076 498 to 968
037 or, expressed in percentage, by 10.08%.
However, this number may be higher, as requirements concerned with the professional
skills of a welder are higher than in 25 other
jobs.
The data obtained by means of the survey and confirmed by the statistics reveal
that individual US industries suffer from a
shortage of skilled welders, welding engineers, and other joining and quality control
specialists. For instance, the main issue in
the petrochemical industry is the shortage of qualified engineers and specialists
in joint quality inspections, whereas the
heavy machinery industry has an insuf-
No. 2/2013
ficient number of welders (Table 1). The
shortage of qualified personnel is strictly
connected with the problem of personnel
training system improvement as well as
with the development and implementation of the system ensuring permanent improvement of qualifications of specialists
representing all professions [4].
The major source of innovation is research and scientific, experimental and design engineering works. American experts
believe that the total global outlays on research funding in 2012 will increase by 5.2
% and an amount of 1.4 trillion USD, where
the US share will be 36%, i.e. 436 billion
USD. Research is financed by industry
(64%) and the federal government (29%).
Table 3 presents data related to the structure
of financing research and scientific, experimental and design engineering works in the
USA, broken down into the main financing
sources and contractors.
Research in the USA, similarly as the
world over, is an area open to collaboration.
The data from Table 3 show that industry in-
creasingly finances its own research as well
as basic research conducted by scientific institutions for the industry. The federal government also puts considerable funds into
research activities initiated by industry and
institutions. According to the results of a survey conducted by the R&D Magazine, 80%
of industrial companies finance research carried out along with research centres and other
organisations. It should be emphasized that
not only industry but the federal government also shows a growing interest in profiting from outlays on research. A few years
ago only 10% of companies planned and
calculated profit from such “investments”.
Today, over 50% of businesses consider this
factor as a key indicator of their activity.
The Bayh-Dole Act of 1980 created the
basis for a new US research and technical
policy, the aim of which is to increase the
competitiveness of the national economy.
The law allowed passing the right of intellectual property financed from governmental resources to other non-federal research institutions such as universities, private businesses
Table 3. Ratio “fund source – research contractor”, 2012, m USD (change in relation to 2011- %)
Fund source
Federal government
Governmental
Federal
funds, national
government
centres and
laboratories
29 152
-2,51%
Industry
14 666
-3,69%
202
2,20%
Contractor
National
science fund
Non-profit
Total
Industry
and other
organisations
academic
institutions
125 652
37 577
37 440
6 817
-1,61%
-2,42%
0,93%
-2,29%
279 685
237 487
3 868
2 129
3,75%
3,37%
26,49%
8,89%
National science fund
and other academic
institutions
Other governmental
institutions
Non-profit
organisations
Total
No. 2/2013
29 152
-2,51%
14 868
-2,36%
311 063
2,63%
12 318
2,85%
12 318
2,85%
3 817
2,72%
3 491
2,70%
60 934
2,85%
3 817
2,72%
14 546
2,70%
436 018
2,07%
11 055
2,80%
20 001
1,55%
33
Treatment
Additive
technologies
Joining
Quality
control
Press
forming
Electronics
assembly
34
Casting
Automation
and other entities as well as enabled making companies interested in the development of
invention licences available on the basis of new advanced technologies. The Consortiexclusivity, which is the primary condition um members specify the basic technological
problems which must be solved, agree on the
for their commercialisation.
This law, other governmental decisions programme of a project and choose the manadopted later, and state-run programmes agement of the consortium group. In order
stimulated the integration of basic and ap- to solve various special tasks the Consortium
plied research, increased industrial compa- can invite centres for the development and
nies’ interest in basic research, contributed implementation of industrial technologies,
to the expansion of interdisciplinary research research laboratories, commercial companies
and changed the attitude to research infra- and other organisations as collaborating contractors. The state supports the development
structure [5, 6].
In order to stimulate technological research of innovations until their commercialisation
related to materials joining, strengthen mutual by means of state programmes. Industrial
relations between science and industry, sig- implementations of innovations require significantly shorten implementation time and nificant involvement of industrial and other
expand areas of innovative solutions, EWI to- resources. Table 4 presents the scheme of
gether with the US Institute for Industrial Pro- collaboration for the Consortium and centres
ductivity worked out a successful model of for the development and implementation of
developing and industrial implementation of industrial technologies.
technological innovations related to join- Table 4. Scheme of collaboration of the Focused Industry Consortium and centres for
the development and implementation of industrial technologies
ing technologies. The
basis of the model is
Centres for the development and implementation
of industrial technologies
the idea of creating
new organisational Focused Industry
structures favouring Consortium
closer integration of
all the participants
in the innovative Production of metal
process – from con- for aviation industry Х
Х
Х
Х
Х
cept to development, using additive
technologies
commercialisation
Car body mass
Х
Х
Х
Х
Х
and extensive indus- reduction
trial implementation Quick assembly of
Х
Х
Х
Х
of innovation. Such batteries
structures can be the Environmentally
Х
Х
Х
following: Focused friendly production
Industry Consortia, of electronics
and Manufacturing Production of units
for nuclear power
Х
Х
Х
Х
Х
Technology Applica- stations
tion Centres.
Automation of
The
Consorti- production of
Х
Х
Х
Х
Х
um is a temporary machines for heavy
union of industrial machinery industry
No. 2/2013
The scheme above presents one of the
main ideas related to the functioning of the
Consortium, i.e. the possibility of involving
specialised centres for the development and
implementation of industrial technologies,
supported by their highly qualified experts
and possessing necessary funds, to solve
specific tasks during the development of
specific innovations.
The purpose of the consortium model developed by EWI is to demonstrate a demand
for new materials joining technologies, invent such technologies and create a programme of partnership-based collaboration
for the development and quick industrial
application of new technologies. An example demonstrating how this model functions
in practice is the Additive Manufacturing
Consortium and Nuclear Fabrication Center
created by EWI in 2010.
For instance, the Additive Manufacturing
Consortium joined efforts of large US space
industry corporations taking advantage of
the research provided by EWI and other private, non-profit and state organisations interested in the development and widespread
industrial implementation of leading additive technologies. The consortium included
24 industrial companies and research centres. The industrial partners of the consortium were users and manufacturers, whereas
the research partners included universities
and such organisations as the Army, the Air
Force, the Navy, NIST and NASA. The development and implementation of this model was supported by the state. In order to
implement the project, the state of Ohio established a multimillion dollar grant.
If the purpose of the consortia is to solve
strategic and organisational tasks aiming
No. 2/2013
to develop new technologies, the main collaborating contractors in given projects are
centres dealing with the development and
implementations of industrial technologies.
Such centres should be global market leaders in their sectors, possessing cutting-edge
equipment and employing highly qualified
personnel. An example of such a centre in
materials joining is EWI, which within the
scope of its activity collaborates both with
universities and industrial companies. Such
an approach favours the development of innovative solutions successfully implemented in production. Since 1984 EWI has been
taking advantage of state aid within the
confines of the Ohio Edison Programme.
Permanent development, efficient solutions
and high return on outlays are the factors
which attract private investors. In 2010 private investments in the research works of
EWI were almost 20 times higher than the
contribution from the state [7].
The model of developing and implementing technologies has been approved by the
US government. In 2011, on the basis of
this model, the National Institute of Standards and Technology at the US Department
of Commerce adopted a new national programme for supporting the development
of technological innovations in the USA
named ”Advanced Manufacturing Technology Consortia (AMTech)”. In 2012 the
budget of this programme amounted to 12
m USD. The aim of the programme is to
support innovation-oriented tasks such as
robotics technology, nanomaterials, new
advanced materials and new production
technologies. In total, in 2012 the state supported innovative programmes with a sum
of 75 m USD [8].
35
References:
1. Bucher K.: US Competitiveness Ranking
Continues to Fall; Emerging Markets Are
Closing the Gap. www.weforum.org
2. Strengthening Manufacturing Competitiveness. Report from the 2010 Conference on the Future of Materials Joining in
North America. February 2, 2011. EWI.www.ewi.org
3. Conrardy C.: Materials Joining and Technology. www.weldingandgasestoday.org
4. State of the Welding Industry
Report: Executive Summary. WeldEd.
www.welded.org
36
5. 2012 Global R&D Funding Forecast.
Battelle. The Business of Innovation.
www.battelle.org
6. Дежина И. Поддержка
фундаментальной науки в США:
уроки для России? Бытие науки. 2011,
№ 94. с.6-7
7. Revitalizing American’s Manufacturing
Innovation Infrastructure. Response to
the NIST AMTech Request for Information. EWI. www.ewi.org
8. President Obama Launches Advanced
Manufacturing Partnership.
www.nist.gov
No. 2/2013
Antoni Sawicki
Damping Factor Function in AC Electrical Arc Models
Part 1: Heat Process Relaxation Phenomena,
their Approximations and Measurement
Introduction
Temporary changes of the column
cross-section radius and the distribution of
energy along and across an arc significantly affect the operation of electrotechnological devices and electric appliances. They are
decisive for the possibility of breaking and
re-igniting the arc. A quantity characterising
such possibilities in the most complete manner is a function for the factor of damping
energetic processes in an electric arc plasma.
A coarse, and yet very comfortable approximation of this function is the time constant.
The knowledge of the arc time constant is
necessary for the following:
• selecting economical operating conditions
of electrotechnological devices (welding,
electrothermal etc.);
• optimum influence on the arc in switching
devices (breakers, switches, contactors,
relays);
• maximally intense influence on the arc in
overcurrent and overvoltage protective
devices (fuses, lightning protectors).
The arc may stop during voltage reduction, excessive stretching of the column, excessive current reduction, cooling of the gas
area (sometimes also of the electrode) or as
a result of contracting the column with a diaphragm. The above phenomena trigger deionisation in the plasma column and cooling of
the active areas in the electrodes. Depending
on the intensity and duration of these processes, re-ignition of the arc may be difficult
or even impossible. For this reason, the time
constant should be the following:
• low in electric appliances so that the arc
can stop relatively quickly due to disturbances;
• high in electrotechnological equipment, in
order to prevent undesired terminating of
the arc due to disturbances.
Heat process relaxation phenomena
in electric arc
Arc columns have a heat capacity, and yet
they constitute certain resistance to thermal
current. For this reason they have finite times
of response to forced changes of thermal
states. The amount of internal energy, accumulated in the arc, depends on many factors:
• plasma volume (radius, length and shape
of the column);
• temperature distribution in the column;
• pressure of the plasma-forming gas;
• type of plasma-forming gas, degree of
plasma-forming gas ionisation etc.
The enthalpy of the arc Q changes exponentially in time in accordance with the time
constant [1]
θ=
1
dP0 dR 2
−
I
dQ dQ
(1)
where P0 – dissipated power; R – arc resistance. Hence one can see that θ depends on
current I. In the case of alternating current,
the factor θ depends on the momentary value
of current.
As the electric current is characterised by
thermal inertia, the changes of the thermal
state and geometrical dimensions of the col-
dr hab. inż. Antoni Sawicki, professor at Częstochowa University of Technology - Faculty
of Electrical Engineering
No. 2/2013
37
umn during current rushes or changes in the
column length are not immediate, but proceed with a certain time constant. That is why
arc resistance during the changes of current
r(i) alters exponentially from the stabilised
value r(I0) to the value corresponding to new
current r(I1). The time constant of the arc in
constant pre-set heat transfer conditions is a
time after which the arc column changes its
resistance e–times after energy is no longer
supplied to the column.
The time constant of the arc is defined by
the column cooling conditions. In plasmatrons it is shorter by 2÷3 orders (10-6 ÷10-7
s) if compared with ordinary free arcs. Even
at a frequency of 200 kHz the arcs of AC
plasmatrons have a hysteresis, which reflects
their low time constant. The higher the flow
rate of the gas flowing around the column or
the rate of the arc motion in gas, the lower
the time constant is. In the case of high gas
flow rates, the arc time constant does not depend on the gas chemical composition or the
type of electrodes [2].
When the arc is ignited or terminated the
energy accumulated in the plasma volume
unit is greater than the energy dissipated
from this volume. For this reason, in furnaces with a high temperature of the atmosphere
it is easy to observe the thermal breakdown
of the inter-electrode gap with relatively low
voltage. In turn, if the temperature of the atmosphere is low, the reproducing strength
of the inter-electrode gap rises quickly to a
certain initial value according to a certain exponential dependence. In order to be able to
reliably re-ignite the arc, the rate of power
source voltage reproduction should be higher
than the critical value [1]. This principle is
the basis for testing and assessing the quality of dynamic characteristics of welding
sources. The lower limit of the reproducing
voltage is defined by the arc ignition voltage.
The rate of increasing reproducing strength
of the arc column is defined by the inertia of
38
heat processes, and especially by the arc time
constant θ. In turn, the dynamic properties of
the source are affected by the design and settings of the control system as well as by the
passive conservative elements (inductance
and capacitance) of its circuits.
The gas suppression ability is defined not
only by its time constant but also by its electric strength. The reproduction of the electric
strength of the inter-electrode section strongly depends on the falling rate of temperature
T of the plasma left after the arc column. This
can be roughly determined using the following dependence [1]:
T = Tot − (T0 − Tot ) ⋅ e
−t
θ
(2)
where Tot – ambient temperature; T0 – temperature on the arc axis at the beginning of
the process; θ - arc time constant. The electric strength of the inter-electrode gap is inversely proportional to temperature [1]
ET =
Tot
Eot
T
(3)
where ET – electric strength at heightened
temperature T; Eot – electric strength at ambient temperature Tot.
Approximating the factor of energetic process damping in electric arc
In the majority of simplified mathematical models, very roughly approximating the
physical properties of the electric arc, the
damping factor value of transitory processes (thermal and electric) is adopted as a constant quantity (the so-called time constant).
It is the proportion of two quantities; the
numerator is the heat capacity of the plasma
channel, whereas the denominator is made
up of parameters specifying the properties of
energy dissipation [3]. More accurate models (e.g. Cassie-Mason or Lowke’s) bind proportionally the time constant value with the
cross-sectional area of the plasma column
[4, 5]
No. 2/2013
2
In most operation regimes the gas atmosphere of the DC arc furnace is made up by
where p - gas density; cp – specific heat of carbon oxides. Another formula for the digas of pressure p; λ - gas heat conduction ameter was provided by R.T. Jones and Q.G.
Reynolds [14]:
factor; ra – arc column radius.
The confirmation of this assumption was
also attempted using the approximation of d = 2r ⋅ 3,2 − 2,2 exp − z 
(7)
K 
 5r 
experimental data [6], where one should bear a
K 


in mind that the column cross-sectional area
depends not only on electric current intensi- where the cathode spot radius is
ty but also on the type and pressure of plasI
(8)
ma-forming gas, the temperature of the gas in rK =
, cm
π
j
K
the discharge area, the rate and direction of
gas flow, the diameter of the discharge duct, and jK = 3500 A/cm2 – current density in the
the amplitude and frequency of the magnetic cathode spot.
field etc. In some other models of the arc the
However, the function of the column ditime constant is adopted as being proportion- ameter da(i) is not monotonic in the range
al to the column diameter (e.g. the model by of weak currents. Plasma does not disappear
A.A. Woronin [7-10]).
along with the momentary reduction of curIn relation to relatively long arcs the short rent to zero. Yet, the weakening of the pinch
conical part at the cathode is negligible and effect may cause its expansion, accompathe shape of the whole plasma column can be nied however by some cooling and deterioassumed as cylindrical. Theoretical deliber- ration of electric conductance. This, in turn,
ations and experimentation are significantly depends on the conditions of heat exchange
simplified if one determines the geometrical in the environment and the rate of current
dimensions of the DC arc. Results obtained changes during polarisation alteration. Such
in this way can be adopted for low-frequency behaviour is confirmed by the experimenAC arcs, assuming the necessity of maintain- tal tests of the arc time constant, which ining within them the same plasma equilibrium. creases significantly in the range of weak
In welding arcs the diameter of the arc currents, below approximately 18-30 A [15,
column is, first of all, the function of current 16]. The flow of current through such termida = f(I2/3) [11, 12]. This function is very nated arc plasma is possible after applying
voltage from an additional source [17]. If
close to the empirical formula [1]
≅ × , cm
(5) currents are weak, the arc time constants are
not only high but also strongly dependent on
where n = 0.6÷0.7, obtained in the case of gas the type of gas (e.g. in elgas, SF6, θ = 1÷2 µs,
flowing around the arc in a longitudinal man- in the air θ = 100÷200 µs). In the range of
ner. If the arc is in the air, k = 0.27 cm×A-n. strong currents the tendencies of time conIf one considers a strong-current arc, e.g. stant changes are reverse to the changes of
in a steelmaking arc furnace with a graphite column diameter (cross-sectional area). An
cathode, burning in air, the measured column increase in current intensity as well as an
increase (with very weak saturation) in the
diameter amounts to approximately [13]:
column diameter function (formulas (5)-(7))
0,5

z 
d a = 2rK ⋅  0,864 − 0,253 
(6) are accompanied by a decrease in the time
r
K 

constant, which stabilises at the lowest lev-
∝
No. 2/2013
(4)
39
el. Time constants approaching one another
and constituting approximately 10-4 s correspond to the values of current counted in
hundreds and thousands of amperes, flowing
through arcs in various gases. Therefore, the
direct binding of the damping factor function
by the proportionality formula (4) with the
cross-sectional area θ(S(i)) or the diameter
θ(da(i)) of the arc was not confirmed in practice (Fig. 1). Such an outcome can reflect the
significant influence of the variability of other plasma parameters e.g. θ(da(i), cp(i), p(i),
λ(i)), although in analytical deliberations it is
not openly expressed [18].
a)
The data obtained in experimentation
[15] reveal that the column structure can be
strongly heterogeneous. In such a case the
core of the column is composed of plasma
characterised by a very high temperature
(significantly over 8000 K), very low viscosity and low time constant θf. In turn, the plasma subsurface layer of a lower temperature
(6000-8000 K) has the highest viscosity and
a higher time constant θs. In calculations it is
usually assumed that θf < θs = θ. Depending
on the type of gas, arcs burning in the areas where the atmosphere temperature is high
take on a diffusive form and have one time
constant.
As there is no specific universal analytical
expression describing dynamic current-voltage characteristics of the arc, there is no resultant specific expression for the time constant either. In such a situation, probably the
most appropriate approximation is a function
dependent on current in a non-linear manner
e.g. [19]:
(| |) =
0
+
1
(− | |) ≈
| | small
0 , if | | large
1 , if
(9)
where α > 0, θ1 >> θ0 – constant approximation factors. The necessity of taking into consideration the non-linear damping factor
function is especially visible in the hybrid
models of the arc [19, 20] as they more precisely reproduce dynamic characteristics in
wide ranges of current intensity. Another
popular solution is to make the time constant
dependent on arc conductance (Fig. 1b).
Usually used for such purposes are the
Schwarz-Avdonin models [8], in which the
dependence has the following form:
b)
θ = θ0 g α
Fig. 1. Arc column dynamic characteristics: a) as current
intensity functions; g) as conductance functions (PM –
power of the Mayr-Schwarz model, UC, PC – voltage
and power of the Cassie-Schwarz model)
40
(10)
where θ0, α - approximation factors.
If the basic factor facilitating the termination of AC arc is an insufficiently high momentary value of the current intensity mod-
No. 2/2013
ule, at such a moment this value matches the
maximum damping factor value. The said
value can be determined by means of appropriate measurement methods [21] utilising the delay and increase in the re-ignition
voltage. From the point of view of ensuring
the arc burning stability and the continuity
of electrotechnological device operation it
is justified to experimentally determine such
highest value θ. Due to the range of weak
current, this constant can be used in the Mayr
model. However, the time constant determined in the weak-current range is sometimes
also used in the Cassie model [22], which
may lead to discrepancies of experimentation and calculation results, especially in the
range of strong currents. Therefore, if it is
necessary to precisely reproduce the courses
in circuits by means of the universal model
of strong-current arc (e.g. hybrid TWV), the
whole damping factor function θ(i) should
be expressed by means of dependence.
Experimental methods for determining AC arc dynamic parameters
A characteristic feature of experimental
methods for determining arc dynamic parameters is to take into account the whole
range of physical phenomena taking place in
the column, near-electrode areas and in electrodes themselves. The separation of individual components of energy processes is very
difficult but sometimes possible by means of
analysis [11, 23, 24]. Depending on the design and the principle of operation of electrotechnological arc or plasma-arc devices,
applied technologies and operation modes,
one can observe various levels of disturbances both as to the amplitude and the range of
frequency, which can even exceed the values
allowed by related standards. Such disturbances may originate from sources which are
difficult to identify or eliminate and which
disturb processes taking place in the arc column, electrodes, and even in the circuits of
No. 2/2013
measurement systems. As the quantities u(t)
and i(t) are registered along with random disturbances, calculating the values of conductance g(t) on their basis comes down to solving
a badly conditioned task. An improvement
in the quality if input data can be obtained
using appropriate methods for filtering and
smoothing time courses [25]. However, prior
to undertaking such actions it is necessary to
solve the issue of recognising types of disturbances in order to weaken only the impact of
natural disturbances and leave disturbances
triggered on purpose.
Methods for determining the dynamic parameters of the AC arc can be divided into
several groups:
1) methods using natural periodic courses of
current and voltage;
2) methods introducing additional disturbances to periodic courses, using additional current sources;
3) methods introducing disturbances of the
arc column length (voltage);
4) methods introducing disturbances of conditions of energy dissipation from the column [7, 21].
In the case of the methods utilising natural
electric courses it is assumed that there is an
unequivocal functional relationship between
arc parameters and current, resistance or conductance. It means that, in specified conditions of heating and cooling the column, one
set of model parameters corresponds to one
value of current or arc conductance. Using
properly processed (filtered and/or smoothed)
data, one can apply one of the analytical or
analytical-graphic methods (known as Amsinck, Ruppe, Asturian, Rijanto, Zuckler, Tajew, generalised etc.) in order to determine
the simple parameters of the Mayr or Cassie
models [22]. More complex models require
the use of numerical methods.
There are also possibilities of the direct
determination of the arc time constant, not
requiring the calculations of the remaining
41
parameters of specific mathematical models.
In the low-voltage arcs of sinusoidal alternating current and in the conditions of
relatively low cooling intensity, before and
after the passage of current through zero, it
is possible to observe moments at which the
first voltage derivative, in relation to time,
equals zero. Using the measurement of time
t0 from the moment at which sinusoidal current passes through zero until the moment of
arc ignition or termination it is possible to
determine the whole time constant using the
following formula [21]:
the time constant is carried out at point i = 0,
when conductance g has an indefinite value,
the value of the time constant in the expression (13) is calculated from the following interpolation:
g=
g (t0 − ∆t ) + g (t0 + ∆t )
2
(15)
where t0 – time instant in which i = 0 A; Δt –
time interval, at which the recording of current and voltage values takes place.
According to another method, net current
does not have to pass through zero. In such
a case the value of power is determined ust0
(11)
θ=
ing the formula (14), and the time constant is
2
calculated from the dependence below:
Another simple method uses the harmonic

g  i2
(16)
analysis of arc voltage [26]

θ =−
− 1
dg  gPM


1 1
dt
 − χ 
θ=
(12)
4ω  χ

It is also possible to determine the damping
where χ = A2n+1 / A2n-1 < 1 amplitudes of the factor function on the basis of the reaction of
closest harmonic odds of voltage (usu. χ = the arc column on various length disturbances. They should also be appropriately synA3 / A1 ).
A special method for testing the AC arc chronised and shifted in the phase in relation
consists in “placing” a properly selected to the course of current. In laboratory condihigh-frequency current (as to the amplitude tions the changes of length can be relatively
and phase) on the current flowing through the easy to induce by means of properly selected
arc [7]. For the purpose of testing air switches rotating electrodes (commutators) [7]. The
this frequency usually amounts to 20 kHz. In excitation of high-frequency disturbances by
the case of switches with elgas the frequency electrode vibrations is more difficult, espeis much higher and equals 70 kHz. In this cially if electrodes are massive. It also facilimanner one can trigger additional transition tates the sputtering of electrode material and
processes in the areas of net current passage increases electrode erosion. A relatively high
through zero. After registering the courses, frequency of such changes can be obtained
the parameters of the Mayr model [7] can be by the crosswise action of variable magnetic
field on the arc [27]. Due to some binding of
calculated from the following dependence:
the column to electrode spots (especially of
= −
, if = 0
(13) the cathode) only the central part of a long
arc is the preferable area of this action.
In an electric arc with stabilised current
(14) it is possible to trigger voltage changes by
= , if
=0
modifying the conditions of heat exchange
where PM – constant value of the dissipated with the surroundings [21]. The longitudinal
power of the model. As the determination of pulse flow of gas around the column caus42
No. 2/2013
es momentary changes of dissipated power
and of the arc column diameter. In turn, the
transverse or slant pulse flow of gas around
the column causes its temporary elongation
and contraction. The changing motion of the
arc in relation to the gaseous environment
can also be triggered by means of a modulated magnetic field properly synchronised
with the course of discharge current. The use
of parallel ring electrodes with moving arc
spots enables maintaining almost the whole
length of the column.
Artificially introduced arc disturbances
should be characterised by a limited range
of amplitude due to the strong non-linearity of static and dynamic characteristics as
well as because of discharge instability. For
this reason, the depth of modulation usually
amounts to a few percent. Too high an amplitude of current disturbances changes the
character of arc discharge from the “dc-” to
“ac-type” [18]. Also, too high a frequency
of current disturbances changes the character of discharge from the ac-type arch discharge with thermal plasma to the “RF-type
discharge” with non-equilibrium plasma. For
this reason there should be an inverse proportionality between the amplitude of periodic disturbance and its frequency.
It is also technically possible to carry out
synchronised disturbance of the arc column
with two or more types of external factors
at the same time. In this manner one can
obtain the deepened modulation of courses,
which however may facilitate the occurrence
of discharge instability. Due to this fact such
solutions are not applied for testing electrotechnological devices. In addition, the greater complexity of the design and operation of
the testing station entail the greater complexity of necessary analyses and measurements.
No. 2/2013
Such inputs cannot be compensated by the
improved accuracy of obtained results. In
such situation, the methods introducing
electric disturbances are the simplest, most
accurate and, consequently, most popular
[6, 15, 16, 18, 28].
The second part of the article focuses on
the assessment of the usability of methods
used for measurements of dynamic characteristics by simulating processes in circuits
with modified and hybrid models of the electric arc.
Conclusions:
1. The results of the so-far experimentation and theoretical analysis of such physical
quantities as the damping factor function and
arc geometrical dimensions often do not confirm adopted assumptions, nor do they offer
the possibility of obtaining simple and direct
relationships between θ and the diameter da
or the cross sectional area S of the column.
2. Most of the experimental methods for
determining the dynamic characteristics
of the electric arc enable only the determination of the time constant in the areas of
current decay. The constant constitutes the
maximum value of damping function and, as
such, is predominantly useful for modelling
the joining arc.
3. The pursuit of more and more precise
reproduction of processes in the circuits of
welding and electrothermal equipment with
an electric arc causes the popularisation of
complex hybrid models increasing the usability of non-linear damping functions.
The research work has been financed
from the funds for science
in 2010-2013 as research project
no. N N511 305038.
43
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дународная конференция «Обратные
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