Conference Proceedings

Transcription

NEW METHODS OF DAMAGE
AND FAILURE ANALYSIS OF
STRUCTURAL PARTS
September 8 – 11, 2014
Ostrava
Czech Republic
NEW METHODS OF DAMAGE AND FAILURE ANALYSIS
OF STRUCTURAL PARTS
Book of Abstracts
6th International Conference
September 8 – 11, 2014 Ostrava
edited by
Prof. Ing. Bohumír Strnadel, DrSc.
VŠB – Technical University of Ostrava
17. listopadu 15/2172
708 33 Ostrava
Czech Republic
Vydavatelství Vysoká škola báňská – Technická univerzita Ostrava
Published in cooperation with the project Regional Materials Science and Technology Centre,
CZ.1.05/2.1.00/01.0040.
i
Book of Abstracts
6th International Conference on NEW METHODS OF DAMAGE AND FAILURE ANALYSIS
OF STRUCTURAL PARTS.
VŠB - Technical University of Ostrava, Czech Republic
8 - 11 September 2014
Published by Vydavatelství VŠB – TU Ostrava
17. listopadu 15, 708 33 Ostrava
Czech Republic
Cover image: Corrosion products on fracture surface of steel AISI 304 after stress
corrosion cracking test, by Doc. Ing. Jan Siegl, CSc., image magnification of 1000x.
Proceedings were designed by
Ing. Pavel Židlík
Ing. Daniela Vedrová
ISBN 978-80-248-3488-7
ii
CONTENTS
Proposal of Proximity Rule on Transformation from Subsurface to Surface Flaw .................. 1
K. HASEGAWA, Y. LI, R. SERIZAWA, M. KIKUCHI
Improvement of the Herbert Pendulum Hardness Tester .......................................................... 3
T. KABURAGI, R. SUZUKI, M. MATSUBARA, T. KOYAMA, T. TASHIRO
Decreasing Thermal Stresses in Steam Generator Collector Weld’s Area Using
External Cooling ....................................................................................................................... 5
R. KRAUTSCHNEIDER, L. JOCH
Collapse Mechanism of Rectangular Tubes Subjected to Pure Bending .................................. 7
K. MASUDA
The Evaluation of Actual Material Properties of Low Alloy CrMoV Steel from the
Results of Small Punch Tests .................................................................................................... 9
K. MATOCHA, L. KANDER, O. DORAZIL, K. GUAN, Y. XU
Extra High Strength Steel Plates. Production, Possibilities and Limits .................................. 11
I. MIKA
Precision Evaluation for Realiability of Power Module Using Coupled ElectricalThermal-Mechanical-Analysis ................................................................................................ 13
H. MORITA, Q. YU
Comparative Strain Analysis of 34CrMo4 Steel and Inconel 738LC ..................................... 15
M. ŠTAMBORSKÁ, M. LOSERTOVÁ, K. KONEČNÁ, V. MAREŠ, L. HORSÁK, R. GALACZ
Internal Crack Growth Simulation Using S-version FEM ...................................................... 17
M. KIKUCHI, R. SERIZAWA, S. YAMADA
Collapse Evaluation of Double Notched Stainless Pipes Subjected to Combined
Tension and Bending ............................................................................................................... 19
R. SUZUKI, M. MATSUBARA, S. YANAGIHARA, M. MORIJIRI, A. OMORI, T. WAKAI
Development of a Crack Opening Displacement Assessment Procedure Considering
Change of Compliance at a Crack Part in Thin Wall Pipes Made of Modified
9Cr-1Mo Steel ......................................................................................................................... 21
T. WAKAI, H. MACHIDA, M. ARAKAWA, S. YOSHIDA, S. YANAGIHARA, R. SUZUKI,
M. MATSUBARA, Y. ENUMA
Stress Relaxation Small Punch Testing of P92 Steel .............................................................. 23
P. DYMÁČEK, F. DOBEŠ, M. JEČMÍNKA
ABI Testing of Reactor Pressure Vessel Steel ........................................................................ 25
P. HAUŠILD, J. SIEGL, A. MATERNA
Characterization of Heterogeneous Weldments ...................................................................... 27
V. ŠEFL, R. NOVÁKOVÁ, J. BYSTRIANSKÝ
iii
Resistance to Corrosion Cracking of Steel 100Cr6 in Humid Air under Higher Tensile
Stresses .................................................................................................................................... 29
S. LASEK, V. ČÍHAL, M. BLAHETOVÁ, E. KALABISOVÁ
Relationships between KIC and CVN at Temperatures Lower than NDT............................... 31
M. HABASHI, M. TVRDY
Biomechanics - Probabilistic Reliability Assessment of Femoral Screws .............................. 33
K. FRYDRÝŠEK
Finite Element Human Model for Crash safety Assessment of Automobile in Frontal
Collision .................................................................................................................................. 35
Y. ZAMA
Biomechanics – Problematic Loosening of Locking Screws from Plates .............................. 37
R. ČADA, K. FRYDRÝŠEK
Biomechanics – Safety Factor Evaluation of Anterolateral Plates for Distal Tibia
Fractures .................................................................................................................................. 39
G. THEISZ, K. FRYDRÝŠEK
Cyclic Bending Deformation and Fracture of Al and Al-1.0mass%Mg Alloy ....................... 41
H. IKEYA, H. FUKUTOMI
Cyclic Instability of Steel-Titanium Bimetallic Composite Obtained by Explosive
Welding ................................................................................................................................... 43
A. KAROLCZUK, T. ŁAGODA
Including of Ratio of Fatigue Limits from Bending and Torsion for Estimation
Fatigue Life under Cyclic Loading ......................................................................................... 45
M. KUREK, T. ŁAGODA
Evaluation of Fatigue Crack Growth in Alpha Titanium Alloys ............................................ 47
O. UMEZAWA, M. HAMADA, T. TATSUMI
The Current Status of New Czech Corrosion Fatigue Evaluation Proposal for WWER
Nuclear Power Plants .............................................................................................................. 49
L. VLČEK
Effect of Repeated Heating on One-Point Rolling Contact Fatigue of High-Carbon
High-Chromium Steel Bar ...................................................................................................... 51
K. MIZOBE, R. SEGAWA, T. SHIBUKAWA, K. KIDA
Statistical Analysis of Accidents Due to Fatigue and Corrosion at Facilities Producing
High Pressure Gas ................................................................................................................... 53
T. SHIBUTANI, N. KASAI, H. KOBAYASHI, H. AKATSUKA, T. TAKAHASHI, T. YAMADA
Influence of Inductive Hardening on Wear Resistance in Case of Rolling Contact ............... 55
M. ŠOFER, R. FAJKOŠ, R. HALAMA
Effect of Surface Quality of Machined Railway Wheels on Fatigue Strength ....................... 57
R. FAJKOŠ, T. TKÁČ
iv
Application of Ultrasonic Impact Treatment (UIT) for Improvement of Fatigue Life ........... 59
T. ISHIKAWA, K. HAYASHI
Cyclic Plastic Properties of Class C Steel Including Ratcheting: Testing and
Modelling ................................................................................................................................ 61
R. HALAMA, A. MARKOPOULOS, M. ŠOFER, P. MATUŠEK
Characterization of Vermiculite Particles after Mechanical Treatment .................................. 63
K. ČECH BARABASZOVÁ, G. SIMHA MARTYNKOVÁ
Advanced Numerical Modelling Methods for 1D Periodic Plasmonic Structure
Simmulations........................................................................................................................... 65
L. HALAGAČKA, K. POSTAVA, M. VANWOLLEGHEM, B. DAGENS, J. BEN YOUSSEF,
J. PIŠTORA
Molecular Modeling of Antimicrobial Nanocomposites ........................................................ 67
D. HLAVÁČ, J. TOKARSKÝ
Antimicrobial Kaolinite Based Nanocomposites .................................................................... 69
S. HOLEŠOVÁ, M. HUNDÁKOVÁ, E. PAZDZIORA
Volatile Organic Molecules Sorption onto Carbon Nanotubes ............................................... 71
G. SIMHA MARTYNKOVÁ, D. PLACHÁ, L. ROZUMOVÁ, E. PLEVOVÁ
Optical Modelling of Microcrystalline Silicon Deposited by Plasma-Enhanced
Chemical Vapour Deposition on Low-Cost Iron-Nickel Substrate for Photo-Voltaic
Applications ............................................................................................................................ 73
Z. MRÁZKOVÁ, K. POSTAVA, A. TORRES-RIOS, M. FOLDYNA,
P. ROCA I CABARROCAS, V. VODÁREK, J. HOLEŠÍNSKÝ, J. PIŠTORA
Submicron Calcium Phosphate Particles Study Anchored on Clay Supports ......................... 75
L. PAZOURKOVÁ, G.SIMHA MARTYNKOVÁ, M. HUNDÁKOVÁ, M. VALÁŠKOVÁ
Preparation of Submicron Particles of Biologically Active Substances Using
Supercritical Fluids ................................................................................................................. 77
D. PLACHÁ, T. SOSNA, E. VACULÍKOVÁ, M. MIKESKA, R. DVORSKÝ
Preparation of Carbon Nano Fillers for Metalic Composites .................................................. 79
L. ROZUMOVÁ, G. SIMHA MARTYNKOVÁ
TiO2 – Based Sorbent for Lead Ions Removal ........................................................................ 81
J. SEIDLEROVÁ, M. ŠAFAŘÍKOVÁ, L. ROZUMOVÁ, I. ŠAFAŘÍK, O. MOTYKA
Properties of Kaolinite Treated by Different Temperatures.................................................... 83
M. TOKARČÍKOVÁ, K. MAMULOVÁ KUTLÁKOVÁ, J. SEIDLEROVÁ
Influence of Void on the Mechanical Property of Nanomaterial ............................................ 85
K. YODEN, Y. SAITO, Q. YU
Determination of Anisotropic Crystal Optical Properties Using Mueller Matrix
Spectroscopic Ellipsometry ..................................................................................................... 87
K. POSTAVA, R. SÝKORA, D. LEGUT, J. PIŠTORA
v
Observation of Magnetic Fields around Plastic Deformation Area in Low Carbon
Alloy Steel ............................................................................................................................... 89
K. KIDA, M. ISHIDA, K. MIZOBE
Magneto-Plasmonic Properties of Au/Fe/Au Planar Nanostructures: Theory and
Experiments ............................................................................................................................. 91
J. VLČEK, M. LESŇÁK, P. OTIPKA
Image Analysys via Reaction Diffusion System for Edge Detection ..................................... 93
K. NAKANE, H. MAHARA, K. KIDA
Effects of Inclusion on the In-Plane Mechanical Performance of Micro-Lattice Plate .......... 95
K. USHIJIMA, W. J. CANTWELL, D. H. CHEN
Assessment of Structures Loaded at Creep ............................................................................. 97
S. VEJVODA, P. POPPELKA, P. RYŠAVÝ
Diffusion of Hydrogen in the TRIP 800 Steel......................................................................... 99
J. SOJKA, P. VÁŇOVÁ, V. VODÁREK, M. SOZANSKA
Precipitation Reactions in a Copper - Bearing GOES........................................................... 101
V. VODÁREK, A. VOLODARSKAJA, Š. MIKLUŠOVÁ, J. HOLEŠINSKÝ, O. ŽÁČEK
Concept of Damage Monitoring after Grinding for Components of Variable Hardness ...... 103
A. MIČIETOVÁ, J. PIŠTORA, Z. DURSTOVÁ, M. NESLUŠAN
Study on Reliability Evaluation Method ofAdhesion Strength of Resin .............................. 105
O. HONDA, Q. YU
Direct Bonding of Ti/Al by Metal Salt Generation Bonding Technique with Formic
Acid ....................................................................................................................................... 107
T. AKIYAMA, S. KOYAMA
Direct Bonding of SUS304 Stainless Steel by Metal Salt Generation Bonding
Technique with Formic Acid................................................................................................. 109
T. TSUNETO, S. KOYAMA
Direct Bonding of A6061 Aluminum Alloy by Metal Salt Generation Bonding
Technique with Formic Acid................................................................................................. 111
Y. TOMIKAWA, S. KOYAMA
Effect of Surface Modification by Aqueous NaOH Solution on Bond Strength of
A5052 Aluminum Alloy/Al and Cu/Al ................................................................................. 113
X. MA, S. KOYAMA
Direct Bonding of Cu/Cu by Metal Salt Generation Bonding Technique with Formic
Acid and Acetic Acid ............................................................................................................ 115
S. KOYAMA, N. HAGIWARA, I. SHOHJI
Delamination Property of Modelled Air Plasma Sprayed Therma Barriear Coatings:
Effect of Difference in Chemical Composition of Bond Coat .............................................. 117
M. HASEGAWA, S. YAMAOKA
vi
Microstructure Modification of CGDS and HVOF Sprayed CoNiCrAlY Bond Coat
Remelted by Electron Beam .................................................................................................. 119
P. GAVENDOVÁ, J. ČÍŽEK, J. ČUPERA, M. HASEGAWA, I. DLOUHÝ
Response of Alumina Foam to Tensile Mechanical Loading Including Stress
Concentrator Effect ............................................................................................................... 121
I. DLOUHÝ, Z. CHLUP, H. HADRABA, L. ŘEHOŘEK
Nondestructive Magnetic Monitoring of Grinding Damage ................................................. 123
M. ČILLIKOVÁ, B. MIČIETA, M. NESLUŠAN, D. BLAŽEK
vii
viii
NEW METHODS OF DAMAGE AND FAILURE ANALYSIS OF STRUCTURAL PARTS
8 – 11, SEPTEMBER, 2014, OSTRAVA, CZECH REPUBLIC
PROPOSAL OF PROXIMITY RULE ON TRANSFORMATION FROM
SUBSURFACE TO SURFACE FLAW
K. HASEGAWA1*, Y. LI1, R. SERIZAWA2, M. KIKUCHI2
1
Japan Atomic Energy Agency, Tokai-mura, Ibaraki-ken, 319-1195, Japan;
email: hasegawa.kunio@jaea.go.jp
2
Tokyo University of Science, Yamazaki, Noda-shi, Chiba-ken, 278-8510, Japan
KEY WORDS: proximity rule, subsurface flaw, fatigue crack growth, stress intensity factor
If subsurface flaws are detected that are close to the component free surfaces, flaw-tosurface proximity rule is used to determine whether the flaws should be treated as subsurface
flaws as is, or transformed to surface flaws. However, the criteria for the rules on transforming
subsurface to surface flaws differ among fitness-for-service codes.
A subsurface flaw located near a component surface is illustrated in Fig. 1, where a is the
half flaw depth,  the length, and S the distance from subsurface flaw to component surface.
When S is short, the subsurface flaw is transformed to be a surface flaw with the depth of 2a+S.
ASME [1], JSME [2], and Swedish SSM[3] provide the proximity rules as follow;
Y  S / a  0.4 .
(1)
When S and a satisfy Eq. (1), the subsurface flaw is treated as a surface flaw, where Y is
the flaw-to-surface proximity factor.
It is reported that, from fatigue crack growth
experiments, the proximity factor is not constant, as shown
in Eq. (1). It is suggested that the Y should be the function
of the flaw aspect ratio of a/ based on the experimental
data [4]. This is because the stress intensity factor (SIF) at
point 1 in Fig. 1 increases with decreasing the aspect ratio
a/ under constant S. Large interaction between the crack
tip at point 1 and component free surface occurs for small
aspect ratio of a/.
2a
Point 1, K1
S

Fig. 1 Subsurface flaw near component
surface.
1.4
a/=0.125
a/=0.250
1.3
a/=0.375
K1/K2
The SIFs at points 1 and 2 for
subsurface flaws with various shapes in
plates were calculated by FEM analysis.
The applied load was membrane stress. The
relationship between the ratio of SIF at
points 1 and 2, K1/K2, and the distance S is
shown in Fig. 2, as a parameter of a/. The
K1/K2 increases with decreasing the
distance S and a/. When looking at the
same value of K1/K2, interaction of flaw
with small aspect ratio occurs at long
distance. For example, in case of
K1/K2 =1.1, flaw with a/ = 0.125 occurs at
S =2.3 mm, and flaw with a/ = 0.65 occurs
at S =1.0 mm. That is, smaller the aspect
ratio, longer the distance.
Point 2, K2
a/=0.500
1.2
a/=0.625
1.1
1
0
3
4
5
S, mm
Fig. 2 Stress intencity factor ratio for the distance of
subsurface flaw.
1
1
2
Figure 3 shows the constant interaction for distance and flaw aspect ratio. The exponent of
4.83 for K1/K2 came from fatigue crack growth rates [4]. The fatigue crack growth rate
da/dN =2.01×10-14K4.83 was obtained by 0.5CT (compact tension) specimens, where K is
the SIF range. To compare the behaviour of subsurface to surface flaw by fatigue growth,
equivalent interaction of (K1/K2 )4.83 is shown between the distance S/a and the aspect ratio a/.
Again, interaction for small a/ occurs at long distance. The fatigue test data on abrupt changes
from subsurface to surface flaw depths are shown as open circles in Fig. 3 [4]. It can be seen
that the test data are close to the curve of (K1/K2 )4.83 =1.3. Subsurface crack begins to penetrate
free surface when the crack growth rate at point 1 is 30% higher than that at point 2.
1
1
Experiment
0.8
0.8
4.83
=1.2
4.83
0.6
0.4
S/a (=Y)
S/a (=Y)
(K1/K2)
1.3
0.2
(K1/K2)
0.6
0.4
1.3
Y = 0.8-(a/)
0.2
1.5
=1.3
2.0
0
0
0.2
0.4
Aspect ratio, a/
0.6
0.8
Fig. 3. Stress intencity factor ratios for subsurface flaw.
0
0
0.2
0.4
0.6
Aspect ratio, a/
0.8
Fig. 4. Proposal of proximity rule.
From the view point of codification, it is desired to be a simple expression. Based on the
curve of (K1/K2 )4.83 =1.3 and experimental data, a new proposal of proximity rule is given by;
Y = 0.8 – (a/) for 0 < a/  0.6, and
Y = 0.2
for 0.6 < a/.
(2)
It is concluded that the new proximity factor Y based on SIF interaction can be developed
as a function of aspect ratio.
Acknowledgement: The authors gratefully acknowledge the support by K. Saito, Hitachi
GE, and K. Miyazaki, Hitachi Ltd.
REFERENCES
[1]
[2]
[3]
[4]
American Society of Mechanical Engineers Boiler & Pressure Vessel Code Section XI: Rules
for In-service Inspection of Nuclear Power Plant, 2013 Edition.
The Japan Society of Mechanical Engineers S NA1: Rules on Fitness-for-Service for Nuclear
Power Plants (in Japanese), 2008.
Swedish Radiation Safety Authority (SSM): A Combined Deterministic and Probabilistic
Procedure for Safety Assessment of Components with Cracks-Handbook, 2008.
HASEGAWA, K., LI, Y., MIYAZAKI, K. SAITO, K.: Fatigue Crack Growth for Subsurface
Flaws near Component Surface and Proximity Rules, ASME PVP2013-97559, Paris, 2013.
2
IMPROVEMENT OF THE HERBERT PENDULUM HARDNESS TESTER
T. KABURAGI1, R. SUZUKI2*, M. MATSUBARA2, T. KOYAMA2, T. TASHIRO3
1
Gunma Industrial Technology Center, 884-1 Kamesato-machi, Maebashi, Gunma 379-2147, Japan
Faculty of Science and Technology, Gunma University, 1-5-1 Tenjin-cho, Kiryu, Gunma 376-8515, Japan;
email: r_suzuki@gunma-u.ac.jp
3
Department of mechanical system engineering, Faculy of engineering, Gunma University, 1-5-1 Tenjin-cho,
Kiryu, Gunma 376-8515, Japan
2
KEY WORDS: hardness, damping property, pendulum, measurement system
The Herbert hardness tester [1-3] is a typical pendulum-type hardness tester. Hardness of
materials is measured based on the swing angle of the pendulum in relation to the specimen. In
the present study, hardness is measured using a Habara-type Herbert pendulum hardness tester
[4] with a modified measurement system. We investigate the effect of each condition such as
the indenter tip radius of the Herbert pendulum, the swing cycle and the surface roughness of
the specimens on Herbert hardness. Moreover, we discuss the relationship between damping
hardness and conventional Brinell hardness.
Fig. 1 shows the measurement system for
Herbert hardness. The Herbert pendulum
swings on a specimen and its swing angle is
measured continuously with the two laser
displacement meters which are installed
independent from the Herbert pendulum.
Laser displacement meter
Weight
S (t )  S0et
3
S1
S2
Swing angle [degrees]
Fig. 2 shows the swing angle of the
Weight
Indenter
Herbert pendulum as a function of time. The Weight
Specimen
angular amplitude of the Herbert pendulum
decreases exponentially with time. The
original definitions of Herbert hardness is
evaluated by initial swing angle, S1, and the
Fig. 1. Measurement system.
time it takes for 10 swings, T, because the
swing time is measured using a stopwatch
40X
S t   S 0 e t
S0
and the swing angle is measured using a spirit
level by the visual evaluation. Therefore
Herbert hardness evaluated by original
definitions has a large variation. So we
0
improved the measurement system in order to
0
120t
measure the swing angle accurately more
than original measurement system. In
addition, in place of the original definitions,
T
damping hardness is proposed as an
-40
Time [s]
indication of hardness. Since damping
Fig. 2. Characteristics of the swing angle of the tester as
hardness is determined by the damping factor a function of time obtained by Herbert hardness tester.
obtained from the free damped vibration
waveform of the Herbert pendulum, its method of hardness evaluation more reasonable than
the original definitions. The following equation shows the envelope line connecting successive
local maximum amplitudes S(t).
(1)
The exponent α of Equation (1) is referred to as the damping factor and indicates the amount
of damping in the testing system.
Fig. 3 shows the relationship between the damping hardness and conventional Brinell
hardness. The dashed line is predicted value of the Brinall hardness from the damping hardness
and the indenter radius. For all indenter radii, the damping hardness increases with the decrease
in the Brinell hardness number. For a sample of the same Brinell hardness number, the damping
hardness increased with increasing indenter radius. The difference in the damping hardness
associated with the indenter radius increases as the Brinell hardness number decreases. The
increase in the contact area between the indenter and the specimen associated with the increase
in the indenter radius and the decrease in the hardness of the specimen, the resistance to
swinging of the pendulum tester increased, resulting in large damping.
Damping hardness
0.004
R1
R2
R4
累乗
(R2) value
Predicted
0.003
0.002
0.001
0
0
200
400
Brinell hardness
600
800
Fig. 3. Brinell hardness predicted based on the damping hardness.
Here, in order to predict the Brinell hardness number the multivariate analysis is conducted
using the damping hardness. As a result, we obtained the following expression:
HBW 
R 0.448
 1.131
e 0.946 ,
(2)
where HBW is Brinell hardness number, α is the damping hardness and R is the indenter radius.
The correlation coefficient exceeds 0.98. Therefore, it may be possible to accurately predict the
Brinell hardness based on the damping hardness, and this measurement system may be used
practically.
REFERENCES
[1]
[2]
[3]
[4]
HERBERT, E.G.: "Some Recent Developments in Hardness Testing," The Engineer 135:686-68
(1923).
BENEDICKS, C., CHRISTIANSEN: "Investigations on the Herbert Pendulum Hardness
Tester," Journal Iron and Steel Institute 110; 219-248 (1924).
WILLIAMS, S.R.: Hardness and Hardness Measurements, American Society for Metals,
Cleveland (1942).
HABARA, H., KAWAMITSU, T., HARIMOTO, K., AND INOUE, H.: "Restration of the
Herbert Pendulum Hardness Tester and its application (in Japanese)," Journal of Material
Testing Research Association of Japan, 43(4):248-254 (1998).
4
DECREASING THERMAL STRESSES IN STEAM GENERATOR
COLLECTOR WELD’S AREA USING EXTERNAL COOLING
R. KRAUTSCHNEIDER1*, L. JOCH1
1
Institute of Applied Mechanics Brno, Ltd., Resslova 972/3, 602 00 Brno; email: krautschneiderr@uam.cz
KEY WORDS: steam generator, thermal stresses, external cooling, stress corrosion cracking, dissimilar
metal weld
Presented paper deals with possibility of external cooling of steam generator weld’s area to
decrease internal tensile thermal stresses which are one of the causes of cracking in this area.
On various WWER-440 nuclear power plants (NPP) steam generators (SG) occurred quite
serious problem of cracking in weld joints connecting primary collectors to the SG vessel’s
nozzle (Fig. 1). The cause of this cracking is stress corrosion cracking (SCC) mechanism. The
crack rises and grows on the interface between different kinds of material, austenitic steel
08CH18N10T on one side and carbon steel 22K on the other (Fig. 1).
On the SG secondary side there is a space between the SG nozzle and the primary collector,
which is also called “pocket”. In this pocket, due to poor possibilities of effective blowdown,
exist the secondary media of higher corrosive potential. The existence of corrosive media
together with existing stresses can cause intergranular corrosion and cracking in this area.
Existing stresses are combination of thermal stresses from different thermal expansion
properties of austenitic and carbon steel, and external stresses on the SG nozzle from the
primary circuit. Thermal stresses are higher and thus more important.
So there are two approaches how to decrease the possibility of crack occurrence and growth.
The first one is to try keeping the pocket as clean as possible, it means to try to improve the
effectiveness of the pocket’s blowdown. And the second one is to try to decrease existing
(thermal) stresses. Of course both approaches can be combined.
SG
vessel
(22K)
SG nozzle
(22K)
22K
SG collector
(08CH18N10
T)
Dissimilar
metal weld
Fig. 1. PGV-440 steam generator collector’s weld area.
Recently there were presented some studies of external cooling of this area from Russia.
Those studies were done on WWER-1000 steam generators, but from similar reason. And
5
because the SG primary collector’s joint is similar on WWER-440, we tried to make similar
analyses here.
The external cooling of the SG primary collector weld joint is done by placing external
cooing sleeve around it, with inlet and outlet nozzle and with flowing air cooling media of
higher velocities.
Computational fluid dynamics (CFD) analyses and subsequent finite element analyses
(FEM) were done, with the objective to compare stresses in the weld area with and without
external cooling.
Results of these analyses showed that by proper external cooling of the weld area it is
possible not just to decrease significantly existing tensile stresses, but to change them into
compressive ones. This of course would have great impact on crack occurrence and growth.
REFERENCES
[1]
[2]
[3]
[4]
[5]
JOCH, L.: Influence of thermal fields on dissimilar weld DN1105 of NPP Dukovany steam
generator’s collector (in Czech language), IAM Brno Report, 5071/13, Brno, 2013.
LICKA, A.: Determination of residual life of SG primary collectors with graphite gasket. The
evaluation of the measurement results after running the second block and in steady state
operation at nominal power, including recommendations for further operation (in Czech
language), IAM Brno Report, 2643/98, Brno, 1998.
KUTDUSOV, YU.F. et al.: Innovation of devices for stress reduction in welding joint 111 of
welding unit of “hot” heat-transfer manifold and socket DN1200 of PGV-1000 by air blowing
method on Rostov and Balakovo nuclear power plants (in Russian language), 8-th International
scientific and technical conference "safety assurance of NPP with WWER", OKB Gidropress,
ISBN: 978-5-94883-130-5, Podolsk, 2013.
LYAKISHEV, S.L. et al.: Development and justification of measures on assurance of reliable
and safe operation of welded joints no. 111 of steam generator PGV-1000M (in Russian
language), 6-th International scientific and technical conference "safety assurance of NPP with
WWER", OKB Gidropress, Podolsk, 2009.
TRUNOV N.B. et al.: Results of studies of metal fracture causes in the area of primary
collector-to-steam generator vessel welding and development of corrective measures (in
Russian language), 8th International Seminar on Horizontal Steam Generators, OKB
Gidropress, Podolsk, 2010.
6
COLLAPSE MECHANISM OF RECTANGULAR TUBES
SUBJECTED TO PURE BENDING
K. MASUDA1*
1
University of Toyama, Japan; email: masuda@eng.u-toyama.ac.jp
KEY WORDS: FEM, pure bending, rectangular tube, buckling, effective width
Rectangular and square section tubes are widely used in mechanical equipment. Therefore,
a study of the collapse behaviour is important for both the design and analysis of weightefficient safety structures. In the present paper, the collapse behaviours of rectangular tubes
subjected to pure bending are investigated using the finite element method. Such bending
collapse has been investigated extensively [1], [2]. These studies have revealed the existence of
two types of collapse. The first type is a collapse due to buckling at the compression flange,
and the second type is a collapse due to plastic yielding at the flanges. Moreover, another type
of collapse exists. For a rectangular tube in which the web is wider than the flange, collapse
due to buckling occurs at the compression web [3]. In the present paper, previous evaluation
method of the maximum moment is refined, and systemized evaluation method is proposed.
The validity of this method is verified through comparison with the numerical results obtained
by FEM under various conditions.
REFERENCES
[1]
[2]
[3]
KECMAN, D.: Bending collapse of rectangular and square section tubes, International Journal
of Mechanical Sciences, Vol. 25, 1983, pp. 623-36.
LU, G., YU, T. X.: Energy absorption of structures and materials, CRC Press, Section 5, 2003.
MASUDA, K., CHEN, D.H.: Prediction of Maximum Moment of Rectangular Tubes Subjected
to Pure Bending, Journal of Environment and Engineering, Vol. 6(3), 2011, pp. 554-566.
7
8
THE EVALUATION OF ACTUAL MATERIAL PROPERTIES OF LOW
ALLOY CrMoV STEEL FROM THE RESULTS OF SMALL PUNCH TESTS
K. MATOCHA1*, L. KANDER1, O. DORAZIL1, K. GUAN2, Y. XU2
1
MATERIAL & METALLURGICAL RESEARCH, Ltd., Ostrava, Czech Republic;
email: karel.matocha@mmvyzkum.cz
2
East China University of Science and Technology, Shanghai, China
KEY WORDS: small punch test technique, actual mechanical properties, CEN workshop agreement,
residual lifetime
The Small Punch Test Technique is used, at the present time, to obtain actual tensile,
fracture and creep characteristics necessary for estimation and monitoring of residual lifetime
of critical components of industrial plants [1-3]. CWA 15627 “Small Punch Test Method for
Metallic Materials” was developed in Europe in 2007 [4]. Part B: A Code of Practice for Small
Punch Testing for Tensile and Fracture Behaviour is used for determination of yield and tensile
strength, Ductile Brittle Transition Temperature (DBTT) and fracture toughness of the metallic
materials.
The present paper summarizes the results of the bilateral project in the frame of CzechChinese Scientific and Technological Cooperation focused on the determination and
comparison of the empirical correlations for estimation of FATT and yield strength, tensile
strength and JIC at laboratory temperature from the results of Small Punch Tests for low alloy
14MoV6-3 steel.
The participants of the project were MATERIAL & METALLURGICAL RESEARCH,
Ltd. Ostrava, Czech Rep. and East China University of Science and Technology, Shanghai,
China. Both tensile tests, Charpy impact tests and fracture toughness tests using standardized
test specimens and Small punch tests were carried out in both laboratories according to national
standards.
Results of standardized tensile tests and impact tests obtained in both laboratories are in
very good agreement. However the empirical correlations for determination of yield stress,
tensile strength and FATT from the results of Small punch tests are significantly different.
Factors affecting this experimentally proved fact are discussed.
Acknowledgement: The paper has originated during the solution of the project LH 12199
„The Comparison of Codes of Practice for Determination of Mechanical Properties by SP Tests
between EU and China“ in the frame of the programme of the Ministry of Education, Youth and
Sports KONTAKT II.
REFERENCES
[1]
[2]
MATOCHA, K.: Determination of Actual Tensile and Fracture Characteristics of Critical
Components of Industrial Plants under Long term Operation by SPT. Proceedings of the ASME
2012 Pressure & Piping Division Conference PVP 2012, July 15-19, 2012, Toronto, Ontario,
Canada (CD-ROMM).
FOULDS, J.R., JEWETT, C.W., BISBEE, L.H., WHICKER, G.A., VISWANATHAN, R.:
Miniature Sample Removal and Small Punch Testing for In-Service Component FATT.
Proceedings of the Robert I. Jaffee memorial Symposium on Clean Materials Technology.
ASM/TMS Materials Week, 2-5 November 1992, Chicago, Illinois, USA.
9
[3]
[4]
ABENDROTH, M.: Identification of Creep properties for P91 Steels at High Temperatures Using
the Small Punch Test. In: Proc. of 1st International Conference “Determination of Mechanical
Properties of Materials by Small Punch and other Miniature Testing Techniques”. Ostrava,
Czech Rep., August 31 to September 2, 2010, pp. 39-43, ISBN 978-80-254-7994-0.
CEN WORKSHOP AGREEMENT “Small Punch Test method for Metallic Materials” CWA
15627:2007 D/E/F, December 2007.
10
EXTRA HIGH STRENGTH STEEL PLATES. PRODUCTION,
POSSIBILITIES AND LIMITS
I. MIKA1*
1
SSAB Swedish Steel s.r.o., Spartakovců 3, 70800 Ostrava, Czech Republic; email: ivan.mika@ssab.com
KEY WORDS: high strength steel, welding, cutting, fatigue, buckling
The present limit in mass produced, high strength, low alloyed steels is approximately
1300 MPa of yield strength, or 650 HBW of hardness respectively. In principle there is not so
big problem to reach needed high strength of the steel. The biggest challenge now, is to assure
the toughness and technological properties in combination with high strength. The
technological properties especially mean welding and thermal cutting. To assure good balance
between strength/hardness, toughness and weldability of steel means to heat treat the steel to
optimal microstructure. This involves reaching martensitic structure with low possible content
of carbon and alloying elements, the same way as to guarantee extremely clean steel. Therefore
it is important, as to quenching process proceeds very fast. Right metallurgical treatment and
proper selection of raw materials has to assure extremely low contents of the elements like
S, P, Sn, Zn. Deep vacuum treatment is subsequently responsible for decreasing especially
hydrogen in liquid steel, which improves thermal cuttability of ready plates.
Nevertheless, practically always, when strength of the steel is increasing, namely more than
700 MPa of yield point, the steel has to be more alloyed and technological properties are
decreasing. Further, there are mechanical properties, like fatigue strength and stability of steel
construction which are practically not increasing with increasing of steel strength. There is
necessary to apply new technological procedures and constructional principles to overcome
these limitations. Very hard steels are also substantially more sensitive to steel corrosion
cracking which make a challenge in some kinds of structures made from these steels.
11
12
PRECISION EVALUATION FOR REALIABILITY OF POWER MODULE
USING COUPLED ELECTRICAL-THERMAL-MECHANICAL-ANALYSIS
H. MORITA1*, Q. YU1
1
Department of Mechanical Engineering, Graduate School of Engineering, Yokohama National University;
Tokiwadai 79-5, Hodogaya-ku, Yokohama, Japan; email:morita-hirotaka-vf@ynu.jp
KEY WORDS: power-modules, electrical-thermal-mechanical-analysis, reliability
In recent years, with the development of power electronics technology, the power modules
have been used extensively. The problem of the power module is thermal fatigue of the junction
area between Si chip and substrate caused by cyclic temperature change. It is necessary to
evaluate the thermal fatigue life of the power module under the power cycle using coupled
electrical-thermal-mechanical analysis.
Power cycle test fixed current and sets up fixed on-off time. This conditions done regularly.
However, irregular the electric load imposed at the time of an actual vehicle running. A gap
exists between the thermal fatigues added by a power cycle test.
It is thought that the reliance valuation basis of a power module is a very high standard
compared with real usage environment. Then, if the evaluation technique in the conditions near
real usage environment is establishable. It will be thought that it becomes possible to cancel the
over-spec of a product and to make the product whose cost was cut down. In this research,
analysis in the conditions near real usage environment is conducted. It aims at establishment of
the reliability assessment method of the power module near a real operating condition.
Analysis model is shown in Fig. 1. Electrical-Thermal-Mechanical-Analysis tried two
pattern as follow.
1st pattern, 200 A current was impressed on
DC IN shown in Fig. 1. The changing time of the
electrical current is 2.0 s. The cooling time is 18 s.
This 1st pattern named is REGULAR LOAD
TEST.
2st pattern, direct current was impressed on DC
IN shown in Fig. 1. The current conditions
impressed on DC IN shown in Table 1. This 2st
pattern named is IRREGULARITY LOAD TEST.
The result of coupled electrical-thermal
analysis is shown in Fig. 2. Temperature
distribution changes by changing a setup of a
current value or On-Off time. This understands that
it sees Fig. 2.
This temperature distribution is applied to
Thermal-Mechanical analysis. The result obtained
by Thermal-Mechanical analysis is creep and
plastic strain. Based on Manson-Coffin’s Law, the
thermal figue life of solder joints is often evaluated
by the inelastic strain range, which is the sum of the
creep and plastic strain [1].
13
Fig. 1. Analysis model of power module.
Table 1 Analysis condition for coupled
Thermal-Electrical Analysis.
Step
1
2
3
4
5
On [s]
4
6
3
2
5
Off [s]
3
20
18
15
10
Current [A]
50
100
200
150
60
One creep and plasitc strain comes out in a regular load test, and Manson-Coffin’s Law,
fatigue lives solder joint is one solution. But two or more solutions come out in an irregular
load test, and Manson-Coffin’s Law, fatigue lives solder joint are tow or more
solutions.Applyin this strain range to Manson-Coffin’s Law shown as following
  
N  1328  

 0.01 
1.43
(1)
,
 means strain range [1].
Miner’s Law was applied to irregular load test result. Miner’s Law shown as following,
n 
D    i  ,
 Ni 
(2)
ni means actual number of cycles, Ni means fatigue lives of cycles [2].
Strain, arranged randamly, rearrranged in each pattern.Arranging for each pattern, each
pattern of fatigue life was known. The sum of adding fatigue life of each pattern is fatigue life
of irregular load.
140
50000
120
40000
Fatigue life[cycle]
temperature[°C]
The result of fatigue life of regular load test and irredular load are shown in Fig. 3.
Examining Fig. 3, fatigue life of irregular load test was 4 times of the fatigue life of regular
load test. This result of analysis was shown that, Load on a power cycle test is over-spac so
than the load applied at the time of actual running.
100
80
60
40
regular load
20
irregularity load
0
0
20
40
time[s]
Fig. 2. Temperature date.
30000
20000
10000
0
60
regular load
irregular load
Fig. 3. Fatigue life.
In this research, analysis in the conditions near real usage environment of thermal fatigue
life of power-modules. The rearranged by classifying each pattern, an irregular condition, to
apply the Manson-Coffin’s Law to each and apply Miner’s Law. By using method of this
research, it is possible by using the simulation also in operation in various patterns, to evaluate
the fatigue life. And in this research, precisione valuation for reality of power module using
coupled electrical-thermal-mechanical-analysis.
REFERENCES
[1]
[2]
YU, Q., SHIRATORI, M.: Fatigue-Stremgh Prediction of Microelectronics Solder Joints Under
Thermal Cyclic Loading, IEEE Trans.Compon.,Packag.Manuf.Techol.,Part A 20(3),
pp. 266-273, 1997.
MINER, M. A.: Cumulative Damage in Fatigue, Journal of Applied Mechanics, v01.12
pp. 159-164, 1945.
14
COMPARATIVE STRAIN ANALYSIS
OF 34CrMo4 STEEL AND INCONEL 738LC
M. ŠTAMBORSKÁ1*, M. LOSERTOVÁ1, K. KONEČNÁ2, V. MAREŠ3, L. HORSÁK3,
R. GALACZ2
1
Department of Non-ferrous Metals, Refining and Recycling, Faculty of Metallurgy and Materials
Engineering, VŠB-Technical University of Ostrava, Czech Republic;
email: stamborska.michaela@gmail.com
2
Department of Material Engineering, Faculty of Metallurgy and Materials Engineering, VŠB-Technical
University of Ostrava, Czech Republic
3
Center of Advanced Innovation Technologies (CAIT), VŠB-Technical University of Ostrava, Czech Republic
KEY WORDS: 34CrMo4, IN 738 LC, stress analysis
The article was focused on strain analysis of 34CrMo4 steel and IN 738 LC superalloy. The
34CrMo4 steel grade is a low-carbon steel with medium through-hardening for medium-duty
machine parts [1]. The INCONEL 738 LC alloy is Ni-based low carbon superalloy hardened
by fine  precipitates and carbides and used for high temperature applications, as gas turbine
engines [2, 3].
Experimental analysis of plastic deformation on the surface of specimens using contactless
displacement sensing methods is advantageous to obtain deformation fields in pre-selected
areas. The digital image correlation (DIC) is one of the most advanced optical methods of
displacement sensing and subsequent determination of strains on the surface of examined
objects [4, 5]. The strain fields were calculated from the displacement fields by the Vic 2D
program for both above mentioned materials.
Tensile test was carried out on the Zwick / Roel Z150 device with deformation rate of
2.5 x10-3 s-1 on cylindrical specimens having the gauge length of 28 mm and the diameter of
5 mm. Evaluation and compilation of the true stress –strain diagrams for all six specimens were
carried out using image correlation software Vic 2D and scanning was performed using Canon
5D MARK II.
Material characteristics and stress – strain diagrams obtained for 34CrMo4 steel and IN 738
LC superalloy from standard uniaxial tensile test are listed in Table 1 and shown in the Fig. 1,
respectively.
Table 1 The values of the measured mechanical properties for 34CrMo4 steel and IN 738 LC superalloy.
34CrMo4
IN 738 LC
Specimens
1
2
3
Average value
1
2
3
Average value
Y.S. [MPA]
937
958
941
94528
725
721
722
723  4
U.T.S. [MPA]
1040
1041 1041
10411
945
864
914
908  79
εx [-]
0.446 0.436 0.390
0.4410.020
0.184 0.139 0.177
0.167  0.024
Fig. 2 shows the results of strain field εx [-] obtained by the VIC 2D software for 34CrMo4
steel. The values of strains for 34CrMo4 steel obtained by Vic 2D, shown in red in the necking
of the specimens, reach maximum values from 39 to 45%. The average values of tensile
characteristics of 34CrMo4 steel reached of 945 MPa and 1041 MPa for the yield strength and
ultimate tensile stress, respectively. The values of strains for IN738LC superalloy obtained by
Vic 2D, reach maximum values from 14 to 18%. In the case of IN 738LC, the average tensile
characteristics reached of 723 MPa and 908 MPa for the yield strength and ultimate tensile
stress, respectively.
15
a)
b)
Fig. 1. The stress-strain diagram for a) 34CrMo4 steel and b) IN 738 LC superalloy.
a)
b)
Fig. 2. The strain fields obtained by Vic 2D for a) 34CrMo4 steel and b) IN 738 LC superalloy.
Based on the results of comparative strain analysis the DIC method is suitable for applying
to both alloys with good plasticity and with higher fragility. The method allows very well
determining the magnitude of the strain fields and localization of the deformation area as well
as to detect the casting defects in the material affecting mechanical behaviour.
Acknowledgement: This article has been elaborated in the framework of the projects:
"Opportunity for young researchers", reg. Nr. CZ.1.07/2.3.00/30.0016, supported by
Operational Programme Education for Competitiveness and co-financed by the European
Social Fund and the state budget of the Czech Republic and "Regional Materials Science and
Technology Centre - Feasibility Program", reg. Nr. LO1203 funded by Ministry of Education,
Youth and Sports of the Czech Republic.
REFERENCES
[1]
[2]
[3]
[4]
[5]
HENDRYCH, A., KVÍČALA, M., MATOLIN, V., ŽIVOTSKÝ, O., JANDAČKA, P.:
International Journal of Fracture, Vol. 168, 2011, No. 2, pp. 259-266,
DOI: 10.1007/s10704-010-9573-7.
BALIKCI, E., MIRSHAMS, R.A., RAMAN, A.: Tensile Strengthening in the Nickel-Base
Superalloy IN738LC. Journal of Materials Engineering and Performance, Vol. 9(3), June 2000,
pp. 324-329.
LOSERTOVÁ, M., KONEČNÁ, K., JUŘICA, J., JONŠTA, P.: Hydrogen Effect on Mechanical
Properties of IN738LC Superalloy, In: 20th Anniversary International Conference on
Metallurgy and Materials: METAL 2011 Metal 2011, pp. 1039-1043. ISBN 978-80-87294-24-6.
ROSSI, M., PIERRON, F., ŠTAMBORSKÁ, M., ŠIMČÁK, F.: Experimental and Applied
Mechanics, Vol. 4, 2013, pp. 229-235, DOI: 10.1007/978-1-4614-4226-4_27.
ŠIMČÁK, F., ŠTAMBORSKÁ, M., HUŇADY, R.: Deformation of materials by using digital
image correlation, CHEMICKE LISTY, Vol. 105, 2011, No. 4, pp. 564-567, ISSN 0009-2770.
16
INTERNAL CRACK GROWTH SIMULATION USING S-VERSION FEM
M. KIKUCHI1*, R. SERIZAWA2, S. YAMADA2
1
2
Tokyo University of Science, 2641 Yamazaki, Noda, Chiba, 278-8510, Japan; email: kik@rs.noda.tus.ac.jp
Graduation School of Tokyo University of Science, 2641 Yamazaki, Noda, Chiba, 278-8510, Japan
KEY WORDS: fatigue, inner crack, slant crack, fully automatic crack growth simulation system,
S-version FEM
In nuclear power plant, there is a proximity rule to evaluate multiple inner cracks [1]. Inner
cracks are generated inside of the structure in different manners. There are many parameters
which affects the growing processes of inner cracks. They are; locations, slant angles, aspect
ratio of each inner crack and distances between adjacent inner cracks. When multiple inner
cracks are detected, proximity rules are proposed. But due to the complexity of the problem, it
is necessary to verify proposed proximity rules. But experimental study is very difficult due to
existence of many parameters, and crack growth occurs inside of the structure. Numerical
simulation is needed for this purpose. This problem is simulated using S-version FEM, which
is fully automatic crack growth simulation system developed by authors [2]. Using S-FEM,
inner crack is modeled independently from global structure, and crack growth is easily
simulated. In maintenance code of nuclear power plant, initial defects are modeled as elliptical
cracks in a normal plane to tension loading direction, and growth rate is estimated on this plane.
But by using S-FEM, realistic defect shape is modeled, and crack growth by fatigue is
simulated. Usually, such small defects are subjected to multi-axial loading, and crack growth
behaviors are very complicated. Parametric studies are conducted for this problem, and
proximity rules are verified with numerical results. Three problems are simulated. One is effect
of initial defect shape, second is crack growth of slant inner crack and the last one is interaction
effect of multiple inner crakcs.
(1) Effect of initial defect shape.
Four initial defects shapes are assumed, as shown in Fig. 1 (a)-(d). Aspect ratios and initial
areas are same for all models. By the proximity rule by JSME [1], all initial defects should be
modeled as an elliptical shape, which is case (d) in these figures. Figure 2 (a)-(b) shows changes
of crack shape of star shape model. Number of cyclic loading is shown. At first, crack growth
occurs mainly at the smaller part along the crack front, and finally, crack shape becomes
circular. Figure 8 shows relations of crack area and number of cycles for all cases. It shows that
all results are very similar to each other. This results show that JSME code gives good
estimation on crack growth prediction.
(a) Star shape
(b) Diamond shape
(c) Arbitrary shape
Fig. 1. Four different initial defects shapes.
17
(d) Ellipse
(a) 2.2x107
(b) 9.3x107
Fig. 2. Changes of crack shape for star shape model.
Fig. 3. Comparison of crack growth rates.
(2) Effect of slant angle of inner crack.
Inner crack is initially generated from inclusions, inhomogeneities or some other initial
defects. At first, it may not grow in a perpendicular plane to principal stresses, and has some
slant angle to principal stress axis. Then initial slant inner crack is assumed and crack growth
process is simulated. By the JSME code, this slant crack is evaluated after it is mapped on a
plane perpendicular to the principal stress axis. In this case, pure mode I crack growth is
assumed. Simulation results verified this modeling is reasonable. Crack growth rate of slant
inner crack is compared with that of JSME code, and it is again verified that modeling by JSME
code is reasonable.
(3) Evaluation of interaction effect of multiple inner cracks.
Growing processes of multiple inner cracks are simulated. Figure 4 (a) and (b) shows
overlapping processes of two parallel inner cracks. Results are compared with JSME code.
Again it is verified that JSME code gives conservative evaluation for the effect of interaction
between multiple cracks.
(a) 1.3 x 107
(b) 2.3 x 107
Fig. 4. Crack growth processes.
REFERENCES
[1]
[2]
JSME S NAl-2004: Codes for Nuclear Power Generation Facilities – Rules on Fitness-forService for Nuclear Power Plants -, (2004), (In Japanese)
KIKUCHI, M., WADA, Y., SHIMIZU, Y., LI, Y.: Crack Growth Analysis in a Weld-heataffected Zone Using S-version FEM, International Journal of Pressure Vessels and Piping,
90-91, (2011), pp. 2-8.
18
COLLAPSE EVALUATION OF DOUBLE NOTCHED STAINLESS PIPES
SUBJECTED TO COMBINED TENSION AND BENDING
R. SUZUKI1*, M. MATSUBARA1, S. YANAGIHARA2, M. MORIJIRI1, A. OMORI2, T. WAKAI3
1*
Faculty of Science and Technology, Gunma University, 1-5-1 Tenjin-cho, Kiryu, Gunma 376-8515, Japan;
email: r_suzuki@gunma-u.ac.jp
2
Department of Mechanical System Engineering, Graduate School of Engineering, Gunma University,
1-5-1 Tenjin-cho, Kiryu, Gunma 376-8515, Japan
3
4002 Narita-cho O-arai Ibaraki 3111393, JAEA, Japan
KEY WORDS: combined load, integrity assessment, plastic collapse, stainless steel, piping
The stainless steel piping is widely used in light water nuclear reactor plants and chemical
plants. By the end of 2011, approximately 30 percent of nuclear power plants in Japan had been
in operation for more than 40 years [1]. Aging of piping in old nuclear power plants is
significant concern to the safe operation of old nuclear power plants. In order to safely operate
old nuclear power plants, integrity assessment of aging piping is important and maintenance
has to be performed as necessary. Structural integrity assessment procedures for reactor
equipment are specified by the Japanese Society of Mechanical Engineers (JSME) and the
American Society of Mechanical Engineers (ASME) [2, 3]. The structural integrity evaluation
of stainless steel pipes cracked due to aging is generally performed using plastic collapse as a
failure criterion because large plastic deformation occurs in the ligament before the plastic
collapse. The plastic collapse point is obtained by the double elastic slope (DES) method and
the double elastic deformation (DED) method. A cracked pipe is typically subjected to
combined axial tension and bending in structural integrity evaluations. A circumferential crack
located in pipe cross section is more detrimental for guillotine break than an axial crack. In
many cases, circumferential cracked pipe can be treated as a single-edge cracked pipe. In this
study, the plastic collapse strength of asymmetry multiple circumferential notched stainless
steel pipes subjected to combined axial tension and bending is
90°
investigated experimentally and is compared with the theoretical
plastic collapse strength. In addition, the potential is discussed for
° 27°25
38
the simplification of structural integrity evaluation of multiple
°
cracked piping.
Schematic illustration of a pipe with asymmetry multiple
circumferential notches is shown in Fig. 1 (a). The notch angles of
the two circumferential notches are 38 and 25 degrees, respectively.
The notch separation angle between the two notches is 27 degree.
The total notch angle that included the notch separation angle is
90 degrees. Schematic illustration of a single notched pipe with
notch angle of 90 degree is shown in Fig. 1 (b). Theoretical plastic
collapse limit curves were calculated based on elastic-perfectly
plastic body to compare with the experimental collapse strength of
the pipe with asymmetry multiple circumferential notches subjected
to combined axial tension and bending.
t
(a) Multiple notches
90°
t
A specimen with 200 mm length was machined from SUS304
steel pipe with 32 mm diameter and 1.5 mm thickness. Two through
(b) Single noth
wall circumferential notches with notch angle 38 and 25 degrees
Fig. 1. Cross sections of the
were cut in the specimen by a wire saw with 320 m diameter. The (a) multiple and (b) single
notch tip radius was 160 m. A schematic illustration of the
notched pipes.
19
Horizontal potention meter
Gonimeter
Hydraulic cylinder for tensile load
Specimen
Displacement gauge
Load cell
Hydraulic cylinder for bending load
Vartical potention meter
Fig. 2. Statically indeterminate fracture mechanics
experimental equipment.
Corrected bending stress, b/y
statically indeterminate fracture mechanics
testing equipment is shown in Fig. 2. This
equipment can apply arbitrary combined
axial tensile and bending loads to the
structural member [4]. The bending moment
is applied to the specimen using four point
bending method. A spring and two steel
round bars (SS400) were placed inside the
pipe in order to prevent local buckling at
transverse load points. The experiments
were carried out for the various load
patterns. The experimental plastic collapse
points were obtained by DES and DED
methods.
3
2.5
The corrected bending stress, b/y, at
2
plastic collapse point is plotted as a function
1.5
of the corrected membrane stress, m/y, in
Fig. 3. Here, b, m and y are bending
1
stress, membrane stress and yield stress of
0.5
Ccollapse limit for multiple
pipe material, respectively. Theoretical
Collapse limit for single
plastic collapse limit curves calculated for
0
DES
the pipes with multiple notches (Fig. 1. (a))
DED
-0.5
and the single notch (Fig. 1. (b)) are also
-0.5
0
0.5
1
1.5
plotted in Fig. 3. All experimental plastic
Corrected membrane stress, m/v
collapse points are over the theoretical
Fig. 3. Statically indeterminate fracture mechanics
collapse limit curve for the pipe with
experimental equipment.
multiple notches. The theoretical plastic
collapse limit curve for the pipe with multiple notches is similar to the single notch. The
integrity of the asymmetry multiple circumferential notched stainless steel pipes subjected to
combined axial tension and bending can be evaluated conservatively using the theoretical
plastic collapse strength for the pipe with multiple notches calculated based on the elasticperfectly plastic model. Moreover, the integrity assessment of the pipe with multiple notches
with total notch angle 90 degree subjected to combined load can be performed more easily and
conservatively using the theoretical plastic collapse strength for the pipe with single notch with
notch angle 90 degree.
REFERENCES
[1]
[2]
[3]
Nuclear Safety Review for the Year 2012, IAEA, 2012, pp. 32.
SME Boiler and Pressure Vessel Code, Section XI: Rules for Inservice Inspection of Nuclear
Power Plant Components, 2011 edition, July 1.
MATSUBARA, M., IZAWA, S., HIRAO, N., BUSUJIMA, K., KOYAMA, T., MACHIDA. K.,
KAWADA, D., SAKAMOTO, K., NEZU, K.: Development of Testing Equipment for Studying
Statically Indeterminate Fracture Mechanics, Proceedings ICPVT-10, 2003, pp. 481-486.
20
DEVELOPMENT OF A CRACK OPENING DISPLACEMENT ASSESSMENT
PROCEDURE CONSIDERING CHANGE OF COMPLIANCE AT A CRACK
PART IN THIN WALL PIPES MADE OF MODIFIED 9Cr-1Mo STEEL
T. WAKAI1*, H. MACHIDA2, M. ARAKAWA2, S. YOSHIDA2, S. YANAGIHARA3,
R. SUZUKI3, M. MATSUBARA3, Y. ENUMA4
1
4002 Narita-cho O-arai Ibaraki 3111393, JAEA, Japan; email: wakai.takashi@jaea.go.jp
2-37-28 Eitai Koto-ku Tokyo 1350034, TEPCO Systems Corporation, Japan
3
1-5-1 Tenjin-cho Kiryu 3768515, Gunma University, Japan
4
2-34-17 Jingumae shibuya-ku Tokyo 1500001, Mitsubishi FBR Systems Inc., Japan
2
KEY WORDS: leak-before-break, crack opening displacement, Mod.9Cr-1Mo steel
This paper describes a crack opening displacement (COD) assessment procedure used in
Leak-Before-Break (LBB) assessment of sodium pipes of the Japan Sodium cooled Fast
Reactor (JSFR). For sodium pipes of JSFR, the continuous leak monitoring will be adopted as
an alternative to a volumetric test of the weld joints under conditions that satisfy LBB. The
sodium pipes are made of ASME Gr.91 (modified 9Cr-1Mo steel). Thickness of the pipes is
small, because the internal pressure is very low. Modified 9Cr-1Mo steel has a relatively large
yield stress and small work hardening coefficient comparing to the austenitic stainless steels
which are currently used in the conventional plants. In order to assess the LBB behavior of the
sodium pipes made of modified 9Cr-1Mo steel, the coolant leak rate from a through wall crack
must be estimated properly. Since the leak rate is strongly related to the crack opening
displacement (COD), an appropriate COD assessment method must be established to perform
LBB assessment. However, COD assessment method applicable for JSFR sodium pipes - thin
wall and small work hardening material - has not been proposed yet.
Taking non-linearity of the material and the geometry of JSFR pipes, a series of finite
element analyses (FEA) for the pipe containing a circumferential through-wall crack was
conducted. Based on the parametric FEA results, engineering formulae for COD evaluation
were established [1].
In this method, total elasto-plastic COD, EP, was calculated by classifying the components
of COD; elastic, EE, local plastic, LP, and fully plastic, FP, as follows,
 EP   EE   LP   FP .
(1)
In order to estimate these COD components, elastic, elasto-plastic and plastic FEA were
performed.
The elastic COD, EE, was formulated based on the formulae of the GE/EPRI method [3]
with minor corrections. In accordance with the GE/EPRI methods [2], the COD corresponding
to small scale yielding condition is evaluated using an effective crack angle, eff, in the elastic
COD. The effective crack angle expresses the effect of increasing the COD due to local plastic
deformation around the crack tip as the crack length increases. For large work hardening
materials, such as austenitic stainless steels, the relationship between stress and COD can be
described from elastic to plastic regions smoothly. However, for small work hardening
materials, such as modified 9Cr-1Mo steel, the COD increased sharply. Therefore, in the
proposed method, local plastic COD, LP, is calculated separating from elastic COD, EE, by
using the following equations.
21
 LP


 0.2  qX
 
 R V ,
 4  X  
0 D

 
p

(2)

VD  exp m1  m2 X  m3 X 2  m4 X 3 ,
(3)
where  is the crack half angle, X is the ratio of primary load to plastic collapse load, R is the
radius of the pipe and 0 is strain at proportional limit. p, q are material constants. The
coefficients m1, m2, m3 and m4 are given in tabular form based on the parametric FEA results.
Based on the plastic FEA results, the COD produced after reaching fully plastic conditions,
FP, was formulated as follows;
 FP   0 Rh2 X n ,
(4)
where  and n are coefficient and exponent of Ramberg-Osgood stress-strain relation of the
material, respectively. h2 are given in tabular form based on the plastic FEA results.
Fig. 1. Statically indeterminate fracture mechanics test
machine.
Fig. 2. Comparison between calcurations and
observations.
For the verification of the COD assessment method, a series of tests was conducted under
displacement controlled condition. The experimental apparatus is shown in Fig. 1. The
specimen was a tube with 31.8 mm in outer diameter and 1.5 mm in thickness. A
circumferential through wall crack was machined by electric discharging. The specimen was
subjected to tensile load and bending moment sequentially or concurrently. Figure 2 show an
example of the comparison between calculations and experimental results. In this case, as far
as the load was small, the calculated COD was in a good agreement with observations.
Acknowledgement: This paper includes results of “Technical development program on a
commercialized FBR plant” entrusted to JAEA by the Ministry of Economy, Trade and Industry
of Japan (METI).
REFERENCES
[1]
[2]
WAKAI, T. et al.: Development of LBB Assessment Method for Japanese Sodium Cooled Fast
Reactors (JSFR) Pipes (2) -Crack Opening Displacement Assessment of Thin Wall Pipes Made
of Modified 9Cr-1Mo Steel-, ASME PVP 2010-25249.
Electric Power Research Institute: Crack-Opening Area Calculations for Circumferential
Through-Wall Pipe Cracks, NP-5959-SR.
22
STRESS RELAXATION SMALL PUNCH TESTING OF P92 STEEL
P. DYMÁČEK1*, F. DOBEŠ2, M. JEČMÍNKA3
CEITEC-IPM, Žižkova 22, 61662 Brno, Czech Republic; email: pdymacek@ipm.cz
Institute of Physics of Materials AS CR, Žižkova 22, 61662 Brno, Czech Republic
3
CAIT-VŠB-TU Ostrava, 17. listopadu 15, 708 33 Ostrava-Poruba, Czech Republic
1
2
KEY WORDS: small punch test, stress relaxation, P92 steel
Small punch tests (SPT) are using specimens of a thin disc shape prepared from a small
amount of material that can be extracted directly from the surface of exposed parts without their
damage. In these tests, a puncher penetrates through the disc specimen into a hole [1]. Two
variations of this test type seem to have a good potential for use at elevated temperatures. First,
the test in which the puncher penetrates through the disc at a given constant rate of deflection
(i.e., central deflection measured in a direction perpendicular to the disc) and the necessary
force is measured; this test is marked as CDR (constant deflection rate). It has certain analogy
with the conventional tensile test. Second, the CF test (constant force) is a test in which the
puncher penetrates under constant load and the time dependence of the deflection is measured.
This test is similar to a conventional creep test. Both tests are run up to the rupture of the disc.
As a rule, the puncher is a ceramic ball or a bar with a hemispherical tip. In application within
the field of power- or thermal-generation industry, these tests should be performed at elevated
temperatures and they should be conducted in a protective atmosphere (usually Argon).
Recently, two other types of the small punch test have emerged: (i) the tested discs are
furnished with a precisely machined groove and their testing can then be compared with Charpy
impact tests [2, 3] and (ii) the loading mechanism is adjusted and the acting force is oscillating
[4]. In this way, the fatigue mechanisms and fatigue crack propagation can be studied.
New application of SPT could be employed as the stress relaxation testing at elevated
temperatures. Basically, the specimen has to be loaded at a given deflection rate to a specific
central deflection that conforms conditions of the membrane-stretching regime. The deflection
of the disc is then held constant and the force relaxes as the elastic strain is replaced with
inelastic creep strain. The force vs. time response during relaxation can be recalculated to stress
vs. time response, differentiated and divided by elastic modulus to give the creep rate and finally
its dependence on the stress.
As an experimental material for the study was chosen the ferritic-martensitic 9% chromium
steel P92, that was taken from a pipe with outer radius 800 mm and wall thickness 78 mm.
The relaxation small punch test SPT-R was done at temperature 600°C. The specimen was
deformed at constant rate of 0.25 mm/min to a deflection uR = 1.53 mm that produced the initial
force FR = 921 N and the recording provided force-time relation. Analogous dependence was
obtained from uniaxial tensile relaxation test at initial rate of 0.8 mm/min at 600°C. In both
tests was the maximum load reached at time of about 120 s.
It is possible to determine the initial relaxation force FR, residual force FRZ and the force
drop ΔF during the time that is needed for stabilization of the force. The conversion of force to
stress  in uniaxial test is done by dividing the force by specimen cross section.
In small punch relaxation test we can use the parameter Ψ that was found from relation of
creep tests on standard specimens vs. small punch specimens:
F    ,
23
(1)
where F is applied force,  is applied
stress and Ψ is parameter that relates the
force with stress for the same time to
rupture in both types of creep tests. For
steel P92 and stress 300 MPa the factor
Ψ 3.1 [5].
Table 1 Stress characteristics of SPT-R and uniaxial stress
relaxation test of P92 steel at 600°C.
Relaxation
test
SPT-R
Uniaxial
R
RZ

[MPa] [MPa] [MPa]
297
136
161
298
137
162

[%]
54
54
tR
[h]
5.1
4.9
The relaxation curves converted to stress shown in Fig. 1 show very good agreement. The
values of forces converted to stress values of initial stress R, residual stress RZ and stress drop
 are summarized in Table 1. The research will continue on SPT relaxation testing of other
materials to verify the initial promising results.
Fig. 1. Stress relaxation diagram of P92 steel at 600 °C for uniaxial tensile
test and SPT-R.
Acknowledgement: This work was partly realised in CEITEC - Central European Institute
of Technology with research infrastructure supported by the project CZ.1.05/1.1.00/02.0068
financed from European Regional Development Fund.
REFERENCES
[1]
[2]
[3]
[4]
[5]
CEN Workshop Agreement CWA 15627:2007, Small Punch Test Method for Metallic
Materials, Dec. 2007.
CUESTA, I. I., RODRIQUEZ, C., BELZUNCE, F. J., ALEGRE, J. M.: Analysis of different
techniques for obtaining pre-cracked/notched small punch test specimens. Engineering Failure
Analysis, Volume 18, Issue 8, (2011), pp. 2282-2287.
TURBA, K., GULCIMEN, B., LI, Y. Z.; et al.: Introduction of a new notched specimen
geometry to determine fracture properties by small punch testing. Engineering Fracture
Mechanics Vol.: 78 Issue: 16 (2011), pp. 2826-2833.
VILLARRAGA, M. L., EDIDIN, A. A., HERR, M., KURTZ, S. M.: Multiaxial Fatigue
Behavior of Oxidized and Unoxidized UHMWPE During Cyclic Small Punch Testing at Body
Temperature. Crosslinked and Thermally Treated Ultra-High Molecular Weight Polyethylene
for Joint Replacements, ASTM STP 1445, S.M. Kurtz, R.Gsell, and J. Martell, Eds., ASTM
International, West Conshohocken, PA, 2003.
DYMÁČEK, P., MILIČKA, K., DOBEŠ, F.: Analysis of potential factors influencing the
relation between small punch and conventional creep tests. Hunické listy (Metallurgical
Journal) 2010, vol. LXIII, pp. 50-53.
24
ABI TESTING OF REACTOR PRESSURE VESSEL STEEL
P. HAUŠILD1*, J. SIEGL1, A. MATERNA1
1
Czech Technical University in Prague, Faculty of Nuclear Sciences and Physical Engineering, Department
of Materials, Trojanova 13, 120 00 Praha 2, Czech Republic; email: petr.hausild@fjfi.cvut.cz
KEY WORDS: instrumented indentation, automated ball indentation, reactor pressure vessel steel
The reactor pressure vessel is one of the key safety components in the complex safety
assessments of nuclear power plants. The reactor pressure vessel cannot be replaced (from both
technical and economical reasons) so it often becomes the component determining the
operational safety of the nuclear power plants. WWER 440 nuclear reactor pressure vessel is
fabricated by welding of thick walled ring-type components [1]. The reactor wall contains the
base metal (chromium-molybdenum-vanadium low alloy 15Ch2MFA steel), the multilayer
welding seam (10ChMFT steel) and the two-layer cladding (25 chromium/13 nickel nonstabilized austenitic stainless steel Sv 07Ch25N13 and at least 2 passes of 18 chromium/10
nickel niobium stabilized Sv 08Ch18N10G2B austenitic stainless steel).
Mechanical properties can present a gradient through the wall thickness, which can hardly
be assessed by conventional testing such as tensile or Charpy tests.
The elastic-plastic material properties of WWER 440 nuclear reactor pressure vessel steels
were therefore assessed by instrumented indentation tests carried out across the wall thickness
(Fig. 1). The Automated Ball Indentation (ABI) test [2] is based on multiple instrumented
indentation cycles (at the same penetration location) on a polished metallic surface by a
spherical indenter. Each cycle consists of indentation, unload and reload sequences.
Fig. 1. Microstructure through wall-thickness of WWER 440 pressure vessel with position
of indents in base metal (15CH2MFA), weld (10ChMFT) and cladding.
25
Fig. 2. Estimated yield stress determined by the ABI test in the different positions
of the weld, base metal and cladding of WWER 440 pressure vessel wall.
The representative (true) stress and (plastic) strain curves were determined by the ABI test
and the yield stress was estimated in the different positions through the wall thickness in the
weld, base metal and cladding as shown in Fig. 2. True stress – plastic strain curves presented
well defined power law hardening in the base metal, weld and cladding, which proves the
robustness of the method. Estimated yield stress (as well as the whole true stress – plastic strain
curves) is lower in the cladding (around 400 MPa) than in the base metal (around 500 MPa).
The results obtained by ABI were in a good agreement with results obtained by tensile test.
Although the ABI test is a macroscopic technique, it estimates the properties on a small volume
of material (few grains), which is particularly useful in testing e.g. welds and irregularly shaped
heat affected zones. Especially a multilayer welding seam presented a large scatter which can
hardly be assessed by conventional tensile testing.
Acknowledgement: This work was carried out with the financial support of Technology
Agency of the Czech Republic in the frame of the research project TA03011266.
REFERENCES
[1]
[2]
TIMOFEEV, B., BRUMOVSKÝ, M., VON ESTORFF, U.: The certification of
15Kh2MFA/15Cr2MoVA steel and its welds for WWER reactor pressure vessels, European
Commission, EUR 24581 EN, 2010.
HAGGAG, F.M., NANSTAD, R.K., HUTTON, J.T., THOMAS, D. L., SWAIN R.L.: Use of
automated ball indentation to measure flow properties and estimate fracture toughness in
metallic materials, Applications of automation technology to fatigue and fracture testing,
ASTM 1092, A.A. Braun, N.E. Ashbaugh, and F.M. Smith, Eds., American Society for Testing
and Materials, Philadelphia, 1990, pp. 188-208.
26
CHARACTERIZATION OF HETEROGENEOUS WELDMENTS
V. ŠEFL1*, R. NOVÁKOVÁ1, J. BYSTRIANSKÝ2
1
2
Institute of Chemical Technology, Technická 5 Praha 6, Czech Republic; email: vaclav.sefl@gmail.com
Department of Metals and Corrosion Engineering; Institute of Chemical Technology; Prague, Czech
Republic
KEY WORDS: heterogeneous weld, austenitic steel, carbon steel, steam, structure
Temperature
(°C)
Peak temperature Tp
Heterogeneous welds are basically two or more dissimilar metals connected together by
welding. Dissimilar welds are often required in power plant engineering as different parts of
the water circuit are made of different materials. Other application is production of surface with
specific composition for specific environments or repair of damaged parts of components.
Major concern comes from the welding process - the material is annealed during the process
thus affecting the former structure of material. These changes correspond to the
recrystallization, precipitation and also diffusion of elements from two adjacent metals, mainly
chromium and carbon. All of these phenomena have strong effect on mechanical and corrosion
properties; carbides segregated on grain boundaries and grain coarsening strongly affect the
ductility of the material, diffusing carbon and chromium change the corrosion behaviour. Even
when the diffusion is low, dissimilar metals in the weld can form a galvanic cell due to the
different corrosion resistance. This can, combined with the effect of some precipitates, result in
attack of the fussion layer between the two metals, eventually leading to failure of the
component. Incorrect material selection, welding process and subsequent thermal treatment can
also lead to formation of microcracs providing another failure mechanism.
Solidified weld
Solid-liquid transition zone
Coarse prior austenite grains +
fine prior delta ferrite grains
L

1400
L
L

L
CGHAZ
1200
Grain growth zone

Grain refined zone
1000
FGHAZ
Intercritical zone
Over-tempered region
800

Unaffected BM

600
Heat affected zone
0
0.2
0.4
0.6
Carbon (wt %)
0.8
1
Fig. 1. Structure of heterogeneous weld.
In this work, we used model samples of carbon 22K steel welded with various austenitic
materials and samples of heterogeneous welds cut from steam generator of VVER 440 nuclear
power plant after 20 years of exposure. Model samples were thermally treated in order to
simulate different welding conditions and aging during exposure.
Structure of all heterogeneous weldments was studied using standard metallographical
methods, their chemical composition was verified using scanning electron microscope Tescan
VEGA 3 equipped with EDS probe, the carbon distribution was studied with microanalyzer
27
COMEBAX equipped with WDS probe. The corrosion behaviour was mostly studied by
DL-EPR method (dual-loop electrochemical potentiodynamic reactivation); portion of the
model samples were placed in an autoclave and exposed to model environment at elevated
temperature. Kinetics of oxide layer growth and their composition across the weld was studied
by the techniques described above. Vicker´s hardness test, as the only method viable for
measurements in small-areas of the weld, was used to study mechanical properties. Possibility
of galvanic coupling between the two metals with different corrosion resistance was verified.
Acknowledgement: The authors gratefully acknowledge the support by (the Institution).
28
RESISTANCE TO CORROSION CRACKING OF STEEL 100Cr6 IN HUMID
AIR UNDER HIGHER TENSILE STRESSES
S. LASEK1*, V. ČÍHAL1, M. BLAHETOVÁ1, E. KALABISOVÁ2
VŠB-Technical University of Ostrava, 17. listopadu 15/2172, Ostrava-Poruba, 70833, Czech Republic;
email: stanislav.lasek@vsb.cz
2
SRIMT, Prague
1
KEY WORDS: steel 100Cr6, stress corrosion cracking, humid air, deformation
In the automotive industry are introduced and applied high-strength steels for reducing of
weight and overall costs. Under specific conditions and environments the components made of
high-strength steels can be sensitive to stress corrosion cracking (SCC) or corrosion fatigue.
Cracks and failures of unknown origin have occurred on the parts made of high-carbon chrome
100Cr6 steel. The aim of the contribution is comparison and evaluation the resistance to SCC
of selected steel under atmospheric conditions with relative humidity 40-80% at room
temperature.
For SCC testing were used the tensile type of samples (working part Ø5.0 x 33 mm) made
of 100Cr6 steel (EN 1.3505) in two steelworks (A, B). Standard chemical composition of steel
(wt. %): 0.95 - 1.05% C, 1.35-1.65% Cr, 0.25-0.45% Mn, 0.15 to 0.35% Si, Ni ≤ 0.3%,
Cu ≤ 0.3%, P ≤ 0.03%, S ≤ 0.025%. This steel can be treated by soft annealing and/or
quenching.
The values of yield stress of A steel were Rp = 360-380 MPa, and B one Rp = 490-520 MPa.
The stress corrosion tests were conducted on the samples under constant tensile stress in the
range 600-610 MPa (true stress 612–642 MPa) at room temperature in laboratory atmosphere
(35-50% rel. humidity, first test series) and in humid air at 80% r.h. (second tests).
Stereomicroscopy and scanning electron microscopy were used for the surface and
corrosion study. The microscopic valleys (grooves and lines after tooling, like a stress
concentrators) were observed on samples surface after preparation. Under test conditions,
locally rusty spots or stains were observed, while some of them appeared as cracks at
macroscopic observation, Fig. 1. During test time (3700 h) the fracture has not occurred.
+σ
1mm
(a)
(b)
Fig. 1. a) Surface of sample (A3) after exposition in humid atmosphere (3700 h, 25°C, 80% r.h.). Non-uniform
corrosion, rust stains and areas. b) Detail of surface attack, microscopic pits and microcracks.
Microscopic grooves and pits were formed probably during non-uniform local corrosion
under rust spots, see Fig. 1, where possible initiation of microcracks (SCC) is also shown.
Differences between A and B steel with respect to SCC resistance were not found out. The
initiation of microcracks is caused probably by carbide particles and microscopic pits.
29
Small plastic deformations (3.5-5.0% for A steel and 1.6-2.5% for B one) and low
temperature logarithmic creep was detected and measured on the samples, Fig. 2. Registered
creep deformation was smaller in average for steel with higher yield stress (B).
0,7
A4
B3
A3
B4
deformation [%]
0,6
0,5
0,4
0,3
0,2
0,1
0
0
200
400
600
800
1000
time t [h]
Fig. 2. Low temperature creep registered on the samples (second series).
The mean slow strain rate ἑ = ∆ε/∆t was calculated in the range (5-9).10-10s-1. The
recommended initial strain rate that promoted cracking of ferritic or tempered steels in water is
10-6 s-1. The resistance of tested 100Cr6 steel to SCC is relatively high under tested conditions.
Acknowledgement: This paper was created in the project No. L01203 “Regional Materials
Science and Technology Centre” – Feasibility Program. Founded by Ministry of Education,
Young and Sports of Czech Republic.
REFERENCES
[1]
[2]
[3]
ASTM G 49: Standard Practice for Preparation and Use of Direct Tension Stress-Corrosion
Test Specimens. 2000, 5 p.
ISO 7539-4: Corrosion of metals and alloys. Stress corrosion testing. Part 4: Preparation and
use of uniaxially loaded tension specimens. 1989.
LASEK, S., BLAHETOVÁ, M., ČÍHAL, V.: Stress Corrosion Cracking Study of Steam
Generator Bolt Steel. In Workshop: Fracture Damage of Structural Parts, VŠB-TU Ostrava,
2004, pp. 103-111.
30
RELATIONSHIPS BETWEEN KIC AND CVN AT TEMPERATURES LOWER
THAN NDT
M. HABASHI1*, M. TVRDY2
1
2
Research advisor at C.N.R.S. France; email: mahmoud.habashi@orange.fr
Professor at VSB – Technical University of Ostrava, Czech Republic
KEY WORDS: fracture toughness KA or KIC, impact energy CVN, spink's model, internal hydrogen,
mild steel, metastable austenitic 2404 alloyed steel, fracture features
In the upper shelf and lower shelf of the transition curves CVN-T, several results have
shown the existence of different relationships between mechanical toughness KIC measured
either by linear elastic fracture mechanisms (LFEM) or by applying J-integral JIC and impact
energy CVN (Charpy-V notch). In the upper shelf the most commonly used relationship is that
of Barsom-Rolfe:
(KIC /σy )2 = 5(CVN/σy – 0.05).
(1)
However, two analytical equations have been derived respectively by: a) Rithie, Francise
and Server and b) Rithie and Horn:
KA = 2.9 σy [exp ( σf /σy - 1)]1/2 ρ1/2,
(2)
KA = (3/2 σy E σf )1/2 ρ1/2,
(3)
where KA the apparent mechanical toughness and ρ the notch root radius.
Applying Spink's model, and thus plotting the variation of (KA/KIC)[1+(ρ/C)1/2] against
(ρ/C)1/2 with C the notch length, the results issued from literature and obtained at the upper
shelf, have shown that all the relations are linear with slopes which increase as the fracture
stress σf or the yield strength σy is higher. Furthermore; the characteristic distance ρ0 or the
effective notch root radius can be deduced from these slopes. The objective of this work is to
verify the validity of these relations at temperatures lower than the nil ductile temperature
(NDT). Knowing that in this field of temperatures, previous results have shown that CVN is
sensitive to the addition of elements in steel, heat treatments and eventually the existence of
defects such as those caused by internal hydrogen. Mild steel, with and without internal
hydrogen and a metastable austenitic 2404 alloyed steel are studied at -196°C (liquid nitrogen).
Standard Charpy specimens with different notch root radii varying from 0 to 0.7 mm are used
to measure KIC by applying J-integral and also to measure the impact energy CVN. For all;
bending tests, with high strain rate to measure the impact energy or with very slow strain rate,
were performed and the tests temperature was -196°C. For mild steel without internal hydrogen,
the changes in (KA/KIC)[1+(ρ/C)1/2] = Θ(ρ/C)1/2 and [CVN/(CVN)0 [1+(ρ/C)1/2 ] = Φ (ρ/C)1/2
are in good agreement with those obtained, in the upper sheld, by Ritchie et al in AISI 4340
steel in two different heat treatments, figure 1. However; in the case of mild steel severely
charged with internal hydrogen and containing more than 10 ppm H2, which promotes high
density of defects in the grain boundaries (hydrogen attack), the two linear relations are not
similar. The bi-phases 2404 alloyed steel (80% acicular martensite + 20% austenite) shows that
the slopes and the critical notch root radii of the linear relations are fairly the same. The
isothermal transformation of the residual austenite γ to the martensite α' during the
measurements of KIC with low strain rate (1.6710-5 m / s) is assumed to be responsible for this
difference. However; for all three cases studied here, in the lower shelf or from the results, in
the upper shelf, obtained by Rithie et al, the effective notch root radii whether measured by
fracture toughness or by impact energy tests are about the same, figure 2. The fracture type in
31
mild steel free from internal hydrogen is by micro-cleavage while in the presence of internal
hydrogen, micro-cleavage and inter-granular features, with large cracks are observed. After
fracture toughness tests, the fracture surface of the metastable 2404 alloyed steel shows a
mixture of cleavage-inter-granular features. The main conclusion is that by applying, from now
on, the Spink's model described and used above, large dimension specimens satisfying the
standard LFEM criterion are now not necessary. In homogeneous micro-structure steels, KA or
KIC could be deduced by measuring (CVN)o.
0 .2
0. 4
0. 6
0. 8
5
Mild Steel
T= -196°C
4
3
2
2
0
3
A
IC
(K /K )[1+( r/C)
C 1/2]
4
[(CVN)/(CVN) )] 1/2 [1+( r/C)
C 1/2]
0. 0
5
1
1
(K /K )[1+( r /C)
C 1/2]
A
IC
[(CV N)/(CV N) ] 1/2 [1+( r /C)
C 1/2]
0
0
0 .0
0 .2
0,4
0
0. 8
0. 6
( r/C)
C 1/2
Fig. 1. (KA/KIC) and CVN/(CVN)0 against (ρ/C)1/2. Mild Steel.
20
40
60
80
100
120
140
180
180
160
160
140
140
AI SI 4340 St eel(1200-800°C)
200
60
2
40
20
Mild St eel
80
' 2404 Al loy
100
AI SI 4340 S teel(870°C)
120
Mil d St eel (H )
r0 from (KA/KIC), µmm
0
200
120
100
80
60
40
20
0
0
0
20
40
60
80
100
r0 from [(CVN)/(CVN)0)]1/2, µm
120
140
Fig. 2. ρ0 measured from (KA/KIC) and from CVN/(CVN)0 for different steels
in the upper and lower shelds.
32
BIOMECHANICS - PROBABILISTIC RELIABILITY ASSESSMENT OF
FEMORAL SCREWS
K. FRYDRÝŠEK1*
1
VŠB-Technical University of Ostrava, Faculty of Mechanical Engineering, Department of Mechanics of
Materials, Czech Republic; email: karel.frydrysek@vsb.cz
KEY WORDS: biomechanics, traumatology, orthopaedics, femoral neck fracture, femoral screw,
strength analyses, beam, elastic foundation, probability, Monte Carlo Method, reliability assessment
Proximal femoral neck fractures, see
Fig. 1, remain a vexing clinical problem in
traumatology and are one of the common types
of trauma, see [1], [2]. One possible treatment
method for femoral neck fractures, is the
application of femoral screws made of
Ti6Al4V or stainless steel material. This paper
therefore aims to present a numerical model
(i.e. strength analysis) of femoral screws
together with a probabilistic reliability
assessment (i.e. an application of the
Simulation-Based Reliability Assessment
(SBRA) Method, Anthill SW, Monte Carlo
Method, see [3], [4] and [5]), which is a
modern and innovative stochastic trend. The
analytical model is based on the theory of
beams on an elastic foundation, see [4], where
the bone is approximated by the elastic
foundation.
Fig. 1. Femoral screws in femur as beams on elastic
foundation and their loading.
Fig. 2. Examples of histogram for input parameter
Re /MPa/.
Hence, the femoral screw is resting on an
elastic foundation prescribed by stiffness
k /Pa/. Three screws of length L = 90 mm are
applied in parallel positions on the elastic
foundation,
see
Fig. 1a.
The
force
F188.3;2129.2 N acting in one beam
(screw) can be defined via total loading force
Fm /N/, see Fig. 1b, by the equation
F = Fm/n=mkmkdyng/n. The variables are as
follows: m70;145 kg is the entire mass of a
patient; km0.72;0.82 is the coefficient of
mass reduction (i.e. the mass of one lower limb
is not acting, see Fig. 1b); kdyn1;4 is the
dynamic force coefficient; g = 9.81 ms-2 is the Fig. 3. Example of bending moment distribution for
gravitational acceleration; and n2;3 is the cannulated femoral screw (result of 1 Monte Carlo
simulation).
coefficient of inequality in the division of force
Fm into three screws. These variables are defined via truncated histograms (stochastic
approach); see Fig. 2 for examples. The force F can be decomposed into forces F1=Fcos 
and F2=Fsin ; see Fig. 3. The femoral screw angle 5;80 deg, which is defined by the
limiting angles of adduction and abduction, and the yield limit of material Re /MPa/, are
33
likewise defined via histograms; see Fig. 2. According to the 2nd order theory and the theory of
beams on an elastic foundation, three linear differential equations for the intervals x1,2,3, can be
d 4v
d 2v
written as EJ ZT 4i  F2 2i  kvi  0 together with 12 boundary conditions. EJZT /Nm2/ is
dxi
dxi
flexural stiffness, i /m/ is displacement and xi /m/ are coordinates. For more information, see
5 and 6. Hence, bending moments Moi /Nmm/, see Fig. 3, and maximum stresses
MAX–1129.2; –35.5 MPa, see Fig. 4, can be calculated.
Probabilistic reliability assessment can be
carried out via the SBRA Method by means of
the reliability function RF=Re–MAX /MPa/
(depending on load capacity, compared to the
extreme stress values with yield limit, see
Fig. 4). It is then obvious that if RF < 0, plastic
deformation occurs in the beam (i.e. the yield
limit of the material is overcome; see the 2D
histogram in Fig. 4. The probability that plastic
deformation will occur in the beam is
PRF<0=2.5110–5=0.00251%.
According the theory of beams on elastic Fig. 4. 2D histogram of output parameter RF /MPa/.
foundation in connection with SBRA Method,
own stochastic model for strength analyses of femoral screws intended for treatment of femoral
neck fractures was derived. The probability that yield limit is exceeded is 0.00251% and it is
acceptable. Hence, the femoral screws are safe and suitable for patient treatment. The presented
results were compared with 3D FE model with adequate results (deterministic approach).
Acknowledgements: Supported by the Czech projects TA03010804 and SP2014/17.
REFERENCES
[1]
[2]
[3]
[4]
[5]
FRYDRÝŠEK, K., JOŘENEK J., UČEŇ, O., KUBÍN, T., ŽILKA, L., PLEVA, L.: Design of
External Fixators Used in Traumatology and Orthopaedics – Treatment of Fractures of Pelvis
and its Acetabulum, Procedia Engineering, vol. 48, 2012, pp. 164-173.
FARHAD, N., BRADLEY, E.J., HODGSON, S.: Comparison of Two Tools for the
Measurement of Interfragmentary Movement in Femoral Neck Fractures Stabilised by
Cannulated Screws, Robotics and Computer-Integrated Manufacturing, vol. 26, issue 6, 2010,
pp. 610-615.
FRYDRYŠEK, K.: Probabilistic Calculations in Mechanics 1, Department of Mechanics of
Materials, Faculty of Mechanical Engineering, VŠB - Technical University of Ostrava, Ostrava,
Czech Republic, 2010, ISBN 978-80-248-2314-0, p. 1-149, written in Czech language.
FRYDRÝŠEK. K., TVRDÁ, K., JANČO, R. et al: Handbook of Structures on Elastic
Foundation. VŠB - Technical University of Ostrava, Ostrava, Czech Republic, 2013,
ISBN 978-80-248-3238-8, p. 1-1691.
FRYDRÝŠEK. K.: Strength Analyses of Full and Cannulated Femoral Screws Made up from
Stainless Steel and Ti6Al4V, Calculation report, FME VŠB-Technical University of Ostrava,
Ostrava, Czech Republic, 2014, pp. 1-43.
34
FINITE ELEMENT HUMAN MODEL FOR CRASH SAFETY ASSESSMENT
OF AUTOMOBILE IN FRONTAL COLLISION
Y. ZAMA1*
1
1-5-1, Tenjin-cho, Kiryu, Gunma, Japan; email: yzama@gunma-u.ac.jp
KEY WORDS: crash safety, frontal collision, injury mechanism, FEM
In the assessment for crash safety of automobile, crash test dummies such as Hybrid III and
THOR are utilized commonly. However, injury assessment by the dummy is only for specific
parts, which are head, neck chest and so on. Recently, finite element human model (FE human
model) for crash safety assessment has been developed in the world. The FE human model can
mimic human structure and kinetic characteristics of human body precisely as compared with
the crash test dummy. Moreover, cost of the crash safety assessment by using the model can be
reduced more than the crash test with the dummy. Therefore, crash test simulation by the model
have been carried out in order to clarify injury mechanism and evaluate crashworthiness of
automobiles.
Since 2004, the FE human model have been
developed in the project of Japan automobile
manufacture association (JAMA) in order to use
the model as common tool for crash safety
assessment. In this study, FE human model of
occupant in the event of frontal collision was
developed as shown in Fig. 1. The model
consisted of 90,000 elements with finite
element. A physique of the model was 50%tile
of American male, and UMTRI posture was
Fig. 1. FE human model of occupant for frontal
applied as occupant posture. This development
collision.
was performed in Japan automobile research
institute (JARI) as the previous work of author. The contents of this report already have been
presented and published in a journal [1].
As for the frontal collision of automobile, chest injury frequently occurred. Bio-fidelity of
chest part in the model is important in order to investigate injury mechanism in the real world.
Bending characteristics of ribs and clavicles in the model was validated with the experiment
results of bending test of the ribs and clavicles extracted from post mortem human subject
(PMHS)[2][3]. Force-displacement curve of the rib and clavicle in the current model was far
different from that of the PMHS ribs and clavicles, and material properties of the ribs and
clavicles in the model were estimated with force-displacement curve of the experiments. Rib
and clavicle models with the estimated material properties showed the similar forcedisplacement curve of the experiment.
As for compression characteristics of chest part in the model, the characteristics of the
model was validated with experiment results of table top test by using only chest of PMHS. In
the table top test [4], the chest of PMHS was compressed with seat belt, and compression ratio
derived from chest displacement was evaluated. Figure 2(a) shows relationship between chest
compression ratio and force loading to the chest with seat belt for the model and PMHS. The
compression characteristics of the model was also far different from that of PMHS. The material
properties of organs and fresh in the model were modified. As the result, the compression
characteristics of the model showed good agreement with that of PMHS as shown in Fig. 2(b).
35
2000
belt
Avg (PMHS)
1500
＋σ
Model
1000
-σ
Avg (PMHS)
+σ
Model
Force [N]
Force [N]
1500
table
2000
-σ
1000
500
500
0
0
0
5
10
15
Compression [%]
20
0
5
10
15
20
Compression [%]
(a) Current model
(b) Improved model
Fig. 2. Compression charavteristics of chest for the model and PMHS.
Finally, kinematics of the model in the
event of frontal impact was validated with the
result of PMHS frontal sled experiment [5].
In this experiment, trajectories of head,
thoracic spine, pelvis, greater trochanter,
patella and malleolus during frontal impact
were measured by the three dimensional
motion capture system. From the comparison
of the trajectories with the model and PMHS,
forward and upward motions of the model did
not coincide with those of PMHS. In order to
solve the issue, geometry and tensile property
of the fresh around the pelvis in the model
were optimized by considering of PMHS
experiment results. Fig. 3 shows trajectories
of the improved model and PMHSs. As the
results, forward and upward motions of the
improved model were similar to those of
PMHS. However, rotational motions after
100 ms from the collision in the model was
different from those in the PMHS.
(a) Side view
Acknowledgement: The author would
like to express special thanks for Dr. Ejima
and Mr. Mikami in JARI.
(b) Top view
Fig. 3. Trajectories of the model and PMHS.
REFERENCES
[1]
[2]
[3]
[4]
[5]
ZAMA, Y., ANTONA, J., MIKAMI, K., EJIMA, S., KAMIJI, K., YASUKI, T.: Development
of Finite Element Human Model for Events of Frontal Impact, Transactions of JSAE 41, 2010,
pp. 1243-1248.
KINDIG, M.: Tolerance to failure and geometric influence on the stiffness of human ribs under
anterior-posterior loading, Master thesis, University of Virginia, 2009.
DUPREY, S.: Biomechanical response of the clavicle under bending, Proc. 34th Cong. French
Society of biomechanics, 2009.
KENT, R.: Frontal thoracic response to dynamic loading: the role of superficial tissues,
viscera, and the rib cage, Proc. IRCOBI Conf., 2005.
SHAW, C. G, PARENT, D., PURTSEZOV, S., KERRIGAN, J. R., SHIN, J., CRANDALL,
J. R.: Frontal impact PMHS sled tests for FE torso model development, Proc. IRCOBI Conf.,
2009.
36
BIOMECHANICS – PROBLEMATIC LOOSENING
OF LOCKING SCREWS FROM PLATES
R. ČADA1*, K. FRYDRÝŠEK1
1
VŠB  Technical University of Ostrava, Faculty of Mechanical Engineering, Department of Mechanical
Technology, Czech Republic; email: radek.cada@vsb.cz
KEY WORDS: biomechanics, traumatology, orthopaedics, locking bone screw, angularly stable plate,
osteosynthesis, titanium alloy
Angularly stable plates (Fig. 1) have been used in medical and veterinary practice for the
treatment of unstable fractures for several years. The principle is based on the fixation of a
screw in the plate by means of the screw thread. The thread on the screw head must have the
same gradient as the thread on the shank; often the thread on the screw head is finer and has
multiple starts. Angularly stable plates are produced either as straight plates or as anatomically
shaped plates. Plates and bone screws
(Fig. 2) are produced from austenitic steel
W. Nr. 1.4441 according to ISO 5832-1 and
titanium
alloy
Ti6Al4V
according
to ISO 5832-3. For both these materials the
tensile strength is almost identical, at Fig. 1. Angularly stable plate with locking bone screws
and their application (photos a, b Čada).
860 - 1050 MPa. The modulus of tensile
elasticity is 1.135 GPa for titanium alloy and
2.1 GPa for stainless steel. Titanium alloy is
thus 145 % more elastic and can be surfacetreated by anodizing, creating titanium oxide
Fig. 2. Locking self-tapping bone screw and detail
on the implant surface.
of screw head (photos Čada).
In the operating procedure, locking bone
screws in the angularly stable plate (Fig. 3)
are manually tightened using a torque clutch
with torque 1.5 Nm. Even if this procedure is
complied with, when extracting the titanium
alloy implant it is difficult (and sometimes
impossible) to loosen some locking bone Fig. 3. Locking bone screw and locking hole in angularly
stable plate (photo b Čada).
screws from the locking holes in the
angularly stable plate (Fig. 3). Often the
extraction causes wear of the internal
hexagon in the screw head (Fig. 4), leading
to stripping. This then necessitates a
complex process of drilling off the screw
heads and removing the screw shanks from Fig. 4. Internal hexagon in the head of a locking selfthe bone using an extraction set, which
tapping bone screw (a – before, b – after using a
prolongs the length of time spent by the
screwdriver), (photos Čada).
patient in the operating theatre.
In order to solve the above-described problem, the self-locking properties of the thread
connection were verified by calculation, and the effect of the gradient angle of the screw head
on its self-locking properties was evaluated. The experiments were carried out using a
KRAFTWERK torque screwdriver, model 2039, 1-4 Nm (±6.0 %) according to ISO 6789
(Fig. 5), which was supplied with a calibration certificate.
37
The distances between the opposite walls of the
2.5 mm hexagonal bit used in the experiments were
measured 10× during rotation of the bit using a
Mitutoyo digital micrometer: 0-25 mm, 0.001 mm.
Subsequently the mean value and the absolute
measurement error were calculated. The torque
required to strip the internal hexagon of a 3.5 mm
locking bone screw made of titanium alloy Ti6Al4V
was determined experimentally. It was also
determined whether the rotation-slippage of the end
Fig. 5. Torque screwdriver with 2.5 mm bit and
of a hexagonal-bit screwdriver inside the bone angularly stable plate with locking bone screw
screw head when stripping the internal hexagon
(photo Čada).
(Fig. 6a) causes deformation in the circumference
of the screw head and thus increases friction in the thread. A comparison was made of the
measured values of the mean bone screw head diameter both before and after the stripping of
the internal screw head hexagon; absolute measurement errors were also coampared. It was
found that twisting of the shank did not lead to the loosening of the screw head from the
angularly stable plate, but instead caused the shank to fracture, creating a level fracture surface
(Fig. 6b). Experiments were performed to determine the effect of the torque and the influence
of the duration of tightening on the ability to unscrew the bone screw from the hole in the
angularly stable plate. The suitability of using hexagonal forms in screw heads was assessed
from the perspective of the active stresses.
The results of the experiments led to
the following recommendations: the torque
should be reduced from 1.5 Nm to a lower
value; conical (self-holding) screwdrivers
should be used when inserting screws and
cylindrical
(non-self-holding)
screwdrivers should be used for extracting
screws; instead of the internal hexagon, the
Fig. 6. Locking self-tapping bone screw in plate
screw head should use the TORX system, (a – stripped internal hexagon in screw head after using
which gives better torque transfer even screwdriver, b – fracture surface after fracture of the screw
shank), (photos Čada).
when the screwdriver is not fully inserted
into the aperture.
Acknowledgements: The authors gratefully acknowledge the funding from the projects
TA03010804 “Osteosynthesis of leg and arm fractures”, SP2014/193 “Research and
Optimization of Technologies for Higher Utility Properties of New Materials and Mechanical
Engineering Products” and SP2014/17 “Application of numerical and experimental methods
in the field of mechanics and biomechanics”.
REFERENCES
[1]
[2]
ANTOSZEWSKI, B., EVIN, E., AUDY, J.: A study of the effect of type (Cu+Ti) and (Mo+Ti)
electro-spark coatings on fricion in pin-on-disc testing, Journal of Tribology, Vol. 130, No. 1
(2008), pp. 26-31, ISSN 0742-4787.
FRYDRÝŠEK, K., JOŘENEK J., UČEŇ, O., KUBÍN, T., ŽILKA, L. and PLEVA, L.: Design
of External Fixators Used in Traumatology and Orthopaedics – Treatment of Fractures of
Pelvis and its Acetabulum, Procedia Engineering, Vol. 48, 2012, pp. 164-173, ISSN 1877-7058.
38
BIOMECHANICS – SAFETY FACTOR EVALUATION
OF ANTEROLATERAL PLATES FOR DISTAL TIBIA FRACTURES
G. THEISZ1*, K. FRYDRÝŠEK1
1
Department of Mechanics of Materials, Faculty of Mechanical Engineering, VŠB – Technical University
of Ostrava, 17. listopadu 15/2172, 708 33 Ostrava, Czech Republic; email: gunther.theisz@vsb.cz
KEY WORDS: biomechanics, traumatology, safety factor, internal fixation, anterolateral plate,
strength analysis
Two types of fracture osteosynthesis are used in medical
practice – external and internal fixation. This paper analyzes
anterolateral plates; see Fig. 1 (producer: Medin, a.s.). The plate is
used for the internal fixation of distal tibia fractures.
Fig. 1. Anterolateral plate.
The material properties of bone can be described
with sufficient accuracy for individual bone parts using
a homogeneous isotropic material model. This model can be used
Fig. 2. Parts of the tibia
to describe tissue using the modulus of tensile elasticity E /MPa/
including
modulus of elasticity.
and the Poisson number μ /1/ for individual bone parts. In order to
provide an adequate description of reality, the bone model was
divided into cortical and spongy (cancellous) tissues. These
individual parts were each divided into 4 areas (giving a total of 8
areas displaying different material properties); see Fig. 2. In the
condyle areas the bone tissue is considerably harder and stronger
than in other parts of the bone, and so it shows a higher modulus
of tensile elasticity in the model of cancellous bone.
The analysis was performed on a selected simple intra-articular
fracture of the metaphysis and joint surface defined according to
the AO classification 43-C1.1; see Fig. 3. Coulomb friction
between the non-fused bone fragments is defined with friction
coefficient 0.4.
It is very difficult to determine the force acting upon the tibia
during walking. For this reason the force F was selected,
corresponding with the entire weight of the patient mp = 100 kg,
taking into account the dynamic loading of the bone by applying
the dynamic coefficient kDyn = 1.6.
F  k Dyn .m p .g  15691.1N
(1)
The contact between the tibia and the talus is replaced by
the boundary condition of elastic support; this boundary condition
is a suitable replacement for the cartilage and mechanical contact
with the talus.
39
Fig. 3. Stress distribution
according to HMH theory.
Fig. 3 shows the distribution
of reduced stress red according to
von Mises theory for a plate
consisting of material Ti6Al4V.
This calculation is performed for
non-fused bone. The maximum
stress
reaches
the
value
of 641.24 MPa at the hole for
the K-wire. Fig. 4 shows the detail
of the maximum stress in the hole.
Variant calculations were
carried out for fused bone
Fig. 4. Detail of distribution of reduced stress in K-wire hole.
(successful treatment) and nonfused bone (unsuccessful treatment); see Tab. 1.
With regard to the minimum yield strength of the material, the safety factor is calculated
using the following equation:
ky 
 redMAX
,
y
(2)
where redMAX is the maximum reduced stress according to von Mises theory and y
is the minimum yield strength of the given material.
The situation of maximum loading on non-fused bone is an extreme state in which
the fracture fails to heal and the patient places excessive stress on the limb (with loading even
exceeding yield strength – see Tab. 1 – causing plastic deformation). In cases of successful
treatment the safety factor kσy > 10. From this perspective the analyzed anterolateral plate can
be considered safe. In cases of unsuccessful treatment, the safety factor may be kσy < 1. In such
cases the plate is unsafe and the fracture must be re-operated.
Table 1 Safety factor.
Minimum material yield strength σy /MPa/
Maximum calculated stress σred MAX /MPa/
Non-fused bone
(unsuccessful treatment)
Safety factor kσy /1/
Maximum calculated stress σred MAX /MPa/
Fused bone
(successful treatment)
Safety factor kσy /1/
Material
Stainless
Titanium
steel
(Ti6Al4V)
(1.4441)
758
690
641.24
728.61
1.18
0.95
45.76
63.74
16.5
10.82
Acknowledgements: This work was supported by the Czech projects TA03010804 and
SP2014/17.
REFERENCES
[1]
[2]
FRYDRÝŠEK, K., JOŘENEK J., UČEŇ, O., KUBÍN, T., ŽILKA, L., PLEVA, L.: Design of
External Fixators Used in Traumatology and Orthopaedics – Treatment of Fractures of Pelvis
and its Acetabulum, Procedia Engineering, vol. 48, 2012, pp. 164-173.
COWIN, C. S.: Bone Mechanics Handbook. CRC Press, Florida, USA, 2001.
40
CYCLIC BENDING DEFORMATION AND FRACTURE OF
Al AND Al-1.0MASS%Mg ALLOY
H. IKEYA1*, H. FUKUTOMI1
1
Yokohama National University, Japan; email: ikeya-hayato-nf@ynu.jp
KEY WORDS: aluminum, cyclic bending, grain size, work hardening
Reduction in weight is an urgent issue for automobiles in order to improve the fuel
efficiency for the decrease of CO2 gas emission. Replacement with light materials has been
challenged on every assembly of vehicles. Aluminum alloys are light-weight high strength
materials with relatively high electric conductivity and hence much attention has been paid by
automobile and its related industries. As a structural material, the application of aluminum
alloys to car body panels has been examined. In order to use the high electric conductivity
together with the weight advantage, application of aluminum alloys to the electric wires has
been expected. The replacement of electric wires made of copper with aluminum alloys gives
also one of the solutions for the exhaustion of copper resources at the same time.
Application of aluminum wires in vehicles, however, is quite limited at present, because
that strength, toughness, and electrical conductivity of aluminum are not enough in comparison
with copper wires. In order to extend the application area of aluminum wires, it is necessary to
develop aluminum alloys with mechanical properties, especially fatigue characteristics, better
than the present aluminum wires without losing high electrical conductivity.
Although many studies have been conducted on the fatigue of aluminum and aluminum
alloys (see e.g., [1]), studies on cyclic bending fatigue in the circumstances close to the actual
vehicle-fitted condition are limited and hence the fatigue process at the cyclic bending is not
clarified experimentally enough.
In this study, deformation and fracture behavior at the cyclic bending are investigated on
aluminum and Al-1.0mass%Mg specimens with well defined microstructures, as it is known
that grain size and crystal orientation distribution give strong effects on the deformation
behavior of polycrystalline materials. In addition, pure copper is examined for comparison.
Al-1.0mass%Mg alloy was produced by melting 99.99mass% pure Al and 99.9mass% pure Mg
in a high-frequency induction furnace. The purity of aluminum and copper used for cyclic
bending tests is 99.99mass% and 99.9mass%, respectively. Wire-shaped flat plates with the
dimensions of 0.6 mm in thickness and 0.8 mm in width were produced by drawing and used
for the cyclic bending test. The cyclic bending tests were carried out using a cyclic bending
machine at room temperature. Bending angle, bending rate and the maximum bending strain
are 90°, 50 rpm, and 0.02, respectively. Grain size was controlled by heat treatments. Table 1
shows the characteristic of three materials in this study.
Al-1.0mass%Mg
Pure-Al
Pure-Cu
Table 1 Characteristic of three specimens.
Yield strength
Tensile strength
Elongation
(MPa)
(MPa)
(%)
74 – 102
30 – 42
8 – 15
31 – 51
17 – 24
7 – 19
35 – 60
179 – 229
24 – 40
Grain size
(µm)
60 - 210
140 – 300
25 – 85
Figure 1a shows the relationship between grain size and cycles to failure. The maximum
number of cycles to failure is less than 5000 cycles in the present deformation conditions. It is
seen that the number of cycles to failure increases with a decrease in grain size. Thompson et
al. [2] reported that the concentration of dislocation sources increased with an increase in grain
41
size. If the slip activity is enhanced by an increase in grain size, it is expected that the specimens
with large grain size may form a long slip line, and dislocation density becomes high. It seems
that the failure at the cyclic bending is affected by the grain size through these microstructure
changes. Effect of microstructure on the failure can be also seen in the difference of the cycles
to failure for the same grain size. In Fig. 1a, distinct difference of the cycles to failure is seen
for pure Al and Al-1.0mass%Mg with a grain size of about 150 μm and Al-1.0mass%Mg and
Cu with a grain size of about 60 μm. This suggests that the slip line formation as well as the
dislocation microstructure plays an important role for the bending fatigue behavior. According
to the result in Fig. 1a it seems that lower stacking fault energy results in the higher cycles to
failure.
Figure 1b shows the relationship between work hardening exponent and cycles to failure.
Work hardening exponent is evaluated by the stress-strain curve up to a fracture strain. It is
seen that an increase in work hardening exponent results in an increase in cycles to failure.
However, the experimental results given in Fig. 1b suggest that cycles to failure vary
independently of the work hardening exponent in the same material. This indicates that work
hardening exponent might dominate the basic number of cycles to failure and microstructure
can contribute to the improvement in the failure resistance.
b
400
:Al-1.0mass%Mg
: Pure-Al
: Pure-Cu
Grain Size, m
300
Work Hardening Exponent,n
a
200
100
0
0.8
:Al-1.0mass%Mg
: Pure-Al
: Pure-Cu
0.6
0.4
0.2
0
102
103
104
102
Cycles to failure, Cycle
103
104
Cycles to failure, Cycle
Fig. 1. Relationship between the cycles to failure and grainsize (a),
and the cycles to failure and work hardening exponent (b).
REFERENCES
[1]
[2]
KAMP, N., GAO, N., STARINK, M. J., SINCLAIR, I.: Influence of Grain Structure and Slip
Planarity on Fatigue Crack Growth in Low Alloying Artificially Aged 2xxx Aluminum Alloys,
International Journal of Fatigue 29 (2007) 869-878.
THOMPSON, A. W., BACKOFEN, W. A.: The Effect of Grain Size on Fatigue, Acta Met.
Vol. 19 (1971) 597-605.
42
CYCLIC INSTABILITY OF STEEL-TITANIUM BIMETALLIC COMPOSITE
OBTAINED BY EXPLOSIVE WELDING
A. KAROLCZUK1*, T. ŁAGODA1
1
Opole University of Technology, ul. Mikołajczyka 5, 45-271 Opole, Poland; email: a.karolczuk@po.opole.pl
KEY WORDS: explosive welding, cyclic instability, fatigue
In an explosive welding process the energy of detonation is used to join metals. Using this
unique technology metals with highly dissimilar crystal structure can be welded. This advantage
is used for example in the following industry areas: chemical manufacturing, power generation,
shipbuilding, cryogenic gas production. The bonded metals during manufacture process absorb
large portion of kinetic energy and undergo large deformation that results in high hardened thin
joint zone. Fatigue resistance and fatigue phenomena of bimetal manufactured by explosive
welding are purely recognized.
According to the experimental fatigue analysis of bimetallic composite (S355J2+N –
Titanium Grade 1) performed under push-pull loading with force controlled amplitude the
bimetallic specimens exhibit ratcheting phenomena and cyclic instability (softening) [1]. The
example changes in the total strain amplitude a in the function of damage degree n=N/Nexp
(where: Nexp is the final fatigue life, N is current number of cycles) are shown in Fig. 1.
The observed cyclic instability can
be the result of instability of titanium
grade 1 or existence of residual stresses
[2]. The aim of the present article is to
propose fatigue characteristic in the
form of relation between strain
amplitude and number of cycles
associated with the given damage
degree n.
The basic mechanical properties of
parent materials are given in Table 1.
Since the mechanical properties of
parent materials are different the stress
amplitudes generated under push-pull
loading are not equal.
Material
S355J2+N
Tytanium Grade 1
6
x 10
-3
Nexp = 22980
Nexp = 26570
5
Nexp =104820
Nexp =895970
4
3
2
1
0
0.2
0.4
0.6
0.8
1
Fig. 1. The cyclic instability of investigated bimetallic
composite.
Table 1 Basic mechanical materials properties.
ReH, MPa
Rm, MPa
E, GPa
382-395
598-605
204-220
189-215 (Rp02)
308-324
100-104
, 0.27-0.30
0.37
A5, %
24-34
43-56
As a result the stress based fatigue characteristic for the investigated bimetal cannot be
created. Only the strain based fatigue characteristic is reasonable because the elongations of
both materials under push-pull loading are equal. However, the standard Manson-Coffin
characteristic cannot be applied since it is based on separation of the total strain into elastic and
plastic parts and these parts are different in both materials. Based on the mentioned reasons the
new strain based fatigue characteristic is proposed in the following form
 a  p1 2 N f p  p3 2 N f p ,
2
43
4
(1)
where p1, p2, p3, p4 are parameters to be determined experimentally. The form of equation 1 is
very similar to the Manson-Coffin equation but without division into elastic and plastic parts.
The parameters of equation 1 are calculated using experimental data, i.e. total strain amplitudes
a and number of cycles N to the given damage degree n. The identification of parameters has
been done based on the following aim functions:
er n, i    a n, i   p1 2 N n, i  2  p3 2 N n, i  4 ,
(2)
Er n   i 1 er n, i  ,
(3)
p
k
p
2
where: i it is the subsequent specimen, k is the total number of specimen (k=12). Minimization
of function (3) was performed using the Quasi-Newton line search method. The obtained fatigue
curves for three damage degrees n=[0.1; 0.5; 0.9] are presented in Fig. 2. For comparison
purpose the Manson-Coffin curve for S355J2+N steel is also presented in Fig. 2.
Fig. 2. Identified fatigue curves for three damage degrees
n=[0.1; 0.2; 0.3] with the Manson-Coffin curve for S355J2+N steel.
Based on the perforemed analysis the following conclusion are drawn: (i) The strain based
fatigue characteristic of S355J2+N-Titanium Gr. 1clad for n = 0.5 differes in large degree when
compared to steel characteristic; (ii) The proposed characteristic allows to estimate the number
of cycles to the given damage degree of S355J2+N-Titanium Gr. 1 clad under push-pull loading
Acknowledgement: The Project was financed from a Grant by National Science Centre
(Decision No. DEC-2011/03/B/ST8/05855).
REFERENCES
[1]
[2]
[3]
KAROLCZUK A., KOWALSKI M., BAŃSKI R., ŻOK F.: Fatigue phenomena in explosively
welded steel–titanium clad components subjected to push–pull loading, Int. J. Fatigue 48, 2013,
pp. 101–108.
KAROLCZUK A., KLUGER K., KOWALSKI M., ŻOK F., ROBAK G.: Residual stresses in
steel-titanium composite manufactured by explosive welding, Materials Science Forum 726,
2012, pp. 125-132.
KAROLCZUK A., KOWALSKI M., ROBAK G., Modelling of titanium-steel bimetallic
composite behaviour under mechanical cyclic loading, Solid State Phenomena 199, 2013,
pp. 460-465.
44
INCLUDING OF RATIO OF FATIGUE LIMITS FROM BENDING AND
TORSION FOR ESTIMATION FATIGUE LIFE UNDER CYCLIC LOADING
M. KUREK1*, T. ŁAGODA1
1
Faculty of Mechanical Engineering, Department of Mechanics and Machines Design, Opole University of
Technology, ul. Mikołajczyka 5, 45-271 Opole, Poland; e-mail: ma.kurek@po.opole.pl
KEY WORDS: fatigue life, multiaxial criteria
In the literature, there are many criteria of multiaxial fatigue. They are based on various
assumptions and parameters describing the process of fatigue. Among them, there is a special
group of criteria based on the concept of critical plane. Some of them in their equations take
into account the ratio of fatigue limits for bending and torsion. The paper presents the estimation
of the fatigue life under multiaxial cyclic loading of selected construction materials: two
aluminum alloys PA4 (6068) and PA6 (2017A), alloy steel S355JOWP (in past called
10HNAP) and cast iron GGG 40.
For the analysis authors used three different criteria, which are based on the concept of a
critical plane. Coefficients used in the expressions for the equivalent stresses are calculated on
the basis of classical fatigue limits. These are the criteria: the maximum normal and shear
stresses proposed by Macha [1] and the criterion of Carpintieri and Spagnoli [2], where the
critical angle of the plane is increased by the angle
3    af
  1  
2    af





2

 45 ,


(1)
relative to the angle defined by the maximum normal stress.
General form of the equivalent stress according to the proposed criteria can be written as
 eq (t )  B s (t )  K  (t ),
(2)
where K and B are constants used to select a particular form of criteria.
Acknowledgement: The project financed from the funds of the National Centre of Science
– decision number 2011/01/B/ST8/06850.
REFERENCES
[1]
[2]
MACHA E.: Generalization of fatigue fracture criteria for multiaxial sinusoidal loadings in the
range of random loadings, in: Biaxial and Multiaxial Fatigue, EGF 3 (Edited by M.W. Brown
and K.J. Miller), Mechanical Engineering Publications, London 1989, pp. 425–436.
CARPINTERI A., SPAGNOLI A.: Multiaxial high–cycle fatigue criterion for hard metals, Int J
Fatigue 23, 2001, pp. 135–145.
45
46
EVALUATION OF FATIGUE CRACK GROWTH IN ALPHA TITANIUM
ALLOYS
O. UMEZAWA1*, M. HAMADA2, T. TATSUMI2
1
Faculty of Engineering, Yokohama National University, Hodogaya, Yokohama, 240-8501, Japan;
email: umezawa@ynu.ac.jp
2
Graduate School of Engineering, Yokohama National University, Hodogaya, Yokohama, 240-8501, Japan
KEY WORDS: subsurface crack, fatigue crack propagation, titanium alloy
Subsurface crack generation in the high-cycle fatigue of -titanium alloys is dominant at
lower stress regime and lower temperature [1]. The crack initiation sites appear at
crystallographic facets such as the (0001) transgranular cracking [2]. The fatigue crack growth
modelling that based on linear fracture mechanics under the Mode I condition provided a good
estimate of the stress intensity range of subsurface or surface and fatigue crack growth, enabling
the estimation of the crack propagation life for Ti-6Al-4V alloys [3]. The fatigue crack growth
rate calculated using the Paris rule, da/dN=C(KI)m, almost corresponded to the one obtained
from the analysis of the striation on the fracture surface. The calculated crack propagation life
was less than a tenth of the number of cycles to failure over 106. As a result, the subsurface
crack initiation (Stage I crack generation) process consumed a large number of cycles to failure.
In the present study, this evaluation was applied to  or near -type titanium alloys with various
morphologies.
Three kinds of commercially pure titanium, i.e. CPTi JIS type 1 (O: 0.10 mass%),
2 (O: 0.13 mass%) and 3 (0.20 mass%), and four kinds of near α-type Ti-Fe-O (Fe: 0.994,
O: 0.386 mass%) materials, i.e. T specimen (parallel to transverse direction (TD)), L specimen
(parallel to rolling direction (RD)), CR specimen (cross-rolling) and CS specimen (groovedrolling) were used. Each material was hot-rolled and annealed. The L, T and CR materials
showed the pancaked  grains elongated in both RD and TD, which were aligned in prior 
grain. The CR showed equiaxed grains with {hkl}<110> fibre texture. Force-controlling tests
were done at 77 K and 293 K using a servo-hydraulic fatigue machine. The sinusoidal
waveform forcing was uniaxial with a minimum-to-maximum stress ratio, R (σmin/σmax), of
0.01. The fractured specimens showing subsurface fatigue crack initiation were chosen for.
Fatigue crack initiation sites and fracture surfaces were analysed by scanning electron
microscopy. In Ti-Fe-O materials, the planes of aligned facets in an initiation site showed
almost the same inclination against the principal stress axis and microcracking and/or its growth
[4]. Then the initiation site was approximated for an ellipse as initial crack [3]. The crack length,
2a, crack width, 2c, and distance from specimen surface to centre of ellipse, d, were determined,
where the direction of crack length was parallel to the initial crack propagating direction.
Maximum fatigue crack size in Stage II was taken from the ripple mark on fracture surface.
Striation marks on propagating plane were characterized experimentally to examine crack
growth rate, da/dN. A software system “SCAN” based on linear fracture mechanics was
adopted in the system for investigating subsurface crack [5], and then the crack size resulted
from its position gave the stress intensity factor range, KImax = KImax-KImin, using the SCAN.
Figure 1 shows da/dN-KImax relationship for Ti-Fe-O alloy failed at 293 K. Although the
crack growth rate of the L, T and CR were scattered because of their aligned  grain
microstructure, their crack growth rates were approximately the same. The CS showed higher
the crack growth rate than the others. The crack growth rates among three CPTi alloys were
also almost the same where no influence of strength on the rate was detected.
47
Based on the da/dN-KI relationship, fatigue propagation cycle, Np, was calculated. The
Np was a smaller from a tenth to a hundredth than the number of cycle to failure, Nf. As a result,
the subsurface crack initiation process also consumed a large number of cycles to failure in 
titanium alloys as well as Ti-6Al-4V alloys.
10 -2
CS
da/dN (mm/cycle)
L
CR
T
CR
CS
10 -3
L
T
10 -4
10
100
KI max (MPa m)
Fig. 1. Relationship between da/dN and ΔKImax for Ti-Fe-O alloys failed at 293 K.
REFERENCES
[1]
[2]
[3]
[4]
[5]
UMEZAWA, O., NAGAI, K.: Subsurface Crack Generation in High-cycle Fatigue for High
Strength Alloys, ISIJ International 37, 1997, pp. 1170-1179.
YOKOYAMA, H., UMEZAWA, O., NAGAI, K., SUZUKI, T., KOKUBO, K.: Finite Element
Analysis of Composite Materials, Boca Raton: CRC Press 2008. Cyclic Deformation,
Dislocation Structure and Internal Fatigue Crack Generation in Ti-Fe-O Alloy at Liquid
Nitrogen Temperature, Metallurgical Materials Transactions A 31A, 2000, pp. 2793-2805.
HAMADA, M., UMEZAWA, O.: Evaluation of Subsurface Fatigue Crack Life in Forged Ti6Al-4V Alloys at Cryogenic Temperatures, ISIJ International 49, 2009, pp. 124-131.
MORITA, M., UMEZAWA, O.: Slip Deformation Analysis Based on Full Constraints Model
for -Titanium Alloy at Low Temperature, Materials Transactions 52, 2011, pp. 1595-1602.
NAKANISHI, S., IWAMATSU, F., SHIRATORI, M., MATSUSHITA, H.: Estimation of
Fatigue Crack Propagation of Subsurface Cracks by SCAN, Proceedings of the 2006 ASME
Pressure Vessels and Piping Conference PVP2006-ICPVT-11-93248, 2006.
48
THE CURRENT STATUS OF NEW CZECH CORROSION FATIGUE
EVALUATION PROPOSAL FOR WWER NUCLEAR POWER PLANTS
L. VLČEK1*
1
Institute of Applied Mechanics Brno, Ltd., Resslova 972/3, Veveří, 602 00 Brno; email: vlcekl@uam.cz
KEY WORDS: low-cycle fatigue (LCF), Czech NTD A.M.E. standard, air and primary water
environment, nuclear power plant, WWER
Presented paper introduces an innovative principle of fatigue life assessment suggested for
WWER nuclear power plants. The subject of this work is to take into account the corrosion
environment influence in actual methodology of low-cycle fatigue assessment and prediction.
Due to Czech nuclear power plants operator requirement the project focused on base steel
materials, which are used in primary circuit of WWER-440, started in 2010. The basic idea of
Czech environmental fatigue correction factor has been introduced on international PVP
conference in 2013 [1]. The new project linked to the previous one is focused on the additional
area of welding joints. The aim of this paper is to summarize the current status of the Czech
proposal of corrosion fatigue assessment and prediction. Assessment procedures used for
fatigue life evaluation are stated in NTD A.M.E. standard [2]. The purpose is to take into
account the influence of primary water corrosion environment on fatigue life of components
and piping. The decrease of fatigue life due to primary water environment is generally realized
by so called fatigue life environmental correction factor. Such correction factor was originally
introduced in NUREG documents [e.g. 3] as a ratio of fatigue life in air at reference temperature
conditions to fatigue life in water at operating temperature conditions (FEN = Nair, RT / Nwater).
Such way defined environmental correction factor can’t be directly used for fatigue life
assessment and prediction under operating conditions of WWER nuclear power plants. Reasons
are lying on the side of different way of fatigue life assessment and prediction, which is used
on the WWER power plants. Therefore the redefinition of environmental correction factor FPR
was introduced as a ratio of total strain amplitude in air at operating temperature condition to
total strain amplitude in water at operating temperature condition [4]:
FPR   at air  at water ,
(1)
where at air is total strain amplitude in air at operating temperature, at water is total strain
amplitude in water at the same operating temperature as at air.
Based on the new definition there were constructed dependencies of total strain amplitude
vs. environmental correction factor FPR. Environmental correction factor is related to the total
strain coming not from fatigue design curve, but from fatigue curve without the application of
safety factors on stress n = 2 and number of cycles nN =10. Dependencies at air vs. FPR were
constructed for the case of minimal (theoretical) influence and maximal (theoretical) influence
of primary water environment on fatigue life. Theoretical minimal and maximal influence of
corrosion environment on fatigue life is covered by design fatigue curves (so called S-N curves)
proposed by Russian authors [5]. With the aim of direct application in the frame of actual
mathematical description, which describe relations of S-N design curves, the coefficient for
water corrosion environment PR can be defined as the reciprocal value of FPR (PR = 1/FPR).
The project is completed by experimental verifications of proposed environmental
correction factor. Experimental work is based on LCF strain-controlled tests in primary water
environment of WWER-440. In the frame of finished project the dependency at air vs. FPR was
verified for base material, which is austenitic stainless steel 08CH18N10T (AISI 321). LCF
49
tests done in corrosion environment of primary water demonstrated that the use of theoretical
maximal correction proposed for austenitic stainless steel seems to be unfounded. Much closer
to the reality of fatigue life reduction due to effect of corrosion environment for base metal
08CH18N10T is the minimal theoretical proposal of correction.
General proposal of fatigue life assessment and prediction in water environment covers not
only the base steel materials, but also their welding joints. The subject of actually running
theoretical-experimental program covers similar metal welds of austenitic stainless steel
08CH18N10T. Moreover LCF tests in corrosion environment of dissimilar metal welds are
under preparation.
REFERENCES
[1]
[2]
[3]
[4]
[5]
VLČEK, L.: Corrosion Fatigue Evaluation of Austenitic Stainless Steels: the New Proposal to
the Czech Standard in the Area of Nuclear Power Plants Type WWER, PVP 2013, July 14-18,
Paris, France.
NTD A.M.E. standard: Normatively Technical Documentation of Association of Mechanical
Engineers, Section III, Strength assessment of equipment and piping of nuclear power plant type
WWER, NTD_ASI_Sekce_III_2013, č. 1 (in Czech).
Effect of LWR Coolant Environments on the Fatigue Life of Reactor Materials, Final Report,
Argonne National Laboratory, U.S. Nuclear Regulatory Commission Office of Nuclear
Regulatory Research Washington, DC 20555-0001, NUREG/CR-6909, ANL-06/08, February
2007.
VLČEK, L.: Fatigue life assessment of equipment under corrosion environment conditions,
stage I: Analysis of Russian Environmental fatigue Correction of Fatigue Design Curves and
Comparison with American Approach, Proposal of Environmental Correction Factor, IAM Brno
report No. 4736/10, Brno, 2010. (in Czech).
FILATOV, V. M., EVROPIN, S. V., Strength calculation of NPP equipment and pipelines
during operation. Low- and high-cycle corrosion fatigue, International Journal of Pressure
Vessels and Piping, Vol. 81, 2004.
50
EFFECT OF REPEATED HEATING ON ONE-POINT ROLLING CONTACT
FATIGUE OF HIGH-CARBON HIGH-CHROMIUM STEEL BAR
K. MIZOBE1*, R. SEGAWA1, T. SHIBUKAWA2, K. KIDA1
1
2
University of Toyama, Gofuku 3190, Toyama, 930-8555, Japan; email: kmizobe@eng.u-toyama.ac.jp
YSK Co., Ltd. 3103-6 Kitanokawachi-Otsu Nishimatsuura-gun Arita-Chou Saga 849-4166 Japan
KEY WORDS: single-ball rolling contact fatigue, repeated heating, SUJ2, grain refinement
Bearings are used under severe loadings. Flacking failure which is one of the bearing
fractures occur under contact stress. It is difficult to calculate the driving stress of crack tip
because the subsurface crack propagated under compressed stress field.
Lumdberg and Palmgren [1] found that the lifetime of bearings depended on the subsurface
crack initiation process and crack initiation occurred in late stages of bearing lifetime (106,
107 cycles). The present design criteria of the bearings is based on their concept. However, in
1999, Nélias’ research group [2] discovered that some cracks appeared during the early stages
of bearing lifetime (104, 105 cycles). According to their research, we need to focus on the two
bearing lifetime groups, one is the crack initiation at the low cycle stage (104, 105 cycles) and
the other was the crack propagation at the high cycle stage (106, 107 cycles).
Our research groups [3-4] developed the single-ball rolling contact fatigue (RCF) testing
machine in order to directly observe the crack initiation and propagation under contact stress.
This method has many advantages, such as, the single-ball contact stress, the simple shape of
the specimen and the large observation area. Generally, a bearing specimen of the RCF consists
of a retainer and two races. The retainer which includes three balls is sandwiched with the two
races. Fig. 1 shows the subsurface observation process of the RCF test. The balls rotate along
the upper race groove. The small contact area can be observed by straight cutting line. In the
single-ball RCF, we performed rolling contact fatigue between one ball and one shaft bar.
Single-ball RCF enables to observe the large contact area by sectioning the specimen.
In our previous work, we focused on the effect of repeated quenching on the rolling contact
fatigue. Repeated heating was widely used as the refinement method since Grange [6-7], who
investigated the relation between repeated heating and material strength of low carbon steel. In
the case of high-carbon high-chromium steel (JIS-SUJ2), refining the prior austenite grain size
improves material strength. We have investigated the prior austenite grain refinement of SUJ2
material by repeated heating [8-9]. In this study, we performed the single-ball RCF test of
repeatedly-heated SUJ2 bar and observed cracks originating from the non-metallic inclusions.
Fig. 2 is a schematic illustration showing the single-ball RCF mechanism. Fig. 3 is a photograph
of the device. All tests were performed using 17 mm diameter and 300 mm length SUJ2 shafts,
and 3/8 inch diameter SUJ2 balls. RCF tests were performed at 3000 rpm and maximum
Hertzian contact stress of 5.3 GPa. We prepared once-quenched samples and three-timesquenched samples. After the single-ball RCF tests, cracks originating from the non-metallic
inclusions on the specimen’s cross section were observed by using a KeyenceVK9700 laser
confocal microscope. Because the ball always run on the same line, the inner stress was
analysed by Sackfield and Hills’s method [5].
51
Rotation
Sectioning
Observation
Fig. 1. Schematic illustration of subsurface observation under RCF.
Fig. 2. Schematic illustration of single-ball RCF mechanism and
specimen sectioning for subsurface observation.
Fig. 3. Single-ball RCF testing apparatus.
Acknowledgement: This work was supported by Grant-in-Aid for JSPS Fellows 25-4761.
REFERENCES
[1]
[2]
[3]
[4]
[5]
[6]
[7]
[8]
[9]
LUNDBERG, G., PALMGREN, A.: Dynamic Capacity of Roller Bearings (Generalstabens
litografiska anstalts förlag, Sweden 1947).
NÉLIAS, D., DUMONT, M. L., CHAMPIOT, F., VINCENT, A., GIRODIN, D., FOUGERES,
R. FLAMAND, L.: Journal of Tribology, Vol. 121, No. 2, (1999), pp. 240-251.
ROZWADOWSKA, J., KIDA, K., SANTOS, E. C., HONDA, T., KANEMASU, K.,
HASHIMOTO, K.: Advanced Materials Research, Vols. 418-420, (2011), pp. 1613-1617.
HAZEYAMA, S., ROZWADOWSKA, J., KIDA, K., SANTOS, E. C., HONDA, T.,
KANEMASU, K., SHIBUKAWA, T.: Advanced Materials Research, Vol. 566, (2012),
pp. 182-186.
SACKFIELD, A., HILLS, D. A.: The Journal of Strain Analysis for Engineering Design,
Vol. 18, No. 2, (1983), pp. 101-105.
GRANGE, R. A., SHACKELFORD, E. R.: Method of Producing Fine Grained Steel, 1966, US
patent 3, 278, 345.
GRANGE, R. A.: Effect of microstructural banding in steel, Metallurgical and Materials Trans.
A 2, 1971, pp. 417-426.
MIZOBE, K., SANTOS, E. C., HONDA, T., KOIKE, H., KIDA, K.: Observation of nonmetallic inclusions on repeatedly quenched SAE 52100 bearing steel fracture surfaces,
International journal of materials and product technology 44(3/4), 2012, pp. 227-239.
MIZOBE, K., HONDA, T., KOIKE, H., SANTOS, E. C., SHIBUKAWA, T., KIDA, K.:
Relationship between repeatedly quenching and fisheye cracks around TiN and Al2O3 inclusions
in high carbon bearing steel, Material Research Innovations 18, S1, 2014, pp. S60-S65.
52
STATISTICAL ANALYSIS OF ACCIDENTS DUE TO FATIGUE AND
CORROSION AT FACILITIES PRODUCING HIGH PRESSURE GAS
T. SHIBUTANI1*, N. KASAI1, H. KOBAYASHI2, H. AKATSUKA2, T. TAKAHASHI2,
T. YAMADA2
1
Yokohama National University, 79-5 Tokiwadai, Hodogaya-ku, Yokohama, 240-8501, Japan;
email: shibu@ynu.ac.jp
2
Institute of High Pressure Gas Safety, Hulic Kamiyacho Building, 3-13, Toranomon 4-Chome, Minato-Ku,
Tokyo 105-8447, Japan
KEY WORDS: failure mode analysis, high pressure gas safety act, fatigue, corrosion
Industrial plants, such as chemical plants, oil refinery use large amount of high pressure
gas to produce the product. However, facilities to use high pressure gas generally require high
safety level because they become the deterioration by high pressure or low temperature of the
gas. Moreover, the chemical composition of high pressure gas might cause explosion, fire
accident and corrosion.
In Japan, handling high pressure gas is restricted by High Pressure Gas Safety Act. When
an accident has taken place with respect to the high pressure gas, all who handle high pressure
gas shall submit a notification report of the accident to the government. The database of
accidents has been constructed by Institute of High Pressure Gas Safety.
This study provides a statistical analysis of notification reports of accidents from 2008 to
2011. The focus of the analysis is put on accidents at the facilities producing high pressure gas.
Authors have developed a method for classifying accidents of high pressure gas by using a tree
diagram.
Accidental events are classified into explosion, fire, leakage, rupture without the leakage,
and others. Most accidental events are the leakage. The leakage of high pressure gas is classified
into three types: the leakage from the pressurized component, the small leakage from looseness
of bolts, flanges, and valves, and the leakage from other factors such as human factors. As a
result of failure mode analysis, major failure modes of the leakage from the pressurized
component are fatigue and corrosion.
Accidents due to fatigue occur at cold evaporator (CE), compressed natural gas (CNG)
stand, and refrigeration equipment. Detail analysis revealed that fatigue cracking takes place at
brazing joints of pure copper (C1220) tubes at CE and refrigeration equipment. Copper tubes
are often used since copper has an excellent thermal conductivity. However, no fatigue design
is considered for brazing joints on those heat exchangers. Temperature changing or vibration
from a compressor may cause fatigue cracking at brazing joints of copper tubes.
At CNG stands, fatigue accidents occur around the compressor. In particular, some
accidents are related to flexible tubes. The flexible tube is used for an imperfect alignment of
pipes or absorbing displacements. If the flexible tube is used to absorb the vibration of
compressor, it shall be designed to prevent the fatigue cracking. However, many flexible tubes
at CNG stands are used without the fatigue design.
As for corrosion, large number of corrosion of carbon steel under insulator were caused by
the high temperature and wet environment, and the existence of chloride ion. The increased
number of corrosion on nozzles and the pipes in small diameter has been brought because the
parts are generally not paid attention at all.
53
Since many facilities are old, it has been misunderstood that many accidents are caused by
aging. However, the detail analysis in this study pointed out that fatigue accidents are caused
by design error (the lack of fatigue design). Also, corrosion accidents shall be prevented by
appropriate inspection plan (the lack of corrosion management).
Acknowledgement: The authors would like to thank T. SANO, and Y. Ueda for their support
for accident data arrangement.
54
INFLUENCE OF INDUCTIVE HARDENING ON WEAR RESISTANCE IN
CASE OF ROLLING CONTACT
M. ŠOFER1*, R. FAJKOŠ1, R. HALAMA1
1
VŠB-TU Ostrava, 17. listopadu 15/2127, 708 33 Ostrava-Poruba, Czech Republic;
email: michal.sofer@vsb.cz
KEY WORDS: rolling contact fatigue, wear, ratcheting, R8T
The main aim of presented paper is to show how heat treatment, in our case inductive
hardening, will affect the wear rates as well as the ratcheting evolution process below contact
surface in the field of line rolling contact. Used wear model is based on shear band cracking
mechanism [1] and non-linear kinematic and isotropic hardening rule of Chaboche and
Lemaitre. The entire numerical simulations have been realized in C# program. Results from
numerical simulations are subsequently compared with experimental data and metallographic
analysis.
All rolling contact wear tests were
performed in the Rolling Contact Fatigue
Laboratory at the Department of Mechanics
of Materials of VŠB-Technical University
of Ostrava on TUORS testing device [2]. All
eight samples of wheel specimen were made
of R8T steel, whereas the rail specimens
were made of class C steel.
The wheel specimens were organized
into four sets according to the location of
sample´s collection from railway wheel rim
and application of mentioned heat treatment.
The Hertzian contact pressure and the
Fig. 1. TUORS testing device.
creepage were 1200 MPa and 0.75%
respectively. All the wheel specimens realized 105 cycles in total. After each wear test, the
weight and diameter loss of the wheel specimen have been measured. The findings from
metallographic analysis of wheel specimen after realized 105 cycles were subsequently used in
the evaluation process of performed numerical study.
Exploited wear model uses shear band cracking mechanism, which is capable of predicting
the wear process according to the accumulation of plastic shear strain below contact surface.
For ratcheting prediction in particular depths below contact surface and in case of rolling/sliding
two-dimensional contact, the authors have used a non-linear kinematic hardening rule,
introduced by Lemaitre and Chaboche [3]. The authors in the numerical simulations also took
into account the variability of friction coefficient, which significantly influences the evolution
of ratcheting in early stage of the experiment.
The main aim was to compare experimentally and numerically obtained results with respect
to plastic shear deformation profile in the active material layer and the values of wear rates after
specified number of cycles. Relatively good conformity was found between these two
approaches.
Acknowledgement: This work has been elaborated in the framework of the project
Opportunity for young researchers, reg. no. CZ.1.07/2.3.00/30.0016, supported by Operational
55
Programme Education for Competitiveness and co-financed by the European Social Fund and
the state budget of the Czech Republic and in the framework of the IT4Innovations Centre of
Excellence project, reg. no. CZ.1.05/1.1.00/02.0070, supported by Operational Programme
Research and Development for Innovations and funded from the Structural Funds of the
European Union.
REFERENCES
[1]
[2]
[3]
MAZZU, A.: A simplified non-linear kinematic hardening model for ratcheting and wear
assessment in rolling contact, Journal of Strain Analysis 43, 2008, pp. 349–360.
HALAMA, R., FAJKOŠ, R., MATUŠEK, P., BÁBKOVÁ, P., FOJTÍK, F., VÁCLAVEK, L.:
Contact defects initiation in railroad wheels – Experience, experiments and modelling, Wear
271, 2011, pp. 174-185.
LEMAITRE, J., CHABOCHE, J., L.: Mechanics of solid materials, Cambridge University
Press, Cambridge, 1994.
56
EFFECT OF SURFACE QUALITY OF MACHINED RAILWAY WHEELS ON
FATIGUE STRENGTH
R. FAJKOŠ1*, T. TKÁČ2
VŠB-TU Ostrava, 17. listopadu 15/2127, 708 33 Ostrava-Poruba, Czech Republic;
email: rostislav.fajkos@vsb.cz
2
BONATRANS GROUP a.s., Revoluční 1234, 735 94 Bohumín, Czech Republic
1
KEY WORDS: fatigue strength, surface layers, machining technologies, shot peening
Railway transport capacities all over the world have been growing, a phenomenon which
is accompanied by the requirement to increase axle loads of freight rolling stock. Apart from
new wheel designs for higher axle loads, growing have been also the requirements on their
safety and reliability, since these wheels are often used in extreme climactic conditions.
Cruising speeds of passenger trains have also been increasing, which likewise brings more
stringent requirements concerning the quality and safety of the supplied railway wheels.
This paper describes methods of
evaluating fatigue strength of railway
wheel webs and methods of evaluating
the quality of machined railway wheel
webs. Results of fatigue tests performed
on wheels machined in a standard way
are compared with wheels which have
been treated by shot peening, a
treatment frequently used to increase
the fatigue strength of wheel webs of
railway wheelset.
The principle of a fatigue test of
railway wheels is checking whether the
supplied wheels meet the parameters
defined in standard EN 13 262, i.e.
whether they can withstand 10 million
cycles with the test level of radial stress
amplitude set to 240 MPa at the critical
point. Schematically, this type of test is
carried out at BONATRANS GROUP
a.s., preferably on the electro-hydraulic
test equipment illustrated on 3D model
in Fig. 1 below.
In order for us to be able to qualify
the effect of shot peening and
machining quality on the resultant
fatigue strength of railway wheels on
real scale, the following experiment was
devised. In total four wheel variants
were tested, namely a wheel with an
unmachined web, a wheel with a
Fig. 1. A 3D model of the electro-hydraulic test equipment
used for fatigue strength tests of railway wheels.
Fig. 2. Comparison of stress levels of railway wheels with
different final finishing of the wheel web.
57
machined web, a wheel with an unmachined but shot peened web, and a wheel with a machined
and shot peened web.
The results of the tests for each of the above wheels with different machining technologies
and shot peening are for comparison purposes presented in Fig. 2. All the tests were carried out
on an Inova electro-hydraulic fatigue strength test machine at BONATRANS GROUP a.s.
The fatigue strength of wheels manufactured by BONATRANS GROUP a.s. is around
300 MPa, which provides an adequate spare strength capacity when conducting fatigue strength
tests at the stress amplitude level of 240 MPa required by the standard.
Wheels made from steel grades with a higher content of C (grades ER8, Class B and other),
are basically even better off because of the higher strength of their normalised structure which
develop in wheel web with higher content of C. However, this at least a 25% spare strength
capacity is not enough if the quality of the surface machining is substandard. If because of tool
post vibrations, or because of using a blunt cutting tool, or because of similar technological
shortcomings, fissures develop in the cut surface, the fatigue strength of such product decreases
rapidly.
To test the real fatigue strength of wheels machined using different technologies, designed
were flat bars. The designed shape of the test bodies allowed us to better capture the character
of stresses in the given part of the wheel, and at the same time enabled us to collect such bars
from the surface of a wheel with straight or only gently sloping fixed web. The width of the test
bar in the area of the fatigue failure was 24 mm, and the thickness of the sample was 12 mm.
Three variants of final surface treatment of the test samples collected from a wheel web were
selected for the experiment which will be described on full article.
Acknowledgement: This article has been elaborated in the framework of the project
Opportunity for young researchers, reg. no. CZ.1.07/2.3.00/30.0016, supported by Operational
Programme Education for competitiveness and co-financed by the European Social Fund and
the state budget of the Czech Republic.
REFERENCES
[1]
[2]
[3]
[4]
[5]
[6]
[7]
[8]
[9]
ČSN EN 13262+A1: Railway applications–Wheelsets and bogies–Wheels–Product
requirements, May 2009.
AAR M-107/M208.: AAR Manual of Standards and Recommended Practices, Wheels and
Axles, 2011.
OKAGATA, Y., KIRIYAMA, K., KOAT, T.: Fatigue strength evaluation of the Japanese
railway Wheel, Fatigue Fact. Engng. Mater Struct 30, 356-371.
STRNADEL, B.: Material Science II, Material degradation processes and design, Mining
Academy–Technical University of Ostrava, Ostrava 2008.
MORAVEC, V.: Hardness and durability of dynamically loaded machine parts, Mining
Academy–Technical University of Ostrava, Ostrava 2007.
KLESNIL, M., LUKÁŠ, P.: Fatigue at metal materials, Academia, Prague 1975, 222.
BERETTA, S., CARBONI, M., LO CONTE, A., REGAZZI, D., TRASATTI, S., RIZZI, M.:
Crack Growth Studies in Railway Axles under Corrosion.
BERETTA, S., CARBONI, M., FIORE, G., LO CONTE, A.: Corrosion–fatigue of A1N railway
axle steel exposed to rainwater, International Journal of Fatigue 32 (2010) 952–961.
LUKÁŠ, P., KUNZ, L., WEISS, B., STICKLER, R.: Impact of short cracks and minor notches
on fatigue strength, Metal materials 5, 26, Bratislava 1988.
58
APPLICATION OF ULTRASONIC IMPACT TREATMENT (UIT) FOR
IMPROVEMENT OF FATIGUE LIFE
T. ISHIKAWA1*, K. HAYASHI2
1
Nippon Steel & Sumikin Technology Co., LTD., 1-7-1 Yurakucho, Chiyoda-ku, Tokyo, 100-0006, Japan;
email: ishikawa-tadashi@nsst.jp
2
Nippon Steel & Sumikin Technology Co., LTD., KSP A101, 3-2-1 Sakado, Takatsu-ku, Kawasaki,
Kanagawa, 213-0021, Japan
KEY WORDS: fatigue, residual stress, S-N curve, fatigue crack, repair, stress, welded joint
It is an important subject to prevent fatigue crack initiation for the structural integrity of
welded structural parts especially under cyclic loading. Post-weld treatment methods, such as
grinding, tungsten inert-gas (TIG) dressing, hammer peening, ultrasonic impact treatment (UIT)
etc., are applied to welded-toes as improving procedures against the fatigue of weld joints. The
UIT is one of the most powerful solutions for this subject, because of high improved fatigue
performance with high productivity and high durability in construction stage.
The UIT equipment consists of vibration exciter and transducer & horn, as shown in Fig. 1.
The procedure of UIT application to the welded-toe is shown in Fig. 2. The pins impact the
welded-toe with 25-30 m amplitude of 25-27 kHz vibration. The welded-toe is locally
deformed and reformed to the continuously smooth toe which can be confirmed to the eye.
Pin
amp: 25–30 
impact by ultrasonic
welded toe
Vibration exciter
(25–27 kHz)
Transducer & Horn
Fig. 1. Tool of Ultrasonic Impact Treatment.
weld
base plate metal
Before
Continuously smooth
Welded toe formed
pin
Applying UIT
After
Fig. 2. Process of UIT application to welded toe.
Figure 3 shows the example of S-N (stress versus number of loading cycles to failure)
curves of cruciform welded joint of as-welded, toe grinding, and UIT applied conditions [1].
As shown in Fig. 3, UIT can extend fatigue life ten times or more than the as-welded condition.
The mechanism of fatigue initiation property improvement by UIT is summarized in Fig. 4.
Root radius of welded toe increases from naturally formed shape (0.5-1 mm) to 3 mm as same
as pin head radius. It gives smaller stress concentration at the welded-toe. UIT also induces
compressive residual stress up to the yield strength level of steel, as shown in Fig. 5.
Furthermore, the microstructure near the welded-toe becomes ultra-fine grains by local heavy
plastic deformation.
UIT has been widely utilized in ships, bridges, earth moving equipment, crane garters, and
so on. Welded-toe grinding is well-known as improvement methods of fatigue life. For the shipbuilding use, UIT has been certified by ship classifications societies as the alternative method
without any dust and less noise.
As the superior improved fatigue performance with UIT application [1, 2, 3], new S-N
curve with UIT in design codes or standards has been strongly demanded. Current guidance on
59
improved fatigue life prediction with hammer-peened welds, together with a modification to
the S-N curves is recently published in RP-C203
Residual Stress ahead of
welded toe, MPa
600.0
as welded
UIT treated
400.0
200.0
UIT treated
0.0
0
2
8
4
10
12
14
16
6
-200.0
Distance from welded toe [mm]
-400.0
-600.0
Fig. 3. Example of S-N curves of Welded joint [1].
Fig. 4. Residual stress distribution ahead of welded toe.
Fatigue Design of Offshore Steel Structures (2011) published by DNV ship classification
society. The S-N curve with UIT application can be referred as equivalent as ones with hammerpeened welds, and may be expected more benefit in near future. In Japan, UIT has been widely
used for the bridge construction, and listed as recommended new technology of NEITES
authorized by Land, Infrastructure and Transportation Ministry.
Before
Radius
～0.5mm
After
WM
Radius
3mm
Ultra-fine region
Tensile R.S.
BM
HAZ
・Root radius：
sharp （0.5～1.0ｍｍ）
⇒ Stress concentration
・Residual Stress：
Large tensile Res. Stress
⇒ increase appl,stress
・Microstructures：
Coarse grains（ 20-50 μｍ）
⇒ Locally low strength
BM
Comparison of
溶接金属
Comp.Res.Stress
HAZ
１mm
Increase to 3.0ｍｍ（as same as pin-tip)
⇒ Decrease Stress Concentration
Compressive Res.Stress by Plastic flow
⇒ Decrease appl.stress
fatigue crack
Initiation
properties
improving effect
UIT Grinder LTT
TIG
○
◎
－
◎
－
◎
Ultra-fine grains (1μｍ）
△
－
－
⇒ increase strength
(LTT is Low Temperature Transformation welding consumables)
Fig. 5. Mechanisms of improvement of fatigue crack initiation properties by UIT.
REFERENCES
[1]
[2]
[3]
NOSE, T., OKAWA, T.: Approaches for Fundamental Principles 2: Total Solution for Fatigue
of Steel, Nippon Steel Technical Report, No.391, 2011, pp. 156-161.
SHIMANUKI, H., NOSE, T.: Effect of Ultrasonic Impact Treatment on Fatigue Properties of
Structural Model, Proceeding of Japan Welding Society Vol. 81, 2007, No. 342.
KAYAMORI,Y.,et.al.: Applicability of fatigue solutions to floating wind turbine structures,
International Symp. on Marine and Offshore Renewable Energy, 2013.
60
CYCLIC PLASTIC PROPERTIES OF CLASS C STEEL INCLUDING
RATCHETING: TESTING AND MODELLING
R. HALAMA1*, A. MARKOPOULOS1, M. ŠOFER1, P. MATUŠEK2
VŠB-Technical University of Ostrava, Department of Mechanics of Materials, Centre of Excellence
IT4Innovations, 17.listopadu 15, Ostrava, Czech Republic; email: radim.halama@vsb.cz
2
Bonatrans Group, Bohumín, Revoluční 1234, 735 94, Czech Republic
1
KEY WORDS: cyclic plasticity, ratcheting, FEM, low-cycle fatigue
Cyclic plasticity modelling of metals needs individual approach. There are specific theories
for various metallic materials including mainly phenomenological models useful for practical
applications [1].
This paper is focused on the stress-strain behavior of the Class C steel in cyclic plastic
domain and its FE simulation. An experimental study on the wheel steel specimens including
uniaxial as well as multiaxial tests has been realized in the laboratory at Department of
mechanics of materials of VŠB-TU Ostrava.
The main attention in this study was paid to study
ratcheting under nonproportional loading. The
specimens were subjected to tension-torsion tests on
the
reconstructed
test
machine
INOVA
100 kN/1000 Nm (Fig. 1) as in the previous study
performed on ST52 steel [2].
The extensometer EPSILON 3550 with 25.4 mm
gauge length was used to measure axial strain and
shear strain simultaneously. The testing specimen has
tubular testing part with outer diameter of 12.5 mm
and with inner diameter of 10 mm. The specimen was
used also for the case of uniaxial loading.
The uniaxial multistep test was performed under
strain rate of 0.01 s-1. A cyclically stable behavior of
the steel under higher amplitude loading was
observed in the uniaxial multistep test, see Fig. 2.
Fig. 1. Biaxial fatigue testing machine.
The load path in the tension/compression-torsion
tests was applied in accordance with McDowell’s
experiments [3] to obtain similar stress-strain history
as in a point on the surface under rolling-sliding line
contact case. All multiaxial force controlled tests
were realized under sinusoidal wave loading with
frequency of 0.1 Hz.
As a sample, results from the multiaxial
ratcheting test, which was obtained by the symmetric
tension/compression and by repeated torsion, are
presented at the Fig. 3. The case with axial stress
magnitude of 700 MPa and shear stress magnitude of
400 MPa was realized for the wheel steel Class C. As
the consequence of the repeated torsion applied to the
61
Fig. 2. Results of push-pull multistep test.
specimen the increase of the shear strain occurs cycle by cycle in the same direction as the
torque is applied.
The MAKOC model [4], which is based on AbdelKarim-Ohno kinematic hardening rule
and Calloch isotropic hardening rule, has been applied in subsequent finite element simulations.
The numerical results show very good prediction of stress-strain behaviour of the wheel steel.
3 xy
3 a
a
x
Fig. 3. Results of multiaxial ratcheting test.
Acknowledgement: This work was supported by the European Regional Development
Fund in the IT4Innovations Centre of Excellence project (CZ.1.05/1.1.00/02.0070) and by the
OPVK project Opportunity for young researchers (CZ.1.07/2.3.00/30.0016) co-financed by the
ESF.
REFERENCES
[1]
[2]
[3]
[4]
HALAMA, R., SEDLÁK, J., ŠOFER, M.: Phenomenological Modelling of Cyclic Plasticity,
Chapter in: P. Miidla (Ed.), Numerical Modelling, InTech, Rijeka, 2012, pp. 329-354.
HALAMA, R., FOJTÍK, F., MARKOPOULOS, A.: Memorization and Other Transient Effects
of ST52 Steel and Its FE Description, Applied Mechanics and Materials 486, 2013, pp. 48-53.
MCDOWELL, D.L.: Stress state dependence of cyclic ratchetting behaviour of two rail steels.
International Journal of Plasticity 11, 1995, pp. 397-421.
HALAMA, R., ŠOFER, M., FOJTÍK, F.: Choice and Calibration of Cyclic Plasticity Model
with Regard to Subsequent Fatigue Analysis. Engineering Mechanics 19, 2012, pp. 87-97.
62
CHARACTERIZATION OF VERMICULITE PARTICLES AFTER
MECHANICAL TREATMENT
K. ČECH BARABASZOVÁ1,2*, G. SIMHA MARTYNKOVÁ1,2
Nanotechnology Centre, VŠB-TU of Ostrava, 17. listopadu 15/2172, Ostrava-Poruba, Czech Republic;
email: karla.barabaszova@vsb.cz
2
IT4 Innovations Centre of Excellence, VŠB-TU of Ostrava, 17. listopadu 15/2172, Ostrava-Poruba,
Czech Republic
1
KEY WORDS: vermiculite, atomic force microscopy, particles morphology, surface and size
The vermiculite particles are used increasingly for new functional materials. They are
strong contenders for use in polymer nanocomposites. There are many applications of
vermiculite particles as fillers (such as biopolymers nanocomposites [1, 2], lightweight additive
[3], catalyst [4], isolation, ceramics [5] etc.) since the material is natural, inexpensive and
relatively non-harmful for surrounding.
For many of applications is very important the input processing of vermiculite particles.
The vermiculite particles are normally carried out in energy intensive grinding mills such as
planetary mill, oscillating mill and jet mill. Short-time grinding of vermiculite particles requires
to the particle size reduction. But extended grinding lead to an intense structural degradation of
the lamellar shape, lateral size and particle thickness reduction and progressive amorphization
accompanied with formation of hard agglomerates. The changes of the structure and vermiculite
particle size has an influence on the properties of new nanocomposite materials.
Fig. 1. SEM images of the vermiculite particles after ball (VBb) and jet (VBj) milling.
The natural vermiculite particles from Brazil were grinding in jet (VBj) and ball (VBb)
mills. The shape of vermiculite particles has been studied using scanning electron microscopy
(SEM) and atomic force microscopy (AFM). The particles size (PS) changes were characterized
by the median particle size (d50), volume-weighted mean diameter (d43), mode diameter (dm)
and span value.
In Fig. 1 we can see that the VBb particles have the form of platelets with smooth surfaces
with the fact that individual particles showed sharp edges. After jet milling particles (VBj)
showed rounded and corrugated edges. The scanned data from AFM pictures were used for
description of size and thickness of the individual vermiculite particles. With the help of
topographic profiles were measured on the particles parameters of major length and width as
perpendicular profiles (Fig. 2).
63
Fig. 2. 3D AFM images of the vermiculite particles after ball (VBb) and jet (VBj) milling.
From PS parameters was found that grinding of original samples caused reduction of
particles (from 40 µm) after ball milling (VBb), d50 and d43 to values approx. 12 µm together
with the span value to the 1.5. The PS distribution curves had modal character.
Acknowledgement: This work was supported by the project No. SP2014/39 - Functional
nanostructured materials and CZ.1.05/1.1.00/02.0070 - IT for Innovations Centre of Excellence
project.
REFERENCES
[1]
[2]
[3]
[4]
[5]
ZHANG, K., XU, J., WANG, K.Y., CHENG, L., WANG, J., LIU, B.: Preparation and
characterization of chitosan nanocomposites with vermiculite of different modification. Polymer
Degradation and Stability 94, 2009, pp. 2121-2127.
GRYČOVÁ, E., ČECH BARABASZOVÁ, K.: Antibacterial properties of nanostructured
materials. Journal of Nanocomposites and Nanoceramics 3(1), 2012, pp. 7-13.
LING, J., DAI, B.: TiO2 activation using acid-treated vermiculite as a support: characteristics
and photoreactivity. Applied Surface Science 258, 2012, pp. 3386-3392.
QIUQIANG, CH., WU, P. DANG, Z., ZHU, N., LI, P., WU, J., WANG, X.: Iron pillared
vermiculite as a heterogeneous photo-Fenton catalyst for photocatalytic degradation of azo dye
reactive brilliant orange X-GN. Separation and Purification Technology 71, 2010, pp. 315-323.
VALÁŠKOVÁ, M., SIMHA MARTYNKOVÁ, G., SMETANA, B., ŠTUDENTOVÁ, S.:
Influence of vermiculite on the formation of porous cordierites. Applied Clay Science 46, 2009,
pp. 196-201.
64
ADVANCED NUMERICAL MODELLING METHODS FOR 1D PERIODIC
PLASMONIC STRUCTURE SIMMULATIONS
L. HALAGAČKA1,2*, K. POSTAVA1, M. VANWOLLEGHEM3, B. DAGENS2, J. BEN YOUSSEF4,
J. PIŠTORA1
1
Department of Physics and Nanotechnology Centre, Technical University of Ostrava, 17. listopadu 15,
708 33 Ostrava-Poruba, Czech Republic; email: lukas.halagacka@vsb.cz
2
Institut d'Electronique Fondamentale, UMR CNRS 8622, Universite Paris-Sud XI, Orsay, France
3
Institut d'Electronique, Microelectronique et Nanotechnologie, CNRS UMR 8520, Villeneuve-d'Ascq,
France
4
Laboratoire de Magnétisme de Bretagne, Universit de Bretagne Occidentale, EA 4522/CNRS, Brest, France
KEY WORDS: plasmonics, RCWA, magnetooptics, modeling
The RCWA method is well known approach for simulation of optical response of periodic
structures [1]. Depending on a type of simulated structure and required precision, the spectral
simulations could be time consuming. The reduction of computing time is than essential issue
in advanced simulations like a parameter sweep, simulations of thickness inhomogenity,
depolarization effects, etc. Moreover, the reduction is crucial point in data fitting with a model
where a model is recalculated over optimization algorithm.
In this paper we present our implementation of parallel RCWA method in MATLAB for
1D gratings. Our parallel RCWA split an initial spectral problem into series of individual and
independent problems. Those are solved in parallel by slaves and results are collected back by
master. The parallelization is shown on left subplot of Fig. 1 schematically. The Right subplot
shows comparison between ideal linear scaling and measured scaling of our code. A linear
scaling was achieved up to 256 CPUs.
Scalability of spectral problem
250
O(n)
200
parRCWA
Problem definition:
geometry, materials,
spectral domain,...
distribution
to workers
subproblem
#1
subproblem
#2
.......
1/
150
100
results
collection
50
subproblem
# ii
0
Parallel execution
50
100
150
number of CPUs
200
250
Fig. 1. Parallelization of single spectral problem is shown schematically (left). Measured scalability of
implementation is compared with theoretical linear scaling (right).
The performance of the code is demonstrated by fitting of the optical data measured on real
sample with a developed model. The fabricated sample is the 1D periodic gold grating. The
benefit of the gold grating is, that it can support effect of excitation of surface plasmon
polaritons (SPPs). In study of p-reflectivity the excitation of the SPPs appears as a deep sharp
minima [2]. The excitation of SPPs is strongly dependent on the of incidence a beam, therefore
the structure is ideal for study of effect when the incident beam is focused; the angle of
incidence varies around central angle of incidence φ0 over certain interval <φ0-φs ,φ0+φs> .
Since the fabricated structure is not perfect the fitting of the data with model is needed in order
to determine geometry of the structure. In the numerical simulations the focused beam can be
65
approximated with series of plane waves weighted over Gaussian distribution function.
Assumption of 10 partial plane waves is enough for stable numerical simulations. On the other
hand it increases amount of eigenvalue decomposition by factor 10. Moreover, if the bandwidth
of the monochromator is too broad, depolarization occurs due to the wavelength dependence of
the optical properties of a sample. The normalized Gaussian distribution of the wavelengths
around chosen spectral point λ0 with standard deviation σw is assumed [1]. Modeling of the final
spectral resolution requires discretization of the spectral range around λ0 and calculation of the
optical response at all specific wavelengths weighted by corresponding distribution function.
For tabulated optical functions of used material we have used linear spline to obtain proper
values at any wavelength. By numerical test we found, that the use of of only three spectral
points is sufficient to describe depolarization effect from the finite bandwidth. The use of only
three spectral points, namely λ0 - σw, λ0, and λ0 + σw significantly reduces calculation time.
Acknowledgement: Partial support from the projects CZ.1.05/1.1.00/02.0070,
CZ.1.05/2.1.00/01.0040 (RMTVC), CZ.1.07/2.3.00/20.0074 (Nanobase), Czech Science
Foundation 205/11/2137 and SP2013/129 is acknowledged.
REFERENCES
[1]
[2]
[3]
LI, L.: Use of Fourier series in the analysis of discontinuous periodic structures, Journal of
Optical Society of America: A, 13, 1996, pp. 1870-1876.
HALAGAČKA, L., et. al.: Coupled mode enhanced giant magnetoplasmonics transverse Kerr
effect, Optics Express, 21 2013, pp. 21741-21755.
GARCIA CAUREL, E., et. al.: Advanced Mueller ellipsometry instrumentation data analysis.
In Ellipsometry at the nanoscale, Springer, 2013, Engineering, p. 31.
66
MOLECULAR MODELING OF ANTIMICROBIAL NANOCOMPOSITES
D. HLAVÁČ1,2*, J. TOKARSKÝ1,2
IT4Innovations, Centre of Excellence, VŠB-TU Ostrava, 17. listopadu 15/2172, Czech Republic;
email: dominik.hlavac@vsb.cz
2
Nanotechnology Centre, VŠB-TU Ostrava, 17. listopadu 15/2172, Czech Republic
1
KEY WORDS: natural minerals, antimicrobial agents, molecular modeling, adhesion
Nowadays, growing demand of more effective antimicrobial agents is observed in many
areas involving food processing and packaging, water cleaning, hygiene or medicine.
Therefore, except finding new ones, various ways of improvements of existing agents are
searched. One of the most promising ways is the modification of administration (i.e. pure
solution, incorporation into composite material, etc.) especially by their docking on suitable
matrix because controlled release and, therefore, prolonged and more environmentally friendly
activity may be achieved. Wide range of possible matrices varying in effectivity, cost as well
as stability may be used. Natural minerals represent a reasonable choice since they are low-cost
environmentally stable materials. However, because docking capabilities of every mineral
differ from each other (in dependence on the surface area, the host-guest interaction and the
method of preparation), in the first step the selection of best ones may be done according to
knowledge of host-guest interaction. Therefore, docking capabilities of various natural minerals
were investigated using molecular mechanics and dynamics in Materials Studio modeling
environment. Various host matrices were compared according to calculated values of
interaction energies between antimicrobial agents and mineral surfaces. Obtained results were
compared to available experimental data in order to evaluate the possibility of prediction based
on knowledge of host-guest interaction.
Fig. 1. Model structures of antimicrobial agents and natural minerals: a) chlorhexidine, b) nystatin, c) kaolinite,
d) montmorillonite, e) vermiculite.
Acknowledgement: The authors gratefully acknowledge the support by the European
Regional Development Fund in the IT4Innovations Centre of Excellence project
(CZ.1.05/1.1.00/02.0070).
67
68
ANTIMICROBIAL KAOLINITE BASED NANOCOMPOSITES
S. HOLEŠOVÁ1,2*, M. HUNDÁKOVÁ1,2, E. PAZDZIORA3
Nanotechnology centre, VŠB-Technical University of Ostrava, 17. listopadu 15/2172, 708 33 OstravaPoruba, Czech Republic; email: sylva.holesova@vsb.cz
2
IT4Innovations Centre of Excellence, VŠB-Technical University of Ostrava, 17. listopadu 15/2172, 708 33
Ostrava-Poruba, Czech Republic
3
Institute of Public Health Ostrava, Centre of Clinical Laboratories, Partyzánské náměstí 7, 702 00 Ostrava,
Czech Republic
1
KEY WORDS: kaolinite, chlorhexidine, antimicrobial activity
The development of suitable materials with the ability to inhibit the growth of microbes is
one of the current topics of material and medical research. As far as treatment of oral infections
is concerned, the current market lacks any curative form for a local long-acting application that
would enable therapy without need to use systemic treatment. A solution might be offered by
anchoring the drug to a suitable carrier that can provide transport to the designated place in the
body, gradual release and hence suppression of side effects. Recently, increased attention is
paid to so-called inorganic carriers, often based on clay minerals. The use of clay minerals as
excipients in pharmaceutical formulations has been described by many authors [1, 2]. The
antimicrobial nanocomposites based on clay mineral montmorillonite are the most studied
systems. Our team mainly deals with investigation of antimicrobial nanocomposites, when clay
mineral vermiculite is used as a drug carrier [3, 4].
In this study we focused on antimicrobial nanocomposites based on kaolinite, which aren’t
much explored in past. Two series of nanocomposites were prepared. In the first case, kaolinite
(KAO) was used as the carrier for antibacterial drug and in the second case, kaolinite modified
with dimethyl sulfoxide (DMSO) was used. In both series, chlorhexidine dihydrochloride (CH)
acts as an active antimicrobial component. The resultant samples were characterized by X – ray
diffraction (XRD), infrared spectroscopy (IR) and scanning electron microscopy (SEM)
(Fig. 1).
Fig. 1. SEM pictures of KAO (left side) and KAO/DMSO (right side).
The antimicrobial activity of prepared composites against bacteria strains Staphylococcus
aureus, Escherichia coli and against yeast Candida albicans were evaluated by finding
minimum inhibitory concentration (MIC). The dilution and cultivation were performed on the
microtitration plate. Starting dispersion contained 10% (w/v) of nanocomposites and than this
dispersion was diluted by a threefold diluting method to concentrations 3.33%, 1.11%, 0.37%,
0.12%, 0.04% and 0.01%. A volume of 1 µl of glucose suspensions of bacterial strain was
69
added. Antimicrobial activity was monitored after the elapse of 30, 60, 90, 120, 180, 240 and
300 min, and then during 5 days, always in 24 h intervals. The MIC values in selected time
intervals for two bacteria strains and yeast are shown in Table 1.
Table 1 MIC (%) (w/v) values of prepared antimicrobial nanocomposites in selected time intervals
(90 min, 300 min, 1 day, 5 days).
Sample
Staphylococcus aureus
Escherichia coli
Candida albicans
90
300
1
5
90
300
1
5
90
300
1
5
KAO_CH (1:1)
1.11
1.11
0.01
0.01
1.11
1.11
0.01
0.01
0.12
0.12
0.12
0.12
KAO_CH (2:1)
1.11
1.11
0.01
0.01
1.11
1.11
0.01
0.01
0.37
0.12
0.12
0.12
KAO_CH (4:1)
1.11
1.11
0.01
0.01
1.11
1.11
0.01
0.01
0.12
0.04
0.12
0.12
KAO/DMSO_CH (1:1)
0.37
0.37
0.01
0.01
1.11
0.37
0.01
0.01
0.12
0.04
0.12
0.12
KAO/DMSO_CH (2:1)
0.37
0.37
0.01
0.01
1.11
1.11
0.04
0.04
0.12
0.12
0.12
0.12
KAO/DMSO_CH (1:1)
3.33
1.11
0.01
0.01
3.33
1.11
0.01
0.01
0.12
0.12
0.12
0.12
It was found that prepared nanocomposites were very effective and they had different effect
against bacteria strains and yeast. In the case of gram-positive S. aureus we observed very good
efficiency in exposition after 24 h and longer. The MIC values decreased to the lowest
concentration 0.01% w/v. We obtained almost the same results against E. coli. All prepared
samples showed very good efficiency against yeast Candida albicans. We could observe not
only good activity in longer time intervals but the prepared samples, especially with higher CH
concentration, were already very effective at earlier time intervals. Important information was
that treatment with DMSO had not significant effect on antimicrobial activity.
These nanocomposites can be in future used for preparation of drugs for local treatment of
oral cavity with long-acting antimicrobial activity.
Acknowledgement: The authors gratefully acknowledge the support by the project
IT4Innovations Centre of Excellence, reg. no. CZ.1.05/1.1.00/02.0070.
REFERENCES
[1]
[2]
[3]
[4]
CARRETERO, M.I., POZO, M., Clay and non-clay minerals in the pharmaceutical industry
Part I. Excipients and medical applications, Appl. Clay Sci. 46, 2009, pp. 73-80.
AGGUZI, C., CEREZO, P., VISERAS, C., CARAMELLA, C., Use of clays as drug delivery
systems: possibilities and limitations, Appl. Clay Sci. 36, 2007, pp. 22-36.
HOLEŠOVÁ, S., VALÁŠKOVÁ, M., PLEVOVÁ, E., PAZDZIORA, E., MATĚJOVÁ, K.,
Preparation of novel organovermiculites with antibacterial activity using chlorhexidine
diacetate, J. Colloid Interface Sci. 342, 2010, pp. 593-597.
HOLEŠOVÁ, S., SAMLÍKOVÁ, M., VALÁŠKOVÁ, M., PAZDZIORA, E., Antibacterial
activity of organomontmorillonites and organovermiculites prepared using chlorhexidine
diacetate, Appl. Clay Sci. 83-84, 2013, pp. 17-23.
70
VOLATILE ORGANIC MOLECULES SORPTION ONTO CARBON
NANOTUBES
G. SIMHA MARTYNKOVÁ1,2*, D. PLACHÁ2, L. ROZUMOVÁ1,2, E. PLEVOVÁ3
Nanotechnology Centre, VŠB-Technical University of Ostrava, 17. listopadu 15, 708 33 Ostrava-Poruba,
Czech Republic; e-mail: grazyna.simha@vsb.cz
2
IT4Innovations Centre of Excellence, VŠB-Technical University of Ostrava, 17. listopadu 15, 708 33
3
Institute of Geonics AS CR,v.v.i. Ostrava, Studentska 1768, 708 00 Ostrava-Poruba, Czech Republic
1
KEY WORDS: carbon nanotubes, organics, adsorption, X-ray diffraction, molecular modelling
Studying of volatile organic compounds (VOCs) in ambient air and water sample is an
important analytical task because of VOCs’ great contribution to environmental pollution and
their potential threat to human health. The VOCs adsorbed on the adsorbent can be desorbed
by the methods of thermal desorption or solvent extraction, and then analysed using gas
chromatography (GC), or analysed for weight change at temperature using mass spectroscopy
(MS) and thermal gravimetric analysis [1].
Two types of fibrous carbons were studied:
carbon nanofibers and carbon nanotubes. Both
types were purified using acid treatment to remove
non-carbonaceous substances. The adsorption of
formaldehyde, dichlormethane and naphthalene on
the carbons was studied by experimental and
theoretical (molecular simulation) approaches.
Adsorption of organic molecules is the most
intensive at edges, places of defects or doping
atoms sites. Therefore 2 cases of doping were
studied theoretically and so phosphor and boron
atoms using molecular modelling environment of
software Accelerys.
Fig. 1. Molecular model of H2 sorption onto
single wall carbon nanotubes doped with P.
Molecular models and experiment data are in
good agreement proving that higher adsorption has
happened in case of doped nanotubes. Average amount of organic matter was 8wt.% . Structural
changes for full system were observed using analytical methods: X-ray diffraction and infrared
spectroscopy.
Acknowledgement: We are grateful to project CZ.1.05/1.1.00/02.0070 – IT4Innovations
Centre of Excellence for financial support of this work.
REFERENCES
[1]
LI, Q.L., YUAN, D.X., LIN, Q.M.: Evaluation of multi-walled carbon nanotubes as an
adsorbent for trapping volatile organic compounds from environmental samples, Journal of
Chromatography A, 1026 (2004) 283–288.
71
72
OPTICAL MODELLING OF MICROCRYSTALLINE SILICON DEPOSITED
BY PLASMA-ENHANCED CHEMICAL VAPOUR DEPOSITION ON
LOW-COST IRON-NICKEL SUBSTRATE FOR PHOTO-VOLTAIC
APPLICATIONS
Z. MRÁZKOVÁ1,3*, K. POSTAVA2, A. TORRES-RIOS3, M. FOLDYNA3,
P. ROCA I CABARROCAS3, V. VODÁREK4, J. HOLEŠÍNSKÝ4, J. PIŠTORA1
1
Nanotechnology Centre, Technical University of Ostrava, 708 33 Ostrava-Poruba, Czech Republic;
email: zuzana.mrazkova@vsb.cz
2
Department of Physics, Technical University of Ostrava, 708 33 Ostrava-Poruba, Czech Republic
3
LPICM-CNRS, Ecole Polytechnique, 91128 Palaiseau, France
4
Faculty of Metallurgy and Materials Engineering, Technical University of Ostrava, 708 33 Ostrava-Poruba,
Czech Republic
KEY WORDS: in-situ ellipsometry, plasma-enhanced chemical vapour deposition, thin films,
crystalline silicon, solar cells
The ultimate goal of photovoltaic industry is to reduce the price per watt of generated solar
energy to achieve the grid parity. Research in photovoltaics is thus aimed to achieve low-cost
high-efficiency solar cells. The fabrication cost can be reduced by using less expensive
substrates, by deposition of thin silicon layers with good crystallinity, and by using
economically convenient methods of the deposition.
In this work we study thin microcrystalline silicon (μc-Si) films grown on a flexible
low-cost Fe-Ni alloy substrate by a low-temperature (175°) plasma-enhanced chemical vapour
deposition (PECVD) [1, 2]. Since the crystallinity and material quality of the microcrystalline
silicon change during its growth, the deposition results in an inhomogeneous material with a
rather complicated structure. In order to analyse the changing composition of this complex
material the real time spectroscopic ellipsometry has been used. In-situ ellipsometric data taken
at the photon energy from 2.8 to 4.5 eV every 50 seconds enabled us to study the evolution of
crystallinity of the microcrystalline silicon as it grows (shown in Fig. 1).
80
70
Volume fraction (%)
60
50
40
30
void
20
c-Si
c-Si
a-Si
10
0
0
10
20
30
40
50
60
70
80
Measurement number
Fig. 1. Time evolution (each step correspond to 50 s) of the material composition acquired from the optical
modeling of measured in-situ data. The void, the microcrystalline silicon matrix, the crystalline and the
amorphous silicon fractions are marked in the figure, respectively.
73
The transmission electron microscopy has been used to verify the conclusions from optical
modelling, confirming that (i) there is a thin nucleation layer formed at the beginning of the
material deposition on the substrate, from which silicon crystals start to grow, (ii) the volume
fraction of the crystalline silicon gradually increases as the cone crystals become larger in size,
forming a rough surface after their collisions and subsequent high crystalline fraction material
growth.
Acknowledgement: The authors acknowledge the financial support from projects
SP2014/86 and IT4Inovations CZ.1.05/1.1.00/02.0070.
REFERENCES
[1]
[2]
TORRES RIOS, A., DJERIDANE, Y., NATH, M., REYDET, P. L., REYAL, J. P., ROCA I
CABARROCAS, P.: Epitaxial Growth of Crystalline Silicon on N42 Alloys by PECVD at
175°C for Low Cost and High Efficiency Solar Cells, EU PVSEC Proceedings, 2011, p. 2435.
MRÁZKOVÁ, Z., TORRES-RIOS, A., RUGGERI, R., FOLDYNA, M., POSTAVA, K.,
PIŠTORA, J., ROCA I CABARROCAS, P.: In-situ spectroscopic ellipsometry of
microcrystalline silicon deposited by PECVD on flexible Fe-Ni alloy substrate for photovoltaic
applications, under review in Thin Solid Films.
74
SUBMICRON CALCIUM PHOSPHATE PARTICLES STUDY ANCHORED
ON CLAY SUPPORTS
L. PAZOURKOVÁ1*, G.SIMHA MARTYNKOVÁ1,2, M. HUNDÁKOVÁ1,2, M. VALÁŠKOVÁ1,2
Nanotechnology Centre, VŠB – Technical university of Ostrava, 17. listopadu 15, 70833, Ostrava-Poruba,
Czech Republic; email: lenka.pazourkova@vsb.cz
2
IT4 Innovations Centre of Excellence, VŠB – Technical university of Ostrava, 17. listopadu 15, 70833,
1
KEY WORDS: calcium phosphate, clay mineral, wet precipitation
In recent years the synthetic calcium phosphates (mainly hydroxyapatite) are widely
studied due to their similarities to minerals occurring in human body [1]. The preparation
techniques include lot of methods [2], but utilization of clay minerals as support for calcium
phosphate is very spare [3, 4].
The aim of this study is to compare in-situ preparation of calcium phosphate with the main
component of hydroxyapatite (CPH) on pure clay minerals and sodium form of clay minerals,
to further usages and applications as biomaterial. The final samples were characterized using
X-ray powder diffraction (XRD) and scanning electron microscopy (SEM). The CPH particles
of different size and morphology were formed depending on the type of clay mineral and
chemistry of the montmorillonite (Mt) and vermiculite (Ver).
Fig. 1. Schematic for A) mixing and B) sonication preparation of calcium phosphate supported clay mineral.
The samples of CPH supported on clay minerals were prepared by wet precipitation, more
precisely by mixing and sonication. We used vermiculite and montmorillonite as supporting
clay minerals. Fig. 1 shows schema of preparation procedure.
X-ray diffraction patterns of CPH-NaMt composite (Fig. 2c, 3c) and sodium form of
montmorillonite (NaMt) were compared with CPH-Mt prepared in our previous study
(Fig. 2b, 3b) [5]. It was observed that main intensive reflections (d = 0.282 nm) of calcium
phosphate (with majority of hydroxyapatite) are present in both samples of pure and sodium
form of montmorillonite. In the case of vermiculite samples prepared by mixing method, the
75
samples with CPH show absence of d(001) reflection of pure vermiculite [5] and monoionic
sodium vermiculite (NaVer), respectively. The d(002) reflection of monoionic NaVer is shifted
from 1.225 nm to 1.471 nm of CPH-NaVer. This indicates that CPH influences clay structure
in the way of altering interlayer space of clays.
Fig. 2. XRD patterns of samples prepared using
mixing: a) CPH, b) CPH-Mt, c) CPH-NaMt.
Fig. 3. XRD patterns of samples prepared by
sonication: a) CPH2, b) CPH-Mt2, c) CPH-NaMt2.
From the SEM micrographs is evident that CPH is anchored on the surface of both clay
minerals in state of submicron particles. The CPH particles show different size and morphology
in dependence on type of clay mineral and preparation method.
Acknowledgement: The authors gratefully acknowledge the support by projects: Ministry
of Education, Youth and Sport of Czech Republic SP2014/82 and IT4 Innovations Centre of
Excellence project reg.no.cz.1.05/1.1.00/02.0070. Authors thank M. Heliová for SEM
micrographs.
REFERENCES
[1]
[2]
[3]
[4]
[5]
ZHANG, J., LIU, W., SCHNITZLER, V., TANCRET, F., BOULER, J-M.: Calcium phosphate
cements for bone substitution: Chemistry, handling and mechanical properties, Acta
Biomaterialia 10, 2014, pp. 1035-1049.
SADAT-SHOJAI, M., KHORASANI, M-T., DINPANAH-KHOSHDARGI, E., JAMSHIDI, A.:
Synthesis methods for nanosized hydroxyapatite with diverse structures, Acta Biomaterialia 9,
2013, pp. 7591-7621.
AMBRE, A., KATTI, K.S., KATTI, D. R.: In situ mineralized hydroxyapatite on amino acid
modified nanoclays as novel bone biomaterials, Materials Science and Engineering C 31, 2011,
pp. 1017-1029.
ROUL, J., MOHAPATRA, R., SAHOO, S.K., TRIBHUVAL, N.: Design and characterization
of novel biodegradable polymer-clay-hydroxyapatite nanocomposites for drug delivery
applications, Asian Journal of Biomedical and Pharmaceutical Sciences 2, 2012, pp. 19-23.
PAZOURKOVÁ, L., ČECH BARABASZOVÁ, K., HUNDÁKOVÁ, M.: Preparation of
hydroxyapatite/clay mineral nanocomposite, In NANOCON 2013 5th international conference
October 16th – 18th 2013, 2013 Hotel Voroněž I Brno, Czech Republic, Tanger Ltd. Ostrava
2014, pp. 83-88. ISBN: 978-80-87294-47-5. In Press.
76
PREPARATION OF SUBMICRON PARTICLES OF BIOLOGICALLY
ACTIVE SUBSTANCES USING SUPERCRITICAL FLUIDS
D. PLACHÁ1*, T. SOSNA1, E. VACULÍKOVÁ1, M. MIKESKA1, R. DVORSKÝ2
VŠB-Technical university of Ostrava, Nanotechnology Centre, 17. listopadu 15, 708 33 Ostrava Poruba,
Czech Republic; email: daniela.placha@vsb.cz
2
VŠB-Technical university of Ostrava, Faculty of Mining and Geology, 17. listopadu 15, 708 33 Ostrava
Poruba, Czech Republic
1
KEY WORDS: nanoparticles, caffeine, aspirin, supercritical fluids, Spe-ed SFE-4
Numerous conventional methods are used to reduce the particle size, for example milling,
spray drying, re-crystallization using solvent evaporation, sieving and grinding. However, these
methods are characterized by the many disadvantages such as poor control on size, unsuitable
morphology, wide particle size distribution, exposure of particles to the locally high
temperature and loss of particles in spray drying and milling techniques [1]. Supercritical fluids
have wide range of industrial applications. Using of supercritical fluids for particle size decrease
seems to be a very effective method for many reasons, especially for working in mild
temperatures and for an ability of particle size reductions to nanometric levels.
The technique of particles micronization using supercritical fluids is convenient especially
for preparations of pharmaceuticals and other biologically active substances. Most of
pharmaceuticals are poorly soluble or insoluble in aqueous body fluid systems; this fact limits
their bioavaibility. The dissolution rate can be positively influenced by increase of particle
surface area through reduction of their size. Next prerequisites are suitable and uniform
morphology and narrow particle size distribution [1, 2].
Several supercritical fluid techniques for particles size reduction are known, such as RESS
– Rapid Expansion of Supercritical Solution, PGSS – Particles from Gas-Saturated
Solution/Suspension, GAS – Gas Anti Solvent, SAS – Supercritical Anti Solvent, SEDS –
Solution Enhanced Dispersion by Supercritical Fluid [1, 2].
The RESS technique using pure supercritical CO2 is an alternative how to quickly and
naturally reduce the particle size of various materials. Principle of the technique is: A treated
compound is dissolved in a supercritical fluid; consequently the solution is suddenly
depressurized through a nozzle and expands inside a chamber with much lower pressure. The
rapid depressurization of the supercritical phase causes decreased solubility of the solute which
precipitates as a powder in a gas phase [3].
The laboratory extraction system Spe-ed SFE-4 (Applied Separations) was used for size
reducing of biologically active substances such as caffeine (Fig. 1), aspirin and cimetidine with
application of RESS technique and supercritical CO2. This device is primarily determined to be
used as an extractor of non-polar organic compounds from solids; however the producer admits
that a nanoparticle production is possible.
The working pressure was set to 150 bar and several temperatures were applied (45, 50, 80
and 100°C). The CO2 flow was maintained at 5 l.h-1. The resulted particles of treated substances
(formed caffeine particles are presented on Fig. 2) were evaluated by using SEM, FTIR, XRD
and particle size distribution methods. Volumetric particle size distribution confirmed an
influence of the working temperatures on the particle size. The SEM evaluation is necessary to
observe morphology of formed particles (Fig. 1 and 2). No significant changes were observed
in chemical structure as confirmed by FTIR and XRD.
77
Fig. 1. Caffeine particles before treatment with
supercritical CO2.
Fig. 2. Caffeine particles after treatment with
supercritical CO2.
Acknowledgement: The authors gratefully acknowledge the support by the European
Regional Development Fund in the IT4Innovations Centre of Excellence
(CZ.1.05/1.1.00/02.0070), in the ENET Centre (CZ. 1.05/2.1.00/03.0069).
REFERENCES
[1]
[2]
[3]
KESHAVARZ, A., KARIMI-SABET, J., FATTAHI, A., GOLZARY, A., RAFIEE-TEHRANI,
M., DORKOOSH, F. A.: Preparation and characterization of raloxifene nanoparticles using
Rapid Expansion of Supercritical Solution (RESS), The Journal of Supercritical Fluids, 63,
2012, pp. 169-179.
SAMEI, M., VATANARA, A., FATEMI, S., NAJAFABADI A. R.: Process variables in the
formation of nanoparticles of megestrol acetate through rapid expansion of supercritical CO2.
The Journal of Supercritical Fluids, 70, 2012, pp. 1-7.
SFC 526, The Micronization of Drug Particles by the Rapid Expansion of a Supercritical
Solution. Application Note. Applied Separations.
78
PREPARATION OF CARBON NANO FILLERS FOR METALIC
COMPOSITES
L. ROZUMOVÁ1*, G. SIMHA MARTYNKOVÁ1
1
Nanotechnology Centre, IT4Innovations Centre of Excellence, VŠB-Technical University of Ostrava,
17. listopadu 15, 708 33 Ostrava-Poruba, Czech Republic; email: lucia.rozumova@vsb.cz
KEY WORDS: carbon nanotubes, silver
This paper summarises the research work carried out in the field of carbon nanotube (CNT)
metal matrix composites (MMCs). Much research has been undertaken in utilising CNTs as
reinforcement for composite material. However, CNT-reinforced MMCs have received the
least attention. These composites are being projected for use in structural applications for their
high specific strength as well as functional materials for their exciting thermal and electrical
characteristics [1].
The present paper focuses on preparation of metal matrix nanocarbon composites
(MMNCs) that include processing technique. Metal matrix nanocarbon composite (MMNCs)
was prepared high energy ball milling. Time of milling was set on 120 hours.
Fig. 1. XRD pattern of GAg100/120- enriched graphite with Ag (GAg-original sample).
Acknowledgement: We are grateful to project CZ.1.05/1.1.00/02.0070 – IT4Innovations
Centre of Excellence for financial support of this work. This paper has been elaborated in the
framework of the Nanotechnology – the basis for international cooperation project, reg. no.
CZ.1.07/2.3.00/20.0074 supported by Operational Programme 'Education for competitiveness'
funded by Structural Funds of the European Union and state budget of the Czech Republic.
79
REFERENCES
[1]
BAKSHI, S. R., LAHIRI, D., AGARWAL, A.: Carbon nanotube reinforced metal matrix
composites – a review. International Materials Reviews 55, 2010, pp. 1-24.
80
TiO2 – BASED SORBENT FOR LEAD IONS REMOVAL
J. SEIDLEROVÁ1, M. ŠAFAŘÍKOVÁ2, L. ROZUMOVÁ1*, I. ŠAFAŘÍK2, O. MOTYKA1
Nanotechnology Centre, VŠB – Technical University of Ostrava, 17. listopadu 15, 708 33 Ostrava;
email: lucia.rozumova@vsb.cz
2
Institute of Nanobiology and Structural Biology of GCR, Na Sadkach 7, 370 05, Ceske Budejovice
1
KEY WORDS: TiO2, sorption, lead ions
The search of new technologies for removal of toxic metals from wastewaters has focused
the attention to sorption technologies which are based on metal adsorption on biological
materials (for example nut shells [1], maize leaves [2], tree ferns [3], grape stalk wastes [4]),
various composites or clay minerals. TiO2 is widely used in industry as the white pigment in
paints, as the filler and additive in cosmetics and pharmaceuticals, as photocatalyst, and also as
a sorbent [5]. The present study is focused on the sorption of Pb ions. Batch adsorption
experiments were carried out, aiming to remove lead ions from aqueous solutions using
nanopowder of TiO2.
Non-magnetic nanopowder of TiO2 and magnetically modified nanopowder of TiO2 were
employed for the sorption experiments. FeSO4 .7 H2O was used for synthesis of magnetically
responsive TiO2. The suspension of non-magnetic TiO2 and FeSO4 .7 H2O (at pH 12) was
treated by microwave irradiation for 10 min. The formed magnetically responsive composite
was captured using an appropriate magnetic separator or NdFeB magnet. Then magnetic TiO2
formed was air dried at ca 60°C [6]. A detailed study of the process was performed using various
concentrations of lead ions.
Non-magnetic and prepared magnetic material were characterized by using scanning
electron microscopy, X-ray diffraction methods and AFM. The particle size and specific surface
area were determined. The changes of Fe content in magnetically modified material after
sorption experiments were observed as well. A flame atomic absorption spectrometer was used
for determination the Pb and Fe concentration. Adsorption process has been modelled by
various sorption isotherms.
Acknowledgement: Authors thank to the financial support of Projects: GAČR No. 13
13709S/P503. This paper has been elaborated in the framework of the project New creative
teams in priorities of scientific research, reg. no. CZ.1.07/2.3.00/30.0055, supported by
Operational Programme Education for Competitiveness and co-financed by the European
Social Fund and the state budget of the Czech Republic and in the framework of the
Nanotechnology – the basis for international cooperation project, reg. no.
CZ.1.07/2.3.00/20.0074 supported by Operational Programme 'Education for competitiveness'
funded by Structural Funds of the European Union and state budget of the Czech Republic.
REFERENCES
[1]
[2]
[3]
ORHAN, Y., BUYUKGUNGOR, H.: The removal of heavy metals by using agricultural
wastes. Water Sci. Technology, 1993, vol. 28, pp. 247-255.
BABARINDE, N. A. A., BABALOLA, J. O., SANNI, R. A.: Biosorption of lead ions from
aqueous solution by maize leaf. Int. J. Phys. Science, 2006, vol. 1, pp. 23-26.
HO, Y.S., CHIUB, W. T., HSUB, C. S., HUANGA, C. T.: Sorption of lead ions from aqueous
solution using tree fern as a sorbent. Hydrometallurgy, 2004, vol. 73, pp. 55-61.
81
[4]
[5]
[6]
VILLAESCUSA, I., FIOL, N., MARTÍNEZ, M., MIRALLES, N., POCJ, J., SERAROLS, J.:
Removal of copper and nickel ions from aqueous solutions by grape stalks wastes. Water
Researcher, 2004, vol. 38, pp. 992-1002.
PEHLIVAN, E., ALTUN, T., CETIN, S., BHANGER, M. I.: Lead sorption by waste biomass of
hazelnut and almond shell. J. Hazard. Mater. 167, 2009, pp. 1203-1208.
SAFARIK, I., HORSKA, K., POSPISKOVA, K., MADEROVA, Z., SAFARIKOVA, M.:
Microwave assisted synthesis of magnetically responsive composite materials. IEEE Trans.
Magn. 49, 2013, pp. 213-218.
82
PROPERTIES OF KAOLINITE TREATED BY DIFFERENT
TEMPERATURES
M. TOKARČÍKOVÁ1*, K. MAMULOVÁ KUTLÁKOVÁ1, J. SEIDLEROVÁ1
1
Nanotechnology Centre, VŠB-Technical University of Ostrava, 17. listopadu 15/2172,
708 33 Ostrava-Poruba, Czech Republic; email: michaela.tokarcikova@vsb.cz
KEY WORDS: kaolinite, calcination, leaching test
This article deals with the influence of calcination on stability of kaolinite. Clay mineral
kaolinite Al2Si2O5(OH)4 is a suitable material for prepare photoactive composite with TiO2
nanoparticles (NPs). The composite was prepared by hydrolysis of kaolinite (SAK47) and
TiOSO4 (Precheza a.s.) as a TiO2 precursor. Prepared composite was dried at 105°C or
calcination at 600°C. Not only photoactive properties of prepared composite are important but
the stability and possible impact on the environment are important as well. Therefore,
composite stability was studied by leaching test in demineralization water and extraction agents
with different pH. However, determined concentration of leached aluminum was high,
particularly in extract obtained by leaching of calcined composite. Therefore, kaolinite
calcination was studied for deep understanding of the influence of method preparation on
photoactive composite properties and its stability. Kaolinite was calcined at different
temperatures (100°C, 200°C, 300°C, 400°C, 500°C, 600°C, 700°C and 800°C). After the
calcination of kaolinite at 400 – 650°C, the process of kaolinite dehydroxylation results in
formation of metakaolinite:
Al2Si2O5(OH)4 → Al2Si2O5(OH)xO2-x + (2-x/2) H2O
(1)
K
K
K
Q
M
M
Q
M
M
Q
M
K
K
with a low value of x (about 10% of residual hydroxyl groups in metakaolinite). Disordered
structure of metakaolinite possesses a huge reactive potential. X-ray powder diffraction patterns
of kaolinite calcined at 100°C (KA1) and 600°C (KA6) shows Fig. 1.
KA1
KA6
5
10
20
30
40
50
2-Theta - Scale
Fig. 1. XRPD patterns of kaolinite calcined at 100 °C (KA1) and 600 °C (KA6).
Legend: K - kaolinite, M - muscovite, Q - quartz.
Kaolinite treatment by different temperatures was leached in demineralization water and
extraction agent simulated acid rains. Leaching test was prepared in accordance with European
83
technical standard EN 12457-2. The influence of calcination was evaluated by determination of
aluminum, silicon and other elements in final extracts. Atomic emission spectrometry with
inductively coupled plasma (SPECTRO CIROS VISION) was used for determined of elements
concentration. Structure changes of calcined kaolinite were determined by rtg. diffraction
(Bruker D8 Advance diffractometer).
84
INFLUENCE OF VOID ON THE MECHANICAL PROPERTY OF
NANOMATERIAL
K. YODEN1*, Y. SAITO1, Q. YU1
1
Department of Mechanical Engineering, Graduate School of Engneering, Yokohama National University
Tokiwadai 79-5, Hodogaya-ku, Yokohama, Japan; email: kotaro-yoden-wp@ynu.jp
KEY WORDS: Ag-nano, SiC-chip, void, nanomaterial, high-temperature power device
In recent years, many engineers develop electric vehicles. The power device are required
high performance, miniaturization and weight saving. Previous power device use Si-chip, so
the upper limit of the operating temperature is 150C. It enlarge cooler and prevent power
device to be smaller. Now SiC-chip are developed as an alternative to Si chip. SiC-chip can
work 250C or more. It is expected that SiC chip enable power device to become smaller.
However previous soldering melts high operating temperature. It is expected that bonding
technology using nanoparticles solve this problem. Sintering mechanism of nanoparticle is
complex and unexplained. It is known that sintered material of nanoparticle possesses many
micro voids. We have two tasks to use this nano-particle effectively. One is to clarify relation
void and sintering conditions. And two is to clarify the effect void give nanoparticle material.
In this study, I study influence that voids give the mechanical property of nanomaterial.
Fig. 1. Section having the void of a uniform shape.
Fig. 2. Section having the void of a heterogeneous shape.
Fig. 1 and Fig. 2 are SEM (Scanning Electron Microscope) images of a cross-section of
Ag-nano sintered material which are sintered under same condition. The images show that void
shape, size, place and rate are various. To clear relation these elements and mechanical property,
I simulated some model using FEM analyse.
Fig. 3 is model.1 and model.2 which
are given by Fig. 1 and Fig. 2. Model.1
have voids which are uniform shape, and
model.2
have
voids
which
are
heterogeneous shape. These models are
unified by 9.5% void rate. Size of these
models are one side of the square 80 m. I
cut these models in mesh one side is 1 m,
and vertical and horizontal strained force
of 1% in models.1 and 2. We investigated
the effects of void.
Fig. 3. model.1 (left) and model.2 (right).
85
25
25
20
20
Stress[MPa]
Stress[MPa]
15
Vartical
10
Horizontal
5
0
15
Vertical
10
Horizintal
5
0
0
0,2
0,4
0,6
0,8
1
0
Strain[％]
Fig. 4. S-S diagram (model.1).
0,2
0,4
0,6
0,8
Strain[％]
1
Fig. 5. S-S diagram (model.2).
Yong's modulus [GPa]
Fig. 4 shows that when the strain is 1%,
68,0
68,0
vertical stress is higher than horizontal stress by
70,0
3.9% in model.1. Fig. 5 shows that when the
60,0
strain is 1%, vertical stress is higher than
50,0
model.1
36,4
36,0 34,9
horizontal stress by 18.5% in model.2. From
40,0
30,0
model.2
these results, I understood the heterogeneous
30,0
Al(bulk)
shape of void bring about anisotropy. Fig. 6 is
20,0
Young’s modulus which are calculated by
10,0
degree of leaning of the s-s diagram. Fig. 6
0,0
show that Young’s modulus in model.1 are
Horizontal
Vertical
hardly a difference in vertical direction and a
Fig. 6. Young’s modulus.
horizontal direction. And the difference in the
case of bulk is approximately 46%. In model.2,
horizontal Young’s modulus is higher than vertical Young’s modulus by 14.3%. It means
model.2 have anisotropy. And the difference in the case of bulk is approximately 56%.
Thus, the approximately 10% of void lower 45-55% of Young's moduluses. And the
influence of void is depend on shape of void. And heterogeneous shape of voids make
anisotropy. So when metal nanoparticles materials are used, it is necessary to control a not only
rate of void but also shape of void by sintering condition.
REFERENCES
[1]
YAMAGIWA, M., YU, Q., FUJITA, M., SHINOHARA, M., MURAKAMI, Y.:
”ReliabilityStudy of Mouting Structure for HighTemperature Power Semiconductor Device
Chip Using High Purity Aluminium”, Proceedings of Intersociety Conference on Thermal and
Thermomechanical Phenomena in Electronic Systems (ITherm2008), Orlando Florida, 2008.
86
DETERMINATION OF ANISOTROPIC CRYSTAL OPTICAL PROPERTIES
USING MUELLER MATRIX SPECTROSCOPIC ELLIPSOMETRY
K. POSTAVA1,2, R. SÝKORA2*, D. LEGUT2, J. PIŠTORA2
Department of Physics, Technical University of Ostrava, 17. listopadu 15, 708 33 Ostrava – Poruba, Czech
Republic
2
Nanotechnology centre, Technical University of Ostrava, 17. listopadu 15, 708 33 Ostrava – Poruba, Czech
Republic; email: rudolf.sykora@vsb.cz
1
KEY WORDS: spectroscopic ellipsometry, anisotropic crystal, Mueller matrix, permittivity tensor
Recent development in spectroscopic ellipsometry and ellipsometric instrumentation
triggers wide applications of the technique to characterize anisotropic nanostructures, periodic
systems, and also crystals with reduced symmetry. Moreover, the Mueller matrix ellipsometry,
i.e. measurement of all 15 reduced Mueller matrix elements, enables a complete
characterization of reflection properties of the samples, including phenomena as mode
conversion and depolarization [1]. The main task is usually to determine spectra of all
components of the permittivity tensor. The number of independent components depends on
crystal symmetry.
In this paper the Mueller matrix ellipsometry in the spectral range from 0.73 to 6.4 eV
measured using dual rotating compensator ellipsometer RC2 (Woollam company) is applied to
study anisotropic crystals with tetragonal and monoclinic symmetry.
As a typical uniaxial sample we have characterized a rutile (TiO2) tetragonal crystal. The
optical axis of the sample is parallel to its surface. The sample is characterized at variable angle
of incidence and variable azimuthal rotation angle. Figure 1 shows typical Mueller matrix
spectra compared with data fit obtained using the matrix model based on light propagation in
anisotropic stratified media. The angle of incidence of 45° and the azimuthal angle of
approximately 45° were chosen.
The non-zero off-diagonal blocks of the matrix show mode conversion due to the optical
anisotropy. We discuss the sensitivity to determine ordinary and extraordinary optical functions
for various directions of the optical axis.
We also discuss application of Mueller matrix ellipsometry to determine optical functions
of crystals with monoclinic symmetry. The crystal is characterized by four independent
permittivity tensor elements (three diagonal and one off-diagonal) [2]. The crystals of
Cu(H2O)(C2H8N2)SO4 attract strong interest due to their magnetic properties. The crystal
exhibits monoclinic symmetry with the axis inclination of 105.5 degree. Obtained spectra from
Mueller matrix ellipsometry are compared with ab-initio models based on density function
theory (DFT).
87
Mueller matrix components
–0.3
0.08
0.1
–0.4
0
–0.5
–0.1
–0.3
0.98
–0.5
0.1
–0.1
0.08
0
–0.04
2
4
6
0.3
0.2
0.1
0
–0.9
0
–0.1
–0.2
–0.3
0.1
0
–0.1
–0.2
0.04
0.1
0
–0.1
–0.2
–0.8
0.02
0
–0.02
0
0
0.02
0
–0.02
–0.04
1
–0.4
0.04
2
4
6
–0.8
–0.85
–0.9
2
4
6
2
4
6
Photon energy(eV)
Fig. 1. Spetra of Mueller matrix elements of the rutile crystal. The measured data (blue dots) are compared with
the model (red lines).
Acknowledgement: Partial support from the projects CZ.1.05/1.1.00/02.0070
(IT4Innovations), CZ.1.05/2.1.00/01.0040 (RMTVC), IRP 167/2014, and Czech Science
Foundation 13-30397S is acknowledged.
REFERENCES
[1]
[2]
[3]
GARCIA-CAUREL, E., OSSIKOVSKI, R., FOLDYNA, M., PIERANGELO, A.,
DRÉVILLON, B., DE MARTINO IN A.: LOSURDO, M., HINGERL, K. (EDS.): Ellipsometry
at the Nanoscale, Springer-Verlag Berlin Heidelberg 2013.
JELLISON JR., G. E., MCGUIRE, M. A., BOATNER, L. A., BUDAI, J. D., SPECHT, E. D.,
SINGH, D. J., Phys. Rev. B 84 (2011) 195439.
SYKORA, R., LEGUT, D.: J. Appl. Phys. 115, (2014) 17B305.
88
OBSERVATION OF MAGNETIC FIELDS AROUND PLASTIC
DEFORMATION AREA IN LOW CARBON ALLOY STEEL
K. KIDA1*, M. ISHIDA1, K. MIZOBE1
1
University of Toyama, 3190, Gofuku, Toyama City, Toyama Prefecture, 930-8555, Japan;
email: kida@eng.u-toyama.ac.jp
KEY WORDS: magnetic flux density, plastic deformation, scanning hall probe microscopy
Failure of machine elements is mainly caused by fatigue crack growth occurring as results
of cyclic stress concentration. Following the pioneering studies by Orowan and Irwin many
investigations have been carried out in order to understand the fatigue crack growth in steels.
One of the important topics in this field is plastic deformation occurring around the crack tip.
Non-destructive methods that can be related to the plastic deformation around small crack tip
area are necessary to study the crack growth. For our previous works, we developed a scanning
Hall probe microscope (SHPM) equipped with GaAs films and observed fatigue cracks growing
from artificial slits in steels (JIS, SUJ2 [1,2], S45C [3]). Furthermore, we applied this SHPM
technique to contact problem of tool steel (SKS93) [4], fatigue of welding part (SS400) [5],
tensile loading of thin plate (SKS93) [6] and plastic deformation around a Vickers indentation
(SKS93) [7]. A one-dimensional sensor was used in the first research. From the second research,
three-dimensional observations were carried out using three 10 µm-sized Hall films in order to
study the features of magnetic fields under various loadings.
Fig. 1. Vertical component of three-dimensional magnetic flux density (Bz) in a specimen. Observation covers
the area of (x, y) = (6.0 mm, 12.0 mm). The observations of magnetic flux density were done after
magnetization, before the tests (a), and after the second indentation test (b).
Plastic deformation area expands as the crack grows. In the present research we induced
two Vickers indentations along the center line on a low carbon alloy steel plate (JIS, S45C) and
observed the relation between magnetic field and plastic deformation. Demagnetization was
done by using a coil in order to normalize the initial magnetic field of the specimen. After the
demagnetization, a permanent magnet block was slid along the center line area of the specimen.
The size and residual magnetic flux density of the block were 1 – 10 – 10 mm3 and 99 mT,
respectively. Three components of the magnetic field, Bx, By and Bz were compared before and
after the indentation tests.
89
Magnetic component, Bz [mT]
Fig. 1 is an example of the magnetic fields
Initial distribution
1.0
before and after the tests. After the first and second
Vickers indentations were induced on the
specimen surface, they were observed. The load
0.9
and diagonal length of the indentations were
20 kgf and 300 μm. Fig. 1(a) shows a magnetic
field component (Bz) in a direction vertical to the
1st indentation
0.8
specimen surface before the test, and (b) is after the
second indentation test. The distance between
2nd indentation
5.7
mm
6.6
mm
0.7
these indentations was 900 µm. The positions of
0
4
8
12
the first and second indentations were y = 5.7 mm
y [mm]
and 6.6 mm, respectively. We can see clearly the Fig. 2. Change in magnetic field component, Bz,
due to Vickers indentations.
changes in magnetic fields in a circle of Fig. 1(b).
Observation area was divided into 10 µm-width segments whose longitude axes are parallel to
the x-axis, and all of the magnetic components in the segments were compared to their initial
values observed before testing. The segment coordinates are numbered along the y-axis. The
peak to bottom value in each segment was calculated with the maximum and minimum values
in it. After the calculation, distributions of the peak to bottom values of Bx, By and Bz in all
segments were arranged along the y-axis. Fig. 2 shows the changes in the peak to bottom values
of Bz before and after the indentation tests. When comparing the magnetic components after
the first and second indentations, it is found that the bottom peak area expands toward the
second indentation area along the center line (y-axis). The left part of the curve measured after
the first indentation test corresponds to that after the second test. This means the magnetic fields
record the change in plastic deformation.
Acknowledgement: A part of this research was supported by Grants-in-Aid for Scientific
Research, JSPS (KAKENHI, Scientific Research (c), No. 23560089).
REFERENCES
[1]
[2]
[3]
[4]
[5]
[6]
[7]
KIDA, K., TANABE, H., OKANO, K.: Changes in magnetic flux density around fatigue crack
tips, Fatigue & Fracture of Engineering Materials & Structures, 32, 3, 2009, pp. 180-188.
KIDA, K., SANTOS, E. C., HONDA, T., KOIKE, H., ROZWADOWSKA, J.: Observation of
magnetic flux density around fatigue crack tips in bearing steel using a SHPM with a threedimensional small-gap probe, Int. J. Fatigue, 39, 2012, pp. 38-43.
KIDA, K., SANTOS, E. C., URYU, M., HONDA, T., ROZWADOWSKA, J.,
SARUWATARI, K.: Changes in magnetic field intensities around fatigue crack tips of medium
carbon low alloy steel (S45C, JIS), Int. J. Fatigue, 56, 2013, pp. 33-41.
KIDA, K., URYU, M., HONDA, T., SANTOS, E. C., SARUWATARI, K.: Three-Dimensional
Observation of Magnetic Fields in Alloy Tool Steel under Spherical Hertzian Contact, Materials
Research Innovations, 18, Supplement1, 2014, pp. 71-75.
KIDA, K., HONDA, T., SANTOS, E. C., SARUWATARI, K., URYU, M., HOURI, K.,
TANABE, T., KANEMASU, K.: Three-dimensional Magnetic Microscopy of Early Stage
Fatigue in WMZ of Low Carbon Steel Plates (JIS-SS400), Materials Research Innovations, 18,
Supplement1, 2014, pp. 66-70.
KIDA, K., URYU, M., HONDA, T., SANTOS, E. C.: Changes in Three-dimensional Magnetic
Fields of Star shaped JIS-SKS93 Plates Embedded in Clear Acrylic Cold Mounting Resin under
Tensile Loads, Materials Research Innovations, 18, Supplement1, 2014, pp. 89-93.
HONDA, T., SANTOS, E. C., KIDA, K.: Scanning Hall probe microscopy of residual magnetic
fields around plastic deformation of Vickers indentations in carbon tool steel (JIS, SKS93),
Mechanics of Material, 69, 2014, pp. 262–269.
90
MAGNETO-PLASMONIC PROPERTIES OF Au/Fe/Au
PLANAR NANOSTRUCTURES: THEORY AND EXPERIMENTS
J. VLČEK1,3*, M. LESŇÁK2, P. OTIPKA3
National Supercomputing Centre IT4Innovations, VŠB-TU Ostrava, Czech Republic;
email: jaroslav.vlcek@vsb.cz
2
Regional Materials Science and Technology Centre, VŠB-TU Ostrava, Czech Republic
3
Dept. of Mathematics and Descriptive Geometry, VŠB-TU Ostrava, Czech Republic
1
KEY WORDS: magneto-plasmonics; response factors; sensitivity criteria
The non-reciprocity of magnetooptical
reflection response by surface plasmon
excitation in the planar Au/Fe/Au/glass
nano-systems with prism coupling is
studied. These structures are intended as
magnetic field sensor units combining
magneto-optical and surface-plasmonresonance effects. In order to simulate the
diffraction response of discussed structures
to external magnetic field theoretical model
based on matrix algorithm is applied. The
ability of MO-SPR systems to sensing of
magnetic field is analysed using the
response factor
0.3
Au 7 nm, Fe 8 nm, Au 14 nm
0.25
0.2
0.15
K = tan 
0.1
r
F
0.05

0
-0.05
-0.1
-0.15
-0.2
45
45.5
46
46.5
incidence angle [deg]
47
Fig. 1. Sensitivity criteria F and K.
r() 


R pp
 Rpp


R pp
 Rpp
,
(1)
where Rpp denotes the reflectance of p-polarized beam; and, the sign in upper index relates to
the orientation of external magnetic field. Unlike our previous work [1] the newly proposed
sensitivity criteria F and K are applied (see Fig. 1).
Obtained theoretical results are compared with experiments realized using the measuring
device Multiskop (Optrel GbR, Germany). In particular, the intensity of reflected light is
detected in dependence to the angle of incidence at the wavelength 632.8 nm. We completed
this equipment by digitally controlled electro-magnet, which enables production of a predefined
magnetic field in transversal configuration.
REFERENCES
[1]
VLČEK, J., LESŇÁK, M., PIŠTORA, J., ŽIVOTSKÝ, O.: Magneto-optical sensing of
magnetic field, Optics Communications 286, 2013, pp. 372-377.
91
92
IMAGE ANALYSYS VIA REACTION DIFFUSION SYSTEM FOR EDGE
DETECTION
K. NAKANE1*, H. MAHARA2, K. KIDA3
1
Osaka University, Faculty of medicine, 1-7, Yamadaoka, Suita, Osaka, Japan;
email: knakane2008l@gmail.com
2
Chiba University Hospital, 1-8-1, Inohana, Chuo-ku, Chiba-shi, 260-8677, Japan
3
University of Toyama, Gofuku 3190, Toyama1-8-1, Inohana, Chuo-ku, Chiba-shi, 260-8677, Japan
KEY WORDS: reaction diffusion, edge detection, homology
To measure the particle size of the material and to analyze the images of the structure, we
need to detect the edge of the particle. Since the grain boundaries are not so clear, it takes a lot
of time to detect them, manually.
Here, we will present a support method to detect the edge from material images. By solving
non-linear partial differential equations numerically, we make blurry images clear. Combining
ordinary image analysis method, we will take the edge of grain. In this talk, we introduce the
results of our method, and discuss the possibility of this method.
(a)
(b)
(c)
Fig. 1. (a) The image of SUJ2 (Q3T1). (b) The numerical result of reaction diffusion equation.
(c) The superimposed images of (a) and (b).
This method is developed for edge detection and figure-ground separation [1]. This method
can realize a high quality processing on noisy image compared with the Median filter [2].
Nomura et al. extended this algorithm in order to detect edge from a gray-scale image [3]. This
algorithm was applied, recently, for detection of blood vessels in fingertips [4].
The reaction-diffusion equations consist of three variables with diffusion terms. Two of
them are the FitzHugh-Nagumo equations that are a model of nerve membrane [5, 6]. Another
93
one is a variable for averaging threshold that is introduced by Nomura et al.. Precise equations
are described in the reference therein. Fig. 1 shows the result of the edge detection with this
method.
(d)
(e)
Fig. 2. (d) Noise reduction result of (c). (e) The image that (d) is superimposed on (b).
By adding a manual correction (noise reduction and so on), we can detect the edge easily.
For this purpose, we need useful GUI system. To make a GUI that suits the purpose of the
engineer, it would be necessary to devise in each scene. We are very glad, if we could discuss
the direction of development.
REFERENCES
[1]
[2]
[3]
[4]
[5]
[6]
NOMURA, A., ICHIKAWA M., MIIKE H., EBIHARA M., MAHARA H., SAKURAI T.:
Ralizing Visual Functions with Reaction-Diffusion Mechanism: Journal of the Physical Society
of Japan 72, 2003, pp. 2385-2395.
EBIHARA M., MAHARA H., SAKURAI T., NOMURA A., MIIKE H., Image processing by a
discrete Reaction-Diffusion System: Proceeding of Visualization, Imaging and Image Processing
396, 2003, pp. 145-150.
NOMURA, A., ICHIKAWA M., SIANIPAR R. H., MIIKE H.: Edge Detection with ReactionDiffusion Equations Having a Local Average Threshold: Pattern Recognision and Image
Analysis 10, 2008, pp. 289-299.
NAKANE, K. AND MAHARA, H: A numrical method to detect the edge of vessels, in
preparation.
FITZHUGH R.: Impulses and Physiological States in Theoretical Models of Nerve Menbrene:
Biophysical Journal 1, 1961, pp. 445-466.
NAGUMO J., ARIMOTO S., YOSHIZAWA S.: An Active Pulse Transmission Line
Simulationg Nerve Axon: Proceeding of the IRE 50, 1962, pp. 2061-2070.
94
EFFECTS OF INCLUSION ON THE IN-PLANE MECHANICAL
PERFORMANCE OF MICRO-LATTICE PLATE
K. USHIJIMA1*, W. J. CANTWELL2, D. H. CHEN3
1
Tokyo University of Science, 6-3-1 Niijyuku, Katsushika-ku, Tokyo, Japan; email: kuniharu@rs.tus.ac.jp
Khalifa University of Science, PO Box 12778, Abu Dhabi, UAE
3
Jiangsu University, 301 Xuefu Road, Zhenjiang, Jiangsu, P.R. China
2
KEY WORDS: micro-lattice structure, initial stiffness, yield strength, finite element method
Cellular structures, such as honeycombs, lattices and foams have been taken attention and
used for many structural applications owing to their superior mechanical performance per unit
mass. For decades, many researchers have been investigated the mechanical properties of
cellular structures with regular and irregular cells by using numerical, theoretical and
experimental approaches.
One of our authors have developed
the selective laser melting (SLM)
technique for manufacturing microlattice structures at length scales of
microns. The micro-lattice structure can
be produced by using CAD data, so the
micro-architecture of the structure can be
changed easily to enhance the overall
(b) with non-uniform cells
mechanical properties such as initial (a) with uniform cells
Fig. 1. Photos of micro-lattice structure by SLM technique.
stiffness, plastic collapse or buckling
strength and impact energy absorption capacity. Figure 1 shows the example of micro-lattice
block with two types of inner cells.
The effects of hole and inclusions on the mechanical performance of honeycomb structures
have been investigated by other researchers. However, the micro-lattice structure investigated
has much potential for improving the mechanical properties by changing the strand’s length,
angles between adjacent strands.
In this study, the in-plane mechanical properties of the micto-lattice plate with inclusion is
analysed by using numerical analysis, Finite Element Method. The inclusion is centred in the
plate, and subjected to in-plane tension load. In our discussion, the effects of inclusions on the
initial stiffness and plastic collapse strength are discussed. The inclusions investigated here are
modelled by holes, softer (coarser) cells and harder (finer) cells.
Figure 2 shows analytical models investigated in this study. Also, Figures 3 shows
variations of initial stiffness E* and plastic collapse strength σ*pl with the defect width w. It can
be understood that the initial stiffness depends strongly on the shape of holes, and decreases
nonlinearly as the width w increases. On the contrary, the plastic collapse strength for each
plate is almost coincident under the same width w, and decreases linearly as w increases. That
is because the plastic collapse strength is mainly governed by the weakest point of the plate,
and the weakest point can be observed at the edge of the hole.
95
(a) Defect-w
(b) Defect-s
(c) Defect-d
(d) Nu-1_2
Fig. 2. Examples of lattice plates with holes or inclusion.
(e) Nu-2_1
(a) for initial stiffness
(b) for plastic collapse strength
Fig. 3. Variations of defect width w on material properties.
REFERENCES
[1]
[2]
[3]
[4]
[5]
GIBSON, L. J., ASHBY, M. F.: Cellular Solids: Structure and Properties, 2nd ed. Cambridge:
Cambridge Press 1997.
SILVA, M. J., GIBSON, L. J.: The Effects of Non-periodic Microstructure and Defects on the
Compressive Strength of Two-dimensional Cellular Solids, International Journal of Mechanical
Science 39, 1997, pp. 549-563.
GUO, X. E., GIBSON, L. J.: Behavior of Intact and Damaged Honeycombs: a Finite Element
Study, International Journal of Mechanical Sciences 41, 1999, pp. 85-105.
CHEN, C., LU, T. J., FLECK, N. A.: Effect of Imperfections on the Yielding of TwoDimensional Foams, Journal of the Mechanics and Physics of Solids 47, 1999, pp. 2235-2272.
CHEN, C., LU, T. J., FLECK, N. A.: Effects of Inclusions and Holes on the Stiffness and
Strength of Honeycombs, International Journal of the Mechanical Sciences 43, 2001,
pp. 487-504.
96
ASSESSMENT OF STRUCTURES LOADED AT CREEP
S. VEJVODA1*, P. POPPELKA2, P. RYŠAVÝ1
1
2
VÍTKOVICE ÚAM a.s., Mezírka 775/1, 602 00 Brno, Czech Republic; email: stanislav.vejvoda@vitkovice.cz
SES Tlmače, a.s., Továrenská 210, 935 28 Tlmače, Slovak Republic
KEY WORDS: oxide layer (scale), tube, strain, stress, crack, exfoliation, delamination, fracture
toughness, stress intensity factor, creep strain, membrane stress, stress concentration, metal
The operational parameters of classical power plants are continually increased. Many times
there was not respected that used material has the limit of usability for reliable service at lower
operational temperatures than an increased operation temperature only upon 30°C. Parameters
effecting the damage from the oxide layer Fe3O4 (magnetite) on the tube head transfer surface
were analyzed for head resisting steel Cr-Mo-V, specification 15 128.5.
A study was made in accordance with [1] and [2]. The oxide layer continually arose on the
inside tube surface 38x6.1 mm made from steel 15 128.5. For the oxide layer Fe3O4 the
following were used: i = 4.5 [J.m-2]; EOX = 208000 [MPa]; KIc = 1.4 [MPa.m0.5] and Poisson
number  = 0.262 [3].
The study was carried out for following parameters:
thickness of the oxide layer d = 0.01 mm ÷ 1.0 mm;
half-axis of the defect in the oxide layer c = 0.01 mm ÷ 1.0 mm;
asperity of the tube inside surface r = 1.6 m; 3.2 m and 6.3 m;
radius of the delaminated area of the oxide layer created on the tube metal
R = 0.025 mm ÷ 20.0 mm;
- wave length of the rough interface on the boundary oxide/metal  = 0.025 mm ÷ 20.0 mm;
- coefficient f related with geometry, for real defect f =1.
-
Some results of the analysis:
- oxide layer (scale) with the defect of c =10 μm would crack at the limit strain ct = 0.12 %;
- interfacial defect of c = 10 m on the boundary oxide layer/metal would be begin growth
under compressive stresses when the critical strain reaches the value of ci = - 0.1%;
- buckling of the oxide layer under compressive stresses would arrive at the critical strain
cb -1.0 [%], when its thickness is d  1 mm and the radius of the delaminated area is
R  10 mm;
- circumferential shearing under
compressive stresses on the
bulging oxide layer d  1 mm of
the
delaminated
area
R = 20 mm located on the
boundary oxide layer/metal
would not occur at strain
cbf  -0.01 [-];
- critical value of strain cs Fig. 1. Exfoliation of the delaminated oxide layer Fe3O4 on the
inside surface of the tube
necessary for exfoliation of the
oxide layer does not fall under 1.90 [%] when the thickness of the oxide layer is d 1.0 mm, the length wave is  20 mm,
and roughness of r = 1.6 m.
97
Exfoliation of the delaminated oxide layer Fe3O4 on the inside surface of the tube was
analyzed by program ANSYS for nominal operation conditions of the boiler (p = 22.1 MPa,
T = 625°C, t = 30 years). The calculated value of the stress intensity factor KI = 4.07 MPa.m0.5
is greater than the fracture toughness of the oxide layer KIc= 1.4 MPa.m0.5. The calculated strains
of the oxide layer were greater than their critical values for damage of the oxide layer as well.
When creep strain limits for given steels are not know, it is possible to use limits given in
[4]. It means 1% for membrane stresses, 2% for bending stresses and 5% for the stress
concentration area. Creep strains are usually calculated in areas of membrane stresses. The FEM
programs enable calculation of the whole structure, in which geometrical notches are present.
Compressive stresses and strains were calculated at these geometrical notches. The big
“compressive creep deformation” was calculated by the FEM in the small area of these notches
and until after stress relaxation the character of creep deformation was changed to tension. This
problem was discussed with creep specialists in the Czech Republic.
Fig. 2. Detail of analyzed structure;
points A, B, H – stress concentration.
Fig. 3. Point A, 2nd principle strain;
compressive stress, relaxation and growth of c.
Fig. 4. Creep strain intensity c at area
of membrane stresses.
Fig. 5. Creep strain intensity c at the stress concentration
area; boundary c =5 % is in red.
Acknowledgment: The authors gratefully acknowledge the support by Technological
agency of the Czech Republic for project No. TA02011179.
REFERENCES
[1]
[2]
[3]
[4]
SCHÜTZE, M.: Modelling oxide scale fracture. Materials at High Temperatures, Volume 22,
Numbers 1-2, February 2005, pp. 147-154.
SCHÜTZE, M, TORTORELLI, P.F., WRIGHT, I.G.: Development of a Comprehensive Oxide
Scale Failure Diagram. Oxid Met (2010) 73: pp. 389-418
Gercek: Int. J. Rock Mech. Min. Sci. 44 (2007)
Cases of ASME Boiler and Pressure Vessel Code, Case N-47-29, Class 1. Components in
Elevated Temperature Service. Section III, Div. 1, 1990.
98
DIFFUSION OF HYDROGEN IN THE TRIP 800 STEEL
J. SOJKA1*, P. VÁŇOVÁ1, V. VODÁREK1, M. SOZANSKA2
Faculty of Metallurgy and Materials Engineering, VŠB – Technical University of Ostrava, Czech Republic;
email: jaroslav.sojka@vsb.cz
2
Faculty of Materials Engineering and Metallurgy, Silesian University of Technology, Katowice, Poland
1
KEY WORDS: hydrogen electrochemical permeation, TRIP 800 steel, hydrogen embrittlement
The presented paper is devoted to the study of hydrogen diffusion in the C-Mn-Si TRIP 800
steel containing 0.2 wt. % of C, 1.5% of Mn and 1.5% of Si. The steel was tested in three
different states: in as-received state after hot and cold rolling and subsequent heat treatment;
and furthermore after 5% and 10% tensile deformation. The tensile deformation resulted in an
increase of mechanical properties (Re, Rm) and in a decrease of retained austenite content from
11.0% in as-received state to 2.2% after 10% tensile deformation. Hydrogen diffusion
characteristics were studied by means of electrochemical permeation method.
Electrochemical hydrogen permeation
9E-07
tests were carried out using a Devanathan8E-07
Stachurski two-component cell separated
7E-07
first build up
transient
by a steel membrane – working electrode.
6E-07
The exit side of the working electrode was
5E-07
second build up
transient
palladium coated to prevent from
4E-07
3E-07
hydrogen atom recombination during
decay transient
2E-07
permeation
experiments.
Hydrogen
1E-07
charging cell was filled with 0.05M
0E+00
H2SO4, while the exit cell was filled with
without
5% tensile
10% tensile
0.1 M NaOH solution. The exit cell was
deformation deformation
deformation
de-aerated by argon bubbling before and
Fig. 1. Hydrogen diffusion coefficients Deff
for all studied states.
during experiments. Firstly, the entry side
of the specimen was polarized anodically
at a current density of + 35 mA.cm-2. At the end of this period (5 minutes), H2SO4 charging
solution was renewed continuously to eliminate metallic ions from the solution. After that, two
build-up transients (BUT) were recorded, the first one at the charging current density of
-20 mA.cm-2, the second one at the charging current density of -35 mA.cm-2. Before ending the
experiment hydrogen charging was stopped and a decay transient (DT) was also recorded.
Hydrogen diffusion coefficients were calculated using the time-lag method according to
Eq. 1:
D
L2
,
6t L
(1)
where L represents the membrane thickness and tL corresponds to the time where the permeation
current reaches 63% of its steady-state value. Sub-surface hydrogen concentration was
calculated using Eq. 2:
i L
C H0   ,
DF
where i is a steady-state current density and F is Faraday’s constant.
99
(2)
The results of hydrogen diffusion coefficients are presented in Fig. 1. It is obvious that the
lowest values of hydrogen diffusion coefficient were always obtained for the first BUT. It can
be related to hydrogen trapping in both reversible and irreversible traps. The hydrogen diffusion
coefficient corresponding to the first BUT was higher for the state after 10% tensile
deformation. This behaviour confirms the hypothesis that formation of martensite facilitates
hydrogen diffusion in the TRIP steel. Hydrogen diffusion coefficients corresponding to the
second BUT were markedly higher and confirmed that most of traps were filled by hydrogen
during the first BUT. Nevertheless, values of hydrogen diffusion coefficients still remained
lower in comparison with conventional steels having bcc lattice. For decay transients hydrogen
diffusion coefficients were usually situated between values obtained for the 1st and 2nd BUT. It
confirms that during the first BUT hydrogen trapping can be expected and during DT hydrogen
detrapping can be expected, influencing thus values of hydrogen diffusion coefficient.
Measured data were also fitted with the theoretical curves of normalized hydrogen flux Jt/J.
Measured data fitted very well with the theoretical curves for the second BUT in all studied
states. However, for the first BUT and for the DT, the measured data were shifted to longer
time in comparison with the theoretical curves confirming thus the important role of hydrogen
trapping and detrapping.
Hydrogen sub-surface concentration was calculated for the first BUT using Eq. 2. The
obtained results showed that the sub-surface concentration of hydrogen was rather high mainly
if the amount of retained austenite in the structure was not too low. In the as-received state,
hydrogen sub-surface concentration reached 12.6 ppm, while after 10% tensile deformation,
hydrogen concentration was 5.4 ppm. In this way the role of retained hydrogen as an important
and very probably irreversible hydrogen trap was confirmed. The high sub-surface
concentration of hydrogen in the studied TRIP steel can, at least partially, explain its rather high
susceptibility to hydrogen embrittlement [1].
The obtained results can be summarised as follows:
- The values of hydrogen diffusion coefficients in the TRIP 800 steel were rather low and
lay between 1.10-7 cm2.s-1 and 7.8.10-7 cm2.s-1;
- The highest values of hydrogen diffusion coefficient were observed during the 2nd build
up transient where the role of hydrogen trapping was limited;
- A decrease of retained austenite content resulted in an increase of hydrogen diffusion
coefficient;
- The comparison of experimental data with the theoretical model showed a good fitting for
the 2nd build up transient, while during the 1st build up transient and during the decay
transient a shift of experimental data to longer time was observed.
- Rather high sub-surface concentration of hydrogen was determined in the studied steel
especially for states with higher retained austenite content.
Acknowledgement: The authors are grateful to the Ministry of Education of the Czech
Republic for the financial support of the project No. LE13011 “Creation of a PROGRES 3
Consortium Office to Support Cross-Border Co-operation” and the project No. LO1203
"Regional Materials Science and Technology Centre - Feasibility Program".
REFERENCES
[1]
SOJKA, J. at al.: Effect of hydrogen on the properties and fracture characteristics of TRIP 800
steels, Corrosion Science 53, 2011, pp. 2575-2581.
100
PRECIPITATION REACTIONS IN A COPPER - BEARING GOES
V. VODÁREK1*, A. VOLODARSKAJA1, Š. MIKLUŠOVÁ2, J. HOLEŠINSKÝ1, O. ŽÁČEK2
VŠB – TU Ostrava, Faculty of Metallurgy and Materials Engineering, Ostrava, Czech Republic;
email: vlastimil.vodarek@vsb.cz
2
ArcelorMittal Frýdek-Místek, Czech Republic
1
KEY WORDS: GOES, precipitation processes, sulphides, nitrides, TEM
Magnetic properties of grain oriented electrical steels (GOES) depend strongly on the
sharpness of the Goss texture. It is believed that the perfection of the final texture is significantly
affected by structural inheritance during a complex processing route of GOES 1. Factors
which are considered to be very important for the formation of the Goss texture during high
temperature annealing include the size of the initial grains with the Goss orientation, their
orientation with respect to the other grains and the role of minor phases in grain boundary
pinning 2, 3. The processing technology of Cu – bearing GOES comprises of following
production steps: slab reheating – hot rolling and coiling – 1st cold rolling – decarburization
annealing – 2nd cold rolling – high temperature annealing – thermal flattening. The effect of
copper in GOES is not well understood. Copper is believed to play several roles 2:
1. Increase the volume fraction and stability of austenite during hot rolling in the two phase
( + ) region.
2. Small copper rich sulphides could inhibit grain growth during recrystallization processes
(decarburization annealing and high temperature annealing).
3. Precipitation of  - Cu could positively affect distribution of AlN particles, which are
expected to be the most important inhibition phase 2. Ideal conditions for the
precipitation of  - Cu represents a slow heating rate during the initial stages of the high
temperature annealing (less than 30°C/h) combined with the presence of many lattice
defects after the 2nd cold rolling. At temperatures of normal grain growth  - Cu
precipitates are expected to dissolve.
4. Segregation of copper atoms at grain boundaries can modify their mobility.
This paper deals with minor phase evolution in a Cu-bearing GOES during the following
production steps of the AlN + Cu processing technology: hot rolling of slabs, 1st cold rolling +
decarburization annealing and a slow heating to the temperature of primary recrystalization
(620°C). Minor phases were investigated in TEM using carbon extraction replicas.
Chemical composition of the strip after hot rolling is shown in Table 1. Decarburization
annealing after the 1st cold rolling reduced the carbon content in the steel to 0.0029 wt.%.
Table 1 Chemical composition of the steel investigated (after hot rolling), wt.%.
C
0.03
Mn
0.25
Si
3.16
S
0.004
Cr
0.024
Cu
0.50
Altot.
0.014
Ti
0.004
N
0.009
In the state after hot rolling, ferrite grain boundaries were decorated by thin films of iron
carbides. This precipitation took place after coiling. Intragranular particles were mostly formed
by sulphides of manganese or complex sulphides of manganese and copper (up to 12 wt.%Cu).
The size of these particles reached up to several hundreds of nanometres. Copper rich sulphides
dissolved during hot rolling (Cu2S: Tsol = ca 950°C). A small number of precipitates consisted
of TiN or AlN particles. No Cu rich metallic particles were detected.
101
Parameters of hot rolling and 1st cold rolling + decarburization annealing are stated in
Table 2. Precipitation after decarburization annealing was much more pronounced. Grain
boundaries were decorated by both AlN and Si3N4 particles. In the GOES Si3N4 nitrides
represent a metastable phase, which is in the temperature interval of 700 – 900°C gradually
replaced by AlN phase. Intragranular precipitation of AlN was very variable. Local differences
in precipitation could be related to chemical heterogeneity of the strip inherited from hot rolling
in the two phase field. Re-precipitation of copper sulphides (Cu2S) and copper rich complex
sulphides of manganese and copper took place during decarburization annealing. The typical
size of these particles was several tens of nanometres. In many cases nucleation of AlN particles
on sulphides was observed. No copper rich metallic particles were detected. TiN particles were
not affected by decarburization annealing.
Table 2 Parameters of hot rolling and cold rolling + decarburization annealing (DCA).
Thickness mm
2.5
Hot Rolling
Tmax °C
Tmin °C
1250
970
Tcoil. °C
 600
Cold Rolling + DCA
Thickness mm Tdecarb.°C
0.65
840
Laboratory simulation of a slow heating rate to the temperature of the primary
recrystallization start (620°C) was carried out on the specimen after the 2nd cold rolling to the
final thickness of 0.3 mm. Heating rate in a protective nitrogen atmosphere was v = 25°C/h.
TEM analysis revealed identical minor phases as in the state after decarburization annealing,
no  - Cu particles were observed. However, a number density of minor phase particles
decreased. It indicates that coarsening of precipitates occurred during the slow heating to the
temperature of primary recrystallization.
Results of investigations suggest that in the 0.5 wt.%Cu – bearing GOES precipitation of
 - Cu does not play a crucial role. Copper atoms dissolve in sulphides exhibiting significantly
lower thermal stability than sulphides of manganese. Copper rich sulphides dissolve during hot
rolling and re-precipitation of copper rich sulphides takes place during decarburization
annealing. The most important inhibition phase in Cu – bearing GOES is AlN. Some nitrogen
is also bound in a metastable Si3N4 phase. Intensive precipitation of these nitrides occurs during
decarburization annealing. Slow heating to the temperature of primary recrystallization was
accompanied by coarsening of nitrides, no  - Cu particles were detected.
Acknowledgement: This paper was created in the projects FR-TI3-053 and the project
No. LO1203 “Regional Materials Science and Technology Centre – Feasibility Program”
funded by Ministry of Education, Youth and Sports of the Czech Republic.
REFERENCES
[1]
[2]
[3]
BERNIER, N., LEUNIS, E., FURTADO, C., VAN DE PUTTE, T. V., BAN, G.: EBSD Study of
Angular Deviations from the Goss Component in Grain-oriented Electrical Steel. Micron,
Vol. 54, 2013, p. 43.
LOBANOV, M.L.: Upravlenije strukturoj i teksturoj eletrotechničeskoj anisotropnoj stali,
Abstract of Doctoral Thesis, Jekaterinburg 2010, pp. 48 (in Russian).
HUMPHREYS, F. J., HATHERLY, M.: Recrystallization and Related Phenomena, Elsevier,
Amsterdam 2004.
102
CONCEPT OF DAMAGE MONITORING AFTER GRINDING
FOR COMPONENTS OF VARIABLE HARDNESS
A. MIČIETOVÁ1*, J. PIŠTORA2, Z. DURSTOVÁ1, M. NESLUŠAN1
University of Žilina, Faculty of Mechanical Engineering, Univerzitná 1, 010 26 Žilina, Slovak Republic;
email: anna.micietova@fstroj.uniza.sk
2
Nanotechnology Centre, VŠB TU Ostrava, 17. listopadu 15, 70833 Ostrava, Czech Republic
1
KEY WORDS: heat treatment, grinding, Barkhausen noise
Roll bearings are routinely heat treated in a variety of manners. Except induction or casehardening, conventional heat treatment is carried out to impart the high hardness and the
corresponding high resistance against friction and contact wear. Annealing process is always
performed after hardening to reduce the high internal stresses induced during rapid cooling.
Further heat treatment is sometimes required to enhance the toughness of hardened parts. These
parts are exposed to the elevated temperatures for a certain period within hardness of parts
decreases as a result of carbides coarsening, decrease of dislocation density and stress
relaxation. Hardness of such parts can vary in the range 38 to 62 HRC. Grinding operations are
usually involved in production of bearings to achieve the required surface roughness, shape and
dimension accuracy. Nowadays, additional requirements such as surface structure, hardness or
stress state are needed to be fully filled due to its substantial influence on functionality of parts
in operation.
Non-destructive monitoring of critical surfaces has to be carried out to reveal the parts
containing the unacceptable surface integrity. Magnetic method based on Barkhausen noise
(BN) is very often employed for such purpose, especially ground surfaces due to the high
sensitivity of BN emission to the thermally induced surface overtempering. BN originates from
irreversible Bloch Walls (BW) motion during cyclic magnetization due to existence of pinning
sites such as grain boundaries, dislocations, precipitates, other phases, etc. Ground parts can
suffer from thermally induced burn as a result of excessive heat generation in the wheel –
workpiece contact. Being so, BN emission increases in magnitude due to decreased pinning
strength of thermally softened layer produced by improper grinding as a result of carbides
coarsening and decrease of dislocation density (stress state is also altered) [1]. Thermally
softened layer contributing to the more enhanced BN signal received on the frees surface layer
can be easily contrasted and recognized when compared with untouched deeper regions in an
optical image due to reduced resistance against etching.
Concept for monitoring surface damage after grinding is based on the contrast between
poor BN emission of untouched structure and enhanced BN response (its rms value) of the
surface undergoing the elevated temperatures. This surface appears dark under optical
observation [2]. On the other hand, such concept can fail when the hardness of a component
decrease. Then the overtempering effect induced by grinding is shadowed by the previous heat
treatment regime since the both processes can represent nearly the same thermal load of the
surface. BN emission and its evolution (for instance, along with the progressive grinding wheel
wear) depends on the annealing temperature and the temperature in wheel – workpiece contact.
The higher annealing temperature is, the less pronounced contribution of the surface
overtempering induced by grinding itself would be expected which in turn correspond to the
less remarkable contrast between the deeper untouched and the near surface thermally softened
layer.
This study demonstrates that the evolution of BN and the BN features versus progressive
grinding wheel wear for components of low hardness differ from those of high hardness.
103
250
18
16
200
14
BN, mV
10
8
100
6
rms
Peak Position
50
Peak Position, ma
12
150
4
2
0
0
20
25
30
35
40
ring n.
Fig. 1. BN (rms) and Peak Position versus number of ground rings, 40 HRC.
0
5
10
15
Fig. 1 illustrates that the evolution of BN in grinding bearing steel 100Cr6 of hardness
40 HRC opposes to the progressive increase of BN emission usually obtained when parts of
hardness 62 HRC are ground. Fig. 1 shows that BN values decrease when grinding wheel wear
is more developed. It means that the surface altered by grinding process is composed of
structure of high pinning strength considering BW motion. Increasing Peak Position indicates
that effect of surface hardening dominates as a result surface heating (induced by grinding
process) followed by self-cooling. This mechanism causes increasing hardness of near surface
regions when compared to the deeper regions. Being so, the concept for monitoring surface
integrity of such surfaces via BN technique should be reconsidered.
This paper discusses the specific aspects of the surfaces heat treated to variable hardness.
Surface characterization is determined by BN technique as well as the conventional ones such
as metallographic observation, residual stresses measurement and microhardness readings.
Acknowledgement: The authors gratefully acknowledge the support by Vega project n.
1/0097/12 and Regional Centre of Excellence reg. no. CZ.1.05/2.1.00/01.0040.
REFERENCES
[1]
[2]
MOORTHY, V. et. all.: Evaluation of heat treatment and deformation induced changes in
material properties in gear steels using magnetic Barkhausen noise analysis, Conference ICBN
03, Tampere, Finland 2001.
ČILLIKOVÁ, M., MIČÚCH, M., NESLUŠAN, M., MIČIETOVÁ, A.: Nondestructrive
micromagnetic evaluation of surface damage after grinding, Manufacturing technology 2013,
Vol. 13, No.2, pp. 152-157.
104
STUDY ON RELIABILITY EVALUATION METHOD OF
ADHESION STRENGTH OF RESIN
O. HONDA1*, Q. YU1
1
Department of Mechanical Engineering, Graduate School of Engineering, Yokohama National University
Tokiwadai 79-5, Hodogaya-ku, Yokohama, Japan; email: honda-okihisa-yz@ynu.jp
KEY WORDS: resin, young’s modulus, reliability evaluation method, adhesion
The recent spread of hybrid cars and electric cars has been supported by downsizing and
technological advance of the power modules. The power modules are composed by different
materials. Thus, the mismatch by the expansion difference between materials occurs because
of temperature increase in use. In this way, delamination occurs in the interface of encapsulant
(resin) and other parts, and it becomes the reliability issue of the product. Therefore the
establishment of the high heat-resistant resin packagfing technology is in demand. This study
is aimed for proposed of the adhesion reliability evaluation method of the resin packaging.
Specifically, it was intended to establish more quantitative method by improving previous
adhesion evaluation method
The delamination load in the
conventional
resin
adhesion
evaluation method is found by
pushing with a tool, and destroying
the resin part of specimen which
bonded resin to the substrate. The
delamination stress is provided by
inputting the load into analysis. The
test like that is called "pudding-cup
test" because of form of specimen.
The specimen and state of the
examination are shown in Fig. 1.
Fig. 1. Pudding-cup test.
However, the conventional
pudding cup test can evaluate only
limited stress ratio of vertical
direction and shear direction.
Therefore, a new method that can
evaluate a ratio of various stress was
examined
by
improving
a
conventional
test
method.
Specifically, Torsion occur in the
Fig. 2. New pudding-cup test.
adhesion interface of the specimen
by attaching a jig on the specimen
like a cantilever and pushing it. The flat knob district is prepared into the specimen to attach
this jig. This test method is called "a new pudding cup test" as follows, and the state is shown
in Fig. 2.
The strength of the torsion can be controlled by adjusting the load point on a cantilever. If
the test is reproduced on analysis, an interfacial stress state at the time of the fracture becomes
clear. The result of preliminary analysis is shown in Fig. 3. From a Fig. 3, the shear stress in
105
the aspect (xz, yz plane) takes the maximum on the edge lap of the adhesion side, and the normal
stress also takes the maximum in one point in the edge lap.
Thus, the delamination
occurs in an edge lap at a
maximum point of the
normal stress. A fracture
condition is found by
plotting the normal and shear
stress at the starting point of
delamination.
Room
temperature
(25 degrees), a ceramic
substrate, epoxy resin were
used as an examination
condition in this time. Three
points 0[mm], 40[mm] and
Fig. 3. Result of preliminary analysis.
100[mm] from the center of
the pudding cup were chosen as a load point. The
universal Pulling-bending testing equipment was
used.
The plot of the delamination stress that is
found from the experiment and analysis is shown
in Fig. 4. Only the case that normal stress was
dominant was found out in the conventional
examination, but, by the new pudding cup test, the
delamination condition that the shear stress is
dominant became available to be found out. As
overall tendency, the graph is like a flat oval.
Therefore, the normal stress gives the bigger
contribution to the delamination compared to the
shear stress.
Fig. 4. Delamination condition of resin.
Thus, the packaging resin adhesion reliability evaluation method that could cover a wider
case was developed. In this time, it was carried out with room temperature, a ceramic substrate,
but the evaluation in wider condition will be possible in the future, because even if these
conditions are changed, the method of the examination does not change. If graphs such as Fig. 4
from various cases are found, it can be easily checked where delamination occurs in comparison
with a true product. Therefore, it will be able to contribute to the reliability evaluation of the
future product.
REFERENCES
[1]
[2]
MIYOSHI, T., SHIRATORI, M., ODA, J.: Daigakukiso Zairyourikigaku, Zikkyou Shuppan:
2006.
Technical Information Institute Co.,Ltd.: Jushi to Kinzoku no settyaku setsugou gijutsu,
Nihoninsatsu, 2012.
106
DIRECT BONDING OF Ti/Al BY METAL SALT GENERATION BONDING
TECHNIQUE WITH FORMIC ACID
T. AKIYAMA1*, S. KOYAMA2
1
2
Graduate School of Science and Technology, Gunma University, Japan; email: t13802002@gunma-u.ac.jp
Faculty of Science and Technology, Gunma University, Japan
KEY WORDS: surface modification, fracture, bonding strength, Ti, Al, formic acid
In the past, methods such as brazing, friction stir welding and laser welding have been used
for bonding titanium alloy and aluminum alloy. However, these techniques have some
shortcomings: (1) microcracks are developed owing to the softening of the weld zone; (2) the
gap in the weld zone results in corrosion; and (3) a high heat input is required to compensate
for high heat radiation from titanium and aluminum. Moreover, the adverse effect of halogens
present in the flux on the environment is also a cause of concern. In addition, tools used for
friction stir welding have a short lifespan and this translates to higher running costs.
Furthermore, aluminum is an excellent heat radiating and electricity conducting element;
therefore, it is difficult to bond titanium and aluminum using other welding methods. Because
of these limitations, solid-state bonding is considered to the most suitable method for bonding
materials at low temperatures. In recent years, as an alternative to fusion bonding, solid-state
bonding to assemble electronic parts have been proposed as packaging technologies for
miniaturizing medical equipment, and some progress has been made in their practical
application. Such components (low heat resistance and little mechanical strength) require a
bonding method of low bonding temperature and pressure. The problem is that at an actual
bonding surface, there is an oxide film and a machined layer. Therefore, the adhesion between
surfaces at the bond interface is the removal of the oxide film is needed to obtain high strength.
Recently, ultrasonic vibration or plasma processing has been studied as a method for breaking
and cleaning a superficial oxide film. Indeed, in an earlier study, we showed that modification
of an oxide film with formic acid greatly improves the strength of bonding between tins and tin
and copper [1, 2].
In this paper, the effect of metal salt generation processing on the bond strength of the solidstate bonded interface of titanium and aluminum has been investigated by SEM observation of
the interfacial microstructures and fractured surfaces. A cylindrical Ti specimen (Table 1) with
dimensions of φ10 mm × 20 mm and cylindrical Al specimen (99.9% purity) with dimensions
of φ20 mm × 20 mm was used in this experiment. Titanium surfaces were modified by boiling
in formic acid (FA) for predetermined time. The faying surface of the aluminum was finished
by electrolytic polishing. Solid-state
bonding was performed in N2 gas at Table 1 Chemical composition of Titanium used in this study.
C
Ti
H
O
N
Fe
bonding temperature of 733-773 K Elements
wt%
0.0012 0.109 0.004 0.034 0.004 Bal.
under a pressure of 12 MPa (bonding
time of 900 s).
Fig. 1 represents the relationship between metal salt generation processing time and tensile
strength. As shown in Fig. 1, the optimum value of metal salt generation processing time was
observed. It is inferred that the processing time is long; excessive metal salt were generated and
the processing time is short; an oxide film were remained on the bonding surface.
Fig. 2 shows the relationship between the bonding temperature and tensile strength of the
joint. In order to illustrate the effect of metal salt generation processing, the corresponding
relationship for a non-modified joint are also shown. As shown in Fig. 1, the tensile strength
107
Fig. 2. Effect of surface modification on the relation
between tensile strength of joint and bonding temperature.
σ = 11.8 MPa
T = 733 K
Ti
σ = 23.7 MPa
T = 753 K
To examine the factors determining fracture at
the bond interface, the area of the fractured surface
was observed. As shown in Fig. 3, when the surface
modification was not applied and the bonding
temperature of 733 K, substances are not found to
adhere to either surface. With a rise in bonding
temperature, substances came to be observed in a
pair in the fractured surfaces. When the metal salt
generation processing was applied (bonding
temperature of 733 K), the fractured surface started
to show ductile fracture characteristics, although it
was not observed when the metal salt generation
processing was not applied. Thus, it was
hypothesized that high-tensile strength joints were
obtained at a lower bonding temperature with metal
salt generation processing because the contact are
between atomic plane was increased in the bonding
process.
non-modified
Ti
modified (FA)
σ = 21.9 MPa
T = 733 K
increased with bonding temperature irrespective of
metal salt generation processing and approached
about 50 MPa at a bonding temperature
approximately 25 K or more lower than that for the
non-modified joint.
σ = 58.3 MPa
T = 753 K
Fig. 1. Modification time vs. tensile strength.
10 μm
Acknowledgement: This study was supported
by Grant-in-Aid for Scientific Research (c) Fig. 3. SEM images and EDX analysis results of
the fractured surfaces (Ti side).
(26820124) from Japan Society for the Promotion of
Science (JSPS).
REFERENCES
[1]
[2]
KOYAMA, S., AOKI, Y., SHOHJI, I.: Effect of Formic Acid Surface Modification on Bond
Strength of Solid-State Bonded Interface of Tin and Copper, Materials Transactions 51: 2010,
pp. 1759-1763.
KOYAMA, S., OYA, I.: Effect of Formic Acid Surface Modification on Bond Strength of SolidState Bonded Interface of Tin, J. Japan Inst. Metals 73: 2009, pp. 809-815.
108
DIRECT BONDING OF SUS304 STAINLESS STEEL BY METAL SALT
GENERATION BONDING TECHNIQUE WITH FORMIC ACID
T. TSUNETO1*, S. KOYAMA2
1
2
KEY WORDS: metal salt generation bonding, fracture, bonding strength, SUS304, formic acid
In recent years, the demand for energy-efficient devices are increasing as societies around
the world are becoming more environmentally conscious. Efforts to address this demand are
being made in various fields including the medical equipment (automated bio chemical analyser
and artificial heart-lung machine). Therefore, SUS304 stainless steel which is excellent in
corrosion resistance, toughness and workability is widely used for manufacturing medical
devices. We propose assembling medical devices using solid-state bonding rather than the
fusion bonding method; this method is suitable for miniaturized medical equipment. Earlier
research showed that surface modification decreased the bonding temperature required to obtain
a high-bond-strength joint in the solid-state bonding of Cu/Sn [1].
In this investigation, we aimed to obtain a deeper understanding of the effect of metal salt
generation bonding process on the performance of a solid-state bonded joint of SUS304
stainless steel by SEM observation of interfacial microstructure and fractured surfaces. The
specimen to be bonded was a block 20 mm × 15 mm × 5 mm and a 5 mm × 100 mm × 0.178 mm
sheet cut from SUS304 stainless steel plate. SUS304 stainless steel surfaces were modified by
boiling in formic acid (50%) for 660 s. Solid-state bonding was performed in N2 gas at bonding
temperature of 1023-1123 K under a pressure of 147 N (bonding time of 1800 s). In addition,
the specimen was readied for solid-state bonding within 180 s after metal salt generation
processing to avoid oxidation or changes in the bonding surface due to moisture absorption.
After solid-state bonding, the peel strength was evaluated using universal testing machine at
room temperature and a displacement speed of 0.017 mm/s.
Fig. 1 represents the relationship
between the bonding temperature and peel
strength of the joint. In order to illustrate the
effect of metal salt generation processing,
the corresponding relationship for an
unmodified joint and metal salt generation
processing with formic acid is also shown.
As shown in Fig. 1, the peel strength
increase
with
bonding
temperature
irrespective of metal salt generation
processing and approached 600 N at a
bonding temperature approximately 100 K
Fig. 1. Bonding temperature vs Peel strength.
lower than that for the unmodified joint. At
a bonding temperature of 1073 K or more, the peel strength was higher about 300 N than that
of the non-modified joints. Therefore, it is inferred that the effect of metal salt generation
processing were exerted at about 1073 K. Moreover, in the case of the surface were modified
and a bonding temperature of 1123 K, fractures of the base metal a part in the joint occurred.
To examine the factors determine fracture at the bond interface, the area of the fractured
surface was observed with SEM. The results are shown in Fig. 2. As shown in Fig. 2, when the
109
Absorbance(a.u.)
T = 1123 K
T = 1073 K
T = 1023 K
surface was not modified and the bonding
modified
as polished
temperature of 1023 K, substances are not
found to adhere to either the surface. With a
rise in bonding temperature, substances came
to be observed in a part in the fractured
surfaces. When the surface was modified and
the bonding temperature of 1023 K, the
fractured surface started to show ductile
fracture characteristics, although it was not
observed when the surface modification was
not applied. Thus, it is hypothesized that highpeel-strength joints were obtained at a lower
bonding temperature with metal salt
generation processing because the contact area
between SUS304 stainless steel was increased. It is
generally thought that SUS304 stainless steel
exposed to the atmosphere are immediately
10 μm
covered with an oxide film. It is also well
known that these oxides are factors that Fig. 2. SEM micrographs of fractured surfaces of joints
after peel test: as polished and modified.
obstruct the increase in bond strength. It is
known that nickel (II) formate are formed by
1650
boiling or exposing these oxide film and base
metal with formic acid. The GIRAS-IR
1350
spectrum are shown in Fig. 3. Whereas the IR
modified
spectrum of the non-modified sample shows
only spectra characteristic to siloxane
compounds as contaminations during the
as polished
polishing process, the IR spectrum of
modified sample shows IR absorption bands
700
1100
1500
1900
characteristic to carboxylate at 1350 and
-1
1650 cm-1, which indicates the existence of
Wavenumbers, cm
nickel (II) formate at the surface. It is known Fig. 3. GIRAS-IR analysis of Ni surface : as polished
and modified.
that at about 403 K, nickel (II) formate
undergoes an endothermic decomposition
reaction, as shown by following formula, to generate metallic nickel:
Ni(HCOO)2 → Ni + H2↑ + 2CO2↑.
(1)
It is therefore thought that a high-peel-strength joint was obtained at a lower bonding
temperature with metal salt generation process because metal salt such as nickel (II) formate at
the bond interface underwent a decomposition reaction during bonding.
Acknowledgement: This study was supported by Grant-in-Aid for Scientific Research (c)
(26820124) from Japan Society for the Promotion of Science (JSPS).
REFERENCES
[1]
pp. 1759-1763.
110
DIRECT BONDING OF A6061 ALUMINUM ALLOY BY METAL SALT
GENERATION BONDING TECHNIQUE WITH FORMIC ACID
Y. TOMIKAWA1*, S. KOYAMA2
1
2
KEY WORDS: metal salt generation bonding, fracture, bonding strength, A6061, formic acid
In recent years, the ease of recycling material and the demand for energy-efficient devices
are increasing as societies around the world are becoming more environmentally conscious.
Efforts to address this demand are being made in various fields including the automotive
industry where attempts have been made to reduce the weight of cars and vessel. In particular,
A6061 aluminum alloy have high strength, extrudability and are easily recyclable owing their
low melting point. Therefore, A6061 aluminum alloy is widely used for manufacturing various
structures. In solid-state bonding, materials are bonded together by applying heat and pressure
to promote interdiffusion without any significant deformation of the materials. Moreover, since
solid-state bonding involves only low bonding temperatures, the damage to a component is
lesser than that in other methods, thus making this method suitable for bonding precision
assemblies. However, in reality, a bonding surface has various defects such as surface
irregularities, an oxide film and processing layer, which act as inhibitors to successful bonding
[1]. During the early stages of solid-state bonding, surface irregularities on the bonding surfaces
from closing properly. Later, as these gaps close, the oxide films on the bonding surfaces
prevent the surface atoms from coming into contact with each other, thus lowering the bonding
strength. Therefore, it is necessary to pre-treat a bonding surface before solid-sate bonding to
remove any oxide film present on the surface.
In this paper, we examine the effect of acetic acid surface modification on the bond strength
of the solid-state bonded interface of A6061 aluminum alloy was investigated by SEM
observations of interfacial microstructures and fractured surfaces. A cylindrical Al alloy
specimen (Table 1) with dimensions of φ10 mm × 10 mm was used in this study. Before
undergoing metal salt generation, the bonding surface was polished using electrolytic polishing
method. The metal salt generation processing is carried out by boiling the aluminum alloy
surface in a 5% NaOH solution for predetermined time and in a fixed at around 373 K.
Solid-state bonding was performed in the atmosphere under the following conditions: bonding
pressure, 18 MPa; bonding time, 900 s; and bonding temperature, 693-733 K. After solid-state
bonding, the interfacial strength was evaluated using the tensile test. The tensile test was
performed using a universal testing machine at room temperature and a displacement speed
0.167 mm/s. To specify the kind of compound formed on the surface after metal salt generation
processing, the aluminum surface was identified by FT-IR.
Fig. 1 represents the
Table 1 Chemical composition of A6061 aluminum alloy.
relationship between the Elements Si
Fe
Cu
Mn Mg
Cr
Zn
Ti
Al
bonding temperature and mass% 0.68 0.30 0.31 0.11 1.00 0.16 0.05 0.02 Bal.
tensile strength of the joint.
In order to illustrate the effect of metal salt generation processing, the corresponding
relationship for unmodified joints, modified by aqueous NaOH solution joint and metal salt
generation processing with acetic acid is also shown. As shown in Fig. 1, the interfacial strength
increased with bonding temperature irrespective of metal salt generation processing and
approached a target interfacial strength (40 MPa) at a bonding temperature approximately 20 K
lower than that for the unmodified joint and modified by aqueous NaOH solution joint.
111
Moreover, at a bonding temperature of 693 K, the modified joint had interfacial strength
approximately 3 times as large as the unmodified
joint and modified by aqueous NaOH solution joint.
Therefore, it is inferred that the effect of metal salt
generation processing were exerted at about 693 K.
To examine the factors determine fracture at the
bond interface, the area of the fractured surface was
observed with SEM. The results were shown in
Fig. 2. As shown in Fig. 2, when the surface was not
modified and the bonding temperature of 693 K,
substances are not found to adhere to either the
surface. With a rise in bonding temperature,
substances came to be observed in a pair in the
fractured surfaces. When the surface was modified
by aqueous NaOH solution and metal salt generation
processing were applied (bonding temperature of
713 K), the fractured surface started to show ductile
fracture characteristics, although it was not observed
when the surface modification was not applied.
Thus, it is hypothesized that high-interfacial strength
joints were obtained at a lower bonding temperature
with metal salt generation processing because the
contact area between bonding surface was increased.
Fig. 1. Effect of surface modification of the
relation between interfacial strength and
bonding temperature.
Fig. 3 shows the FT-IR results of aluminum
surface that is modified by acetic acid for metal salt
Fig. 2. SEM micrographs of the fractured
surfaces of joint after tensile test.
generation after modifying the surface with aqueous
NaOH solution. As shown in Fig. 3, the difference in
the amount of aluminum oxide between non-metal
salt generated and metal salt generated aluminum
can also be observed. It is known that aluminum and
aluminum oxide are changed into aluminum acetate
by boiling the surface in acetic acid after boiling the
surface in aqueous NaOH solution. As for the
thermal decomposition of aluminum acetate,
literature review shows that the compound thermally
decomposed into Al2O3 granules, H2O, CO and CO2
at temperature ranging from 473-723 K. It is
therefore thought that a high-interfacial-strength
Fig. 3. FT-IR analysis results.
joint was obtained at a lower bonding temperature
with metal salt generation process because aluminum acetate at the bond interface underwent a
decomposition reaction during bonding.
REFERENCES
[1]
KOYAMA, S., KEAT, T. S., AMARI, S., MATSUBARA, K., SHOHJI, I.: Effect of Surface
Modification by Aqueous NaOH Solution on Bond Strength of Solid-State Bonded Interface of
Al, Materials Transactions 54: 2013, pp. 1975-1980.
112
EFFECT OF SURFACE MODIFICATION BY AQUEOUS NaOH SOLUTION
ON BOND STRENGTH OF A5052 ALUMINUM ALLOY/Al AND Cu/Al
X. MA1*, S. KOYAMA2
1
2
KEY WORDS: surface modification, fracture, bonding strength, A5052, Al, Cu, NaOH solution
Aluminum is widely used for manufacturing cars and electronic devices. Currently, the
most common method for bonding aluminum surfaces is brazing. However, brazing requires
positional accuracy and results in the formation of voids by the flux residue; therefore, to avoid
these problems, solid-state bonding methods are considered as a possible alternative. However,
solid-state bonding also suffers from some problems that need to be overcome. One of these
problems is the presence of an oxide film on aluminum surfaces, necessitating the need to
remove or destroy the oxide film without applying high temperature and high load. Moreover,
these techniques have some shortcomings: (1) microcracks are developed owing to the
softening of the weld zone; (2) the gap in the weld zone results in corrosion; and (3) a high heat
input is required to compensate for high heat radiation from aluminum. Furthermore, aluminum
is an excellent heat radiating and electricity conducting element; therefore, it is difficult to bond
aluminum using other welding methods. Because of these limitations, solid-state bonding is
considered to be the most suitable method for bonding materials at low temperatures [1].
In this study, a bonding
Table 1 Chemical composition of A5052 aluminum.
Fe
Cu
Mn Mg
Cr
Zn
Ti
Al
surface was treated with Elements Si
mass%
0.11
0.21
0.04
0.05
2.60
0.18
0.02
0.00
Bal.
NaOH (aq) for removing the
oxide film; moreover, the
effectiveness of this treatment was determined
by observing the bonding interfaces and
fractured surfaces of specimens. A cylindrical Al
specimen (99.9% purity) with dimensions of
φ20 mm × 15 mm, a cylindrical A5052
aluminum alloy (Table 1) with dimensions of
φ10 mm × 15 mm and cylindrical Cu specimen
(99.9% purity) with dimensions of φ10 mm ×
15 mm was used in this experiments. Before
surface modification, the bonding surface was
polished with a #800 emery paper. Surface
modification was carried out by boiling the
aluminum specimen in a 5% NaOH solution for Fig. 1. The effect of surface modification on tensile
20 s with the solution temperature fixed at
strength of joints.
around 373 K. The bonding surface was then
washed with methyl alcohol. Solid-state bonding was performed in N2 gas at bonding
temperature of 753 K under a pressure of 12 MPa (bonding time of 900 s). Furthermore,
grazing-angle incidence reflection-absorption infrared (GIRAS-IR) spectroscopy was
employed to obtain IR spectra for the modified aluminum surface at nanometer-scale depth.
The spectra were obtained using a Fourier transform infrared (FTIR) spectrometer (Thermo
Fisher Inc., Magna-750) equipped with a mercury-cadmium-telluride (MCT) detector using a
single reflection accessory (Harrick Inc., Seagull) at an incidence angle of 80°. An Au-coated
mirror surface was used as the reflecting surface and measurements were carried out at a
113
wavenumber resolution of 8 cm-1. Moreover, the modifying actions of NaOH (aq) on the
aluminum surface, after boiling a 10-µm-thick aluminum foil in NaOH (aq) for 30 s, we
immediately subjected the specimen to differential scanning calorimetry (DSC, SII Seiko
Instruments, DSC6200).
Fig. 1 shows the effect of surface modification
on the bond strength of joint. As shown in Fig. 1,
the tensile strength increased with the surface
modification irrespective of bonding material.
However, when surface modification was applied
for the joint between A5052 aluminum alloy and
aluminum, the A5052/Al joint had a bigger ratio of
rise in strength than the Cu/Al joint. On the basis of
these results, it was inferred that the Cu/Al joint had
low tensile strength because of the formation of an
oxide film between the bonding surfaces, inhibiting
intimate contact between the atomic planes of Cu
and Al. It can be explained from difference in
volume of deformation of the joint after solid-state
bonding.
Fig. 2. FT-IR analysis results of the aluminum
surfaces with/without surface modification.
The GIRAS-IR spectra are shown in Fig. 2.
The IR absorption bands at about 3400 cm-1
correspond to a hydroxyl group, those at 1600 cm-1
correspond to hydrogen carbonate and those at less
than 950 cm-1 correspond to aluminum oxide. The
results shown in Fig. 2 revealed that the oxide
formation was decreased and Al(OH)3 was
generated upon surface modification. From this, it
can be inferred that the bonding surface of the
specimen used in this study is also covered with
Al(OH)3 when it was washed with methanol after
NaOH (aq) surface modification. Moreover, it is
also known from other studies that Al(OH)3
decomposed into particles of Al2O3 and H2O by an Fig. 3. DSC analysis results of the aluminum
sheet with surface modification.
endothermic and dehydration reaction at about
573 K. Actually, as shown in Fig. 3, the generation
and thermolysis of Al(OH)3 were supported by a bigger endothermic peak as recognized by a
gentle endothermic peak. Because the temperature used in solid-state bonding in this study was
at least 573 K or more, it can be inferred that the thermal decomposition of Al(OH)3 to H2O
(gas) and particles of Al2O3 occurred during the bonding process, thus resulting in the exposure
of atomic planes of aluminum. This caused the tensile strength of the joint that undergone
surface modification to increase.
REFERENCES
[1]
114
DIRECT BONDING OF Cu/Cu BY METAL SALT GENERATION BONDING
TECHNIQUE WITH FORMIC ACID AND ACETIC ACID
S. KOYAMA1*, N. HAGIWARA2, I. SHOHJI1
1
2
Faculty of Science and Technology, Gunma University, Japan; email: koyama@gunma-u.ac.jp
Gracuate School of Engineering, Gunma University, Japan
KEY WORDS: metal salt generation bonding, fracture, bonding strength, copper, organic acid
As an alternative to fusion bonding, solid-state bonding to mount electronic parts have been
proposed as packaging technologies for miniaturizing electronic devices. Meanwhile, a number
of mounting parts made of resin have low heat resistance and little mechanical strength. Such
components require a bonding method of low bonding temperature and pressure. The problem
is that at an actual bonding surface, there is an oxide film and a machined layer [1]. Therefore,
the removal of the oxide film is needed to obtain high bond strength. Recently, ultrasonic
vibration or plasma processing has been studied as a method for breaking and cleaning a
superficial oxide film. Indeed, in an earlier study, we showed that modification of an oxide film
with formic acid greatly improves the strength of bonding of tin and copper [2].
In this paper, we examine the effect of organic acid surface modification on the bonding
strength of the solid-state bonded interface of copper was investigated by SEM observations of
interfacial microstructures and fractured surfaces. The specimen to be bonded was a block
15 mm × 15 mm × 5 mm cut from 99.9% ingot and a wire (99.9% purity) dimension of
φ1.2 mm. Copper surfaces were modified by boiling in formic acid and acetic acid for
predetermined time. Solid-state bonding was performed in a vacuum chamber at bonding
temperature of 423-673 K under a pressure of 588 N (bonding time of 0.9 ks).
Fig. 1 represents the relationship between the bonding temperature and peel strength of the
joint. In order to illustrate the effect of metal salt generation processing, the corresponding
relationship for a non-modified joint are also shown. As shown in Fig. 1, the peel strength
increased with bonding temperature irrespective of metal salt generation processing and
approached about 30 N at a bonding temperature approximately 150 K (with formic acid) and
100 K (with acetic acid) or more lower than that for the non-modified joint.
To examine the factors determining
fracture at the bond interface, the area of the
fractured surface was observed. As shown in
Fig. 2, when the surface was not modified,
approximately
0.5 µm-diameter
copper
oxides were observed at the fractured
surfaces. In addition, ductile fracture
characteristics such as tear ridges were not
observed. When the surface was modified
with formic acid and acetic acid, the fractured
surface started to shown ductile fracture
mode, although it was not observed when the
Fig. 1. Effect of surface treatment: relation between
surface modification was not applied. At the
bonding temperature and peel strength.
same time, copper oxide was not observed at
the fractured surfaces. From these results, it was inferred that the joints bonded at low bonding
temperatures had high peel strength because of the removal of oxide film, thus allowing
intimate contact between the atomic planes of copper.
115
It is generally thought that copper exposed
to the atmosphere are immediately covered
with an oxide film. It is known that copper (II)
formate and copper (II) acetate are formed by
boiling the oxide film and the base metal with
formic acid and acetic acid for a long time.
Therefore, it is thought that at least
Cu(HCOO)2 and Cu2(CH3COO)4 were formed
on the surface layer when metal salt generation
process was applied. Moreover, it is known
that copper (II) formate and copper (II) acetate
undergoes an exothermic decomposition
reaction and metallic copper is generated at a
temperature using this study. It is therefore
thought that a high tensile-strength joint was
obtained at lower bonding temperature with
surface modification because copper (II)
formate and copper (II) acetate at the bond
interface underwent a decomposition reaction
after close contact between Cu/Cu was
achieved.
Fig. 3 represents the relationship between
the shelf time of the modified surface and peel
strength of the joint. In case the surface is
modified by formic acid, the peel strength was
decreased in several hours, but the case the
surface is modified by citric acid, the peel
strength was not decreased even for 168 hours.
Generally, it was understood that copper
formate is easy to dissolve in water but copper
acetate is hard to dissolve in water. Thus, it can
be inferred that a bonding surface modified
with acetic acid is also protected from
reoxidation.
Fig. 2. SEM micrographs of the fractured surfaces of
joints after tensile test with or without surface
modification.
Fig. 3. Effect of surface treatment: relation between
shelf time and peel strength.
REFERENCES
[1]
[2]
pp. 1759-1763.
116
DELAMINATION PROPERTY OF MODELLED AIR PLASMA SPRAYED
THERMA BARRIEAR COATINGS: EFFECT OF DIFFERENCE IN
CHEMICAL COMPOSITION OF BOND COAT
M. HASEGAWA1*, S. YAMAOKA1
1
79-5 Tokiwadai, Hodogaya-ku, Yokohama, 240-8501, Kanagawa, Division of Materials Science and
Chemical Engineering, Faculty of Engineering, Yokohama National University, Japan;
email: hasegawa@ynu.ac.jp
KEY WORDS: microstructure, yield stress, delamination toughness
Thermal barrier coatings (TBCs) have been widely used in order to increase the operating
temperature of hot section components in gas turbine blades and vanes [1]. TBCs are usually
composed of outer ceramics thermal barrier coating (TBC) layer and inner intermetallic bond
coat (BC) layer to protect the nickel base superalloy from high temperature and oxidation.
During the service, formation and growth of thermally grown oxide (TGO) occurs between
TBC and BC layer. This growth increases the thermal stress of TBCs, and finally, the increased
thermal stress results in the delamination of the coating. In order to understand the effect of the
difference in chemical composition and microstructure of BC layer on delamination properties,
mechanical properties are examined on modelled TBCs.
Nickel-platinum-aluminides (Ni-Pt-Al) and NiCoCrAlY alloy were selected as BC alloy.
Chemical compositions of BC alloys were Ni-43Al-9Pt, Ni-42Al-9Pt-0.3Hf and
Ni-25Al-19Co-16Cr-0.4Y (mol%). BC alloys were heat treated in a vacuum at 1413 K for
1 hour. After the treatment, modelled TBCs have been formed by air plasma-spraying process.
TBC layer of an 8 mass% Y2O3 partially stabilized ZrO2 was coated on the BC alloy in 250 m
thick. After the process, the TBCs were heat exposed in an air at 1323 K from 10 to 100 hours.
Changes in microstructure during heat exposure were characterized on the polished transverse
section of the TBCs by SEM and EBSD. Yield stresses of BC alloys were decided from the
result of Vickers hardness measurement. To evaluate the delamination toughness of the TBCs
under shear loading condition, pushout tests were performed [2, 3].
Ni-Pt-Al and NiCoCrAlY BC alloys are consist of  single phase and (’) two phases,
respectively.  phase of NiCoCrAlY BC alloy near the TGO/BC interface disappears during
heat exposure, due to the formation and growth of the TGO. In case of the TBCs with
NiCoCrAlY BC alloy, yield stress of the BC alloy decreases with the increase in heat exposure
time. However, as for the TBCs with Ni-Pt-Al BC alloy, yield stress was almost constant
independent of the heat exposure time. TGO thickness increases with the increase in heat
exposure time. TBCs with Ni-Pt-Al alloy shows thinner TGO thickness than that of the TBC
with NiCoCrAlY alloy in a same heat exposure time. Fracture of the TBCs occurs mainly at
TBC/TGO interface. TBCs having Ni-Pt-Al BC alloy shows higher shear strength and
delamination toughness than that of the TBCs with NiCoCrAlY BC alloy.
REFERENCES
[1]
[2]
[3]
PADTURE, N. J., GELL, M., JORDAN, E. H.: Finite Thermal Barreir Coatings for GasTurbine Engine Applications, Science 296, 2002, pp. 280-284.
KIM, S. S., LIU, Y. F., KAGAWA, Y.: Evaluation of Interfacial Mechanical Properties under
Shear Loading in EB-PVD TBCs by the Pushout Method, Acta Mater. 55, 2007, pp. 3771-3781.
HASEGAWA, M., ENDO, T., FUKUTOMI, H.: The Effect of Microstructure Change of Bond
Coat Layer in Air Plasma-Sprayed Thermal Barrier Coating System on Interfacial Mechanical
Property under Shear Loading, J. Japan Inst. Metals 73, 2009, pp. 802-808.
117
118
MICROSTRUCTURE MODIFICATION OF CGDS AND HVOF SPRAYED
CoNiCrAlY BOND COAT REMELTED BY ELECTRON BEAM
P. GAVENDOVÁ1*, J. ČÍŽEK1, J. ČUPERA1, M. HASEGAWA2, I. DLOUHÝ1
1
Brno University of Technology, Institute of Materials Science and Engineering, NETME centre,
Technická 2, Brno, Czech Republic; email: gavendova@fme.vutbr.cz
2
Yokohama National University Faculty of Engineering, Division of Materials Science and Chemical
Engineering, Yokohama, 240-8501, Japan
KEY WORDS: CoNiCrAlY, bond coat, thermal spray, cold kinetic deposition, electron beam
remelting, thermal barrier coating
The efficiency of gas turbine engines can be significantly improved by increasing their
operating temperatures. The engine components, especially the hot-end components, should
maintain their mechanical integrity at high-temperatures. The complex environment of the gas
turbine engine makes it extremely difficult for one material to meet all the different
requirements imposed on the various engine components. Coatings are thus required to provide
increased corrosion (and oxidation) protection at high temperatures in order to assure durability
and field performance of the base alloys. Ni-base turbine blades are generally protected by a
CoNiCrAlY overlay and a ceramic thermal barrier top layer, both thermally sprayed. The
method used for these coatings is usually high-velocity oxygen-fuel (HVOF) spraying, other
methods can be also applied however.
Promising results can be obtained with cold kinetic deposition method and/or cold gas
dynamic spraying (CGDS) [1]. It is a method that enables to get coatings with lower porosity,
higher relative density and better adhesion of the (bond) coating to substrate. From that point
of view it seems that this technique possesses a better potential of for bond coat fabrication
comparing to HVOF etc.
A subsequent remelting of the HVOF by electron beam (EB) irradiation opens another
opportunity to minimize the oxide content (because of vacuum condition needed for EB
technology) and the porosity (thanks to remelting and the melt rapid cooling) of the thermally
sprayed CoNiCrAlY bond coatings. The electron beam remelting process appears to be one of
the most convenient processes to reduce or fully remove the disadvantages of thermal spray
coatings. The bond coats remelted by electron beam produce modifications in the morphology
and phase composition [2] that could be further exploited for TBC performance optimisation.
In the present work two techniques are combined to optimise bond coat properties before
TBC application, the cold gas dynamic spraying (CGDS) and electron beam remelting (EB).
Results of the work focused on comparison of HVOF and CGDS CoNiCrAlY bond coats are
firstly presented. Than the effect of the electron beam remelting of the CoNiCrAlY coating
manufactured by HVOF and CGDS deposition techniques is deeply investigated. Scanning
electron microscopy, light microscopy, and, in addition, X-Ray diffraction techniques were
performed to characterize the phase modification and microstructure composition changes
before and after the treatment. The microstructural and phase analyses have been supported by
microhardness and nanohardness investigation and other necessary supporting techniques.
The bond coat having thickness of about 70 m prepared by both HVOF and CGDS
technique displayed the lower porosity for the CGDS microstructure. The CoNiCrAlY bond
coat to Inconel substrate interface displayed locations with very poor bonding, in larger extent
for the states prepared by HVOF comparing to CGDS. The bond coats prepared by both ways
being EB remelted up to depth of about 90 m are typical by removal of the defects on the
119
substrate to bond coat interface. The microstructure of the bond coat after this treatment has
been is formed by Inconel fine grain layer in depth of 50 m being followed by the surface
layer consisting of elongated dendritic microstructure. The longitudinal axis of dendrites has
been oriented predominantly perpendicularly to the Inconel surface. An increased porosity has
been observed in interdendritical space in larger extent for CGDS samples.
The results obtained thus showed that using the pulsed electron beam surface modification
technique produced positive changes in the bond coat layer as a necessary step for the thermal
barrier coating fabrication. In addition the EB treatment provided a smooth surface, low
porosity level comparing to bond coat surface without this modification. Although the
technology parameters window for successful application of the CGDS BC application and
subsequent EB remelting has been shown to be relatively narrow this procedure supplied good
basis for TBC application when compared to standard HVOF technique and subsequent vacuum
ageing. From the results, it appeared also that there is a potential for improvements of the bond
coat deposition process when applying low-temperature processing methods such as CGDS.
Thermal barrier coatings (TBC) that have been applied on the Inconel substrate with
modified bond coat surface suppose an effective engineering solution for the improvement of
in service performance of gas turbines components. The quality and further performance of
TBC, likewise all thermally sprayed coatings is strongly dependent on the adhesion between
the coating and the substrate as well as the adhesion between the metallic bond coat and the
ceramic top coat layer. The debonding of the ceramic layer or the bond coat layer will lead to
the collapse of the overall thermal barrier system. Though several possible problems can occur
in coating applications like residual stresses, local or net defects (like pores and cracks), one
could say that a satisfactory adhesion is the first and intrinsic need for a good coating. The
coating adhesion is also dependent on the pair substrate-coating materials, substrate cleaning
and blasting, coating application process, coating application parameters and environmental
conditions.
Acknowledgement: The works have been supported by the financial support from the
Operational Programme Education for Competitiveness No. CZ.1.07./2.3.00/30.0005 and
within the project NETME plus centre (Lo1202), project of Ministry of Education, Youth and
Sports under the “national sustainability programme”. Support of Czech Science Foundation
project GACR 13/35890S is further acknowledged.
REFERENCES
[1]
[2]
[3]
[4]
UTU, D., MARGINEAN, G., BRANDL, W., CARTIS, I.: Improvement of the oxidation
behaviour of electron beam remelted MCrAlY coatings, Solid State Sciences 7, 2005,
pp. 459-464.
UTU, D., BRANDL, W., MARGINEAN, G CARTIS, I., SERBAN, V.A.: Morphology and
phase modification of HVOF-sprayed MCrAlY-coatings remelted by electron beam irradiation,
Vacuum 77, 2005, pp. 451-455.
LIMA, C.R.C., GUILEMANY, J. M.: Adhesion improvements of Thermal Barrier Coatings
with HVOF thermally sprayed bond coats, Surface & Coatings Technology 201, 2007,
pp. 4694-4701.
RICHER, P., YANDOUZI, M., BEAUVAIS, L., JODOIN, B.: Oxidation behaviour, of
CoNiCrAlY bond coats produced by plasma, HVOF and cold gas dynamic sprayings, Surface &
Coatings Technology 204, 2010, pp. 3962-3974.
120
RESPONSE OF ALUMINA FOAM TO TENSILE MECHANICAL LOADING
INCLUDING STRESS CONCENTRATOR EFFECT
I. DLOUHÝ1*, Z. CHLUP1, H. HADRABA1, L. ŘEHOŘEK2
1
Institute of Physics of Materials, Academy of Sciences of the Czech Republic, CEITEC IPM, Zizkova 22,
61662 Brno, Czech Republic; email: dlouhy@ipm.cz
2
Military Research Institute, Division of Materials Engineering, Brno
KEY WORDS: tensile test, ceramics foam, open porosity, tensile strength, compression strength
Ceramic foams with open cell porosity are of technological interest because of their
potential use in a number of industrial fields. Applications like catalytic substrates [1], high
temperature filters for melted alloys [2], in tissue engineering as a bone replacement material
[3, 4], insulation materials [5] etc. require high permeability, high surface area and good
insulation characteristics, but also a good response to different types of mechanical loading
typical for given applications. For ceramic foams, mechanical behaviour including the fracture
resistance of thus plays an important role in their potential applications.
At present, it is common to estimate the mechanical performance of ceramic foams by
compressive tests only [5]. The most frequently used mechanical parameter is a
compressive/crushing strength that is evaluated from the compression test curve either as
maximum force at the first relevant peak or as average force of plateau observed on the curve.
Very limited number of works has attempted also a modified bending test [6]. The most
complication supposes the measures to avoid to the crushing between the rollers and the tested
ceramics foam. Suitable thin sheet must be applied into interface between the roller and
specimen surface. It must be rigid and tough but not too much to assure load transfer [7].
Data of tensile tests of the ceramic foams are completely missing in the literature. For
interpretation of the mechanical behaviour and modelling the ceramic foam response [8] in the
given applications some material data are needed however. The aim of the paper thus can be
seen in summarisation of the knowledge obtained with the tensile test of ceramics foams of
different samples geometry, interpretation of data obtained and possible consequences relating
to foam material mechanical behaviour. Tensile loading of ceramic specimens brings
difficulties. It is necessary to solve efficient load transfer and ensure alignment of the specimen
with loading axis of the system which is not simple task. Brittleness of this material brings
difficulties with fixation of material into claws. It is impossible to use any fixing methods using
compression, friction, threaded joint and their combinations. Only one possibility is to employ
adhesion evoked by some kind of adhesive or resin.
Material used in the investigation was alumina based foam (85 vol. % Al 2O3, 14 vol. %
SiO2, 1 vol. % MgO) commercially produced (Vukopor®A) by company Lanik typically used
e.g. for aluminium alloys melt filtration. The foam structure was produced by replication
technique consisting in slurry coating of polyurethane foam. Two types of cell sizes were
applied for investigations, 2.2 (±1.2) mm and 0.8 (±0.3 mm), respectively. The dimensions of
test specimens with both porosity types (10 and 60 PPI) were 10x10x30 mm3 and
15x15x40 mm3 respectively. Samples containing central internal sharp notch located in
10x30 mm cross-section have been also included in investigation.
Special fixture and testing rig was developed to assure transfer of the load to the ceramic
foam specimen. Tensile strength values were determined as maximum force from the loading
diagram related to cross-sectional area of the specimen at fracture.
121
As expected, the scatter of apparent tensile strength values is affected by specimen size.
Specimens with cross section of 15x15 mm2 showed noticeable lower data scatter because of
larger sampling volume averaging the structural heterogeneity. As shown also elsewhere [7, 9]
noticeable scatter of apparent tensile strength is caused (i) by the heterogeneity in distribution
characteristics of the cell sizes in the given sample and, in addition, (ii) by fatal macroscopic
material defects typical for this kind of material and fabrication technology applied. The
observed relations between the ceramic foam microstructure and measured apparent tensile
strength confirm very good susceptibility of the measurements to microstructural differences
however.
The specimen size (in the range of dimensions applied in this investigation) does not affect
substantially the average values of tensile strength as it has been shown based on the Weibull
statistical analysis. From the data sets for the specimen sizes applied and for both cell sizes,
having 10 and 60 PPI, it is obvious that there are no substantial differences in the quality of data
obtained from smaller and larger specimens. Both specimen geometries, having larger and
smaller cross-section exhibit the same slope in Weibull distribution. Specimens with higher PPI
have lower apparent tensile strength comparing to those with lower PPI. The higher cell size
the higher apparent tensile strength is. This is given mainly by strut thickness of the cell which
is significantly lower in samples with 60 PPI comparing to the 10 PPI ones.
The samples with central crack (sharp notch) have included into investigation
experimentally to determine the possible stress concentration effects in open cell porosity
structures. Data from pre-cracked samples were compared with results of samples having the
same cross-section as remaining part of section in the pre-cracked samples. For the pre-cracked
samples, the strength values have been found lower comparing to samples without crack.
Independently of quite large open porosity there is still certain stress concentration effect in
these structures. This must be taken into account in the most critical structural applications.
The quantitative data from the tests supported well the corresponding fractography
observations. Based on the analyses carried out and supporting finite element modelling critical
condition for the struts fracture have been established.
Acknowledgement: The works have been realised in CEITEC centre - infrastructure
supported by the project CZ.1.05/1.1.002/02.0068 financed from Europepan Regional
Development Fund. Support of Czech Science Foundation project GACR 14-11234S is further
acknowledged.
REFERENCES
[1]
[2]
[3]
[4]
[5]
[6]
[7]
[8]
[9]
GARRIDO G. I., PATCAS F.C., UPPER G. ET AL.: Applied Catalysis A: General 338 (2008)
58.
PATCAS F.C., GARRIDO G.I., KRAUSHAAR-CZARNETZKI B., Chemical Engineering
Science 62 (2007) 3984.
MIAO X., TAN L.P., TAN L.S., HUANG X.: Materials Science and Engineering C 27 (2007)
274.
REHOREK L. CHLUP Z., MENG D., et al.: Ceramics International 39 (2013) 8015–8020.
GIBSON L. J., ASHBY M.F.: Cellular Solids, Cambridge University Press, Cambridge, 1999.
BREZNY R., GREEN D.J., Acta Metall. Mater. 38 (1990) 2517.
REHOREK L., DLOUHY I., CHLUP Z. et al.: Ceramics – Silikáty 53 (4) (2009) 237-241.
MARCIAN P. MAJER Z., DLOUHY I., FLORIAN Z.: Chem. Listy Vol. 106 (2012),
pp. 476-477.
DLOUHY, I., REHOREK L., CHLUP Z.: Key Eng. Mat. Vol. 409 (2009), pp. 168-175.
122
NONDESTRUCTIVE MAGNETIC MONITORING OF GRINDING DAMAGE
M. ČILLIKOVÁ1*, B. MIČIETA1, M. NESLUŠAN1, D. BLAŽEK2
University of Žilina, Faculty of Mechanical Engineering, Univerzitná 1, 010 26 Žilina, Slovak Republic;
email: maria.cillikova@fstroj.uniza.sk
2
Nanotechnology Centre, VŠB TU Ostrava, 17. listopadu 15, 70833 Ostrava, Czech Republic
1
KEY WORDS: grinding, Barkhausen noise, surface damage, wear, white layer
Ground surface can suffer from over tempering or overheating due to elevated temperatures
in the wheel – workpiece contact. For this reason, thermally softener or/and re-hardened layers
can occur in the near or sub surface regions. Surface containing the untempered brittle
martensite or thermally softened structure are assumed detrimental to components life due to
early crack initiation and premature failure. Surface integrity expressed in such term as residual
stresses, hardness and the corresponding structure alteration can varies in spite of the cutting
conditions are kept constant. Thus implementation of reliable monitoring concept should be
proposed for the real industrial applications.
Production of large bearing for wind power station involves grinding cycles as the final
operation substantially contributing to the functionality of bearings. Bearings undergo very
rigorous monitoring procedures to reveal unfavourable stress state, structure modifications or
crack initiation originating from grinding cycles or heat treatment. Manufacturer guarantees at
least 20 years failure – free operation of bearing. Monitoring of bearings (mainly raceways) of
diameter in the range 600 to 4000 mm after grinding does not allow implementation of chemical
activation to reveal the unfavourable surface alterations induced by grinding.
Magnetic Barkhausen noise (BN) has found the high industrial relevance for
characterization of ferromagnetic materials. BN originates from irreversible domain and mainly
Bloch Wall (BW) motion during cyclic magnetization. Pulsating magnetization on hysteresis
loop occurs as a result of BW interaction with stress fields and microstructure features such as
dislocations, secondary phases or grain boundaries hindering the smooth BW motion [1]. BW
jumps occur as soon as the strength magnetic field exceed the critical value equal the pinning
strength of the abovementioned pinning sites. BN is the stress and microstructure sensitive
technique. However, while the stress state affects mainly domain and the corresponding BW
alignment, microstructure features affect the free path of BW motion [2]. Although variety of
BN applications were reported, monitoring of grinding burn still dominates as a reliable method
to reveal surface over tempering. While untouched surface emits low BN value, enhanced BN
emission in thermally softened layers is due to thermally induced decrease of dislocation
density, precipitates coarsening and tensile stresses [2].
This paper reports about adoption of BN technique for surfaces of large diameter (for wind
power stations) after grinding. The paper researches the factors taking a key role in grinding
cycles based on BN technique and the corresponding conventional techniques for surface
inspection. To explain the significance of such approach the two main factors of grinding
operations are discussed as follows: grinding wheel wear and the lack of coolant. First aspect
is associated with higher BN values obtained after grinding of bearings of higher diameter,
while the second one represent the unexpected lack of coolant supply randomly occurring in
production.
123
Fig. 1. Influence of grinding wheel wear and lack of coolant on BN and the corresponding micrographs.
Fig. 1 demonstrates the influence grinding wheel wear and lack of coolant on BN.
Progressive grinding wheel wear enhances thermal load of ground surfaces. Higher
temperatures penetrate deeper beneath the surface; thickness of heat affected zone (HAZ)
contributing to the BN emission more than untouched structures values is increasing. Moreover,
compressive stresses induced in the earlier grinding cycles are shifted to the tensile ones. Effect
of grinding wheel wear takes the main role considering higher BN values indicated for bearings
of higher diameters due to higher volume material removed and higher allowances for the
finishing grinding operations, both contributing to the more developed grinding wheel wear in
the final phases of grinding cycles.
On the other hand, effect of coolant supply is more remarkable than that associated with
grinding wheel wear. Lack of coolant is more risky when the grinding wheel is more developed
(during grinding rings of higher diameter - see the red and blue lines in Fig. 1) than the absence
of coolant in the early phases of grinding cycle or grinding smaller rings (see the black line in
Fig. 1). The abrupt increase of BN values in dry grinding is due to accumulation of heat in the
ground surface which in turn corresponds with quite thick HAZ. The following steep decrease
of BN values is attributed to the surface over tempering when austenitizing temperature is
exceeded. The following self-cooling effect results in formation white layer. Surface cracking
which occurs together with white layers is due to high internal stresses induced by rapid cooling.
Implementation of BN technique in the abovementioned production has found the high
relevance serving the technologist and grinding staff to modify the grinding cycles and
eliminates the risky factors of grinding operations.
Acknowledgement: The authors gratefully acknowledge the support by Vega project n.
1/0701/12 and Regional Centre of Excellence reg. no. CZ.1.05/2.1.00/01.0040.
REFERENCES
[1]
[2]
GATELIER-ROTHEA, C., CHICOIS, J., FOUGERES, R., FLEISCHMANN, P.:
Characterization of pure iron and carbon-iron binary alloy by Barkhausen noise
measurements: study of the influence of stress and microstructure, Acta Mater. 46/14, 1998,
pp. 4873-4882.
MOORTHY, V. et all.: Evaluation of heat treatment and deformation induced changes in
material properties in gear steels using magnetic Barkhausen noise analysis, ICBN 03,
Tampere, Finland 2001.
124
Autor
Katedra, institut:
Název:
Místo, rok, vydání:
Počet stran:
Vydala:
Tisk:
Náklad:
Neprodejné
ISBN 978-80-248-3488-7
Prof. Ing. Bohumír Strnadel, DrSc.
Katedra materiálového inženýrství
636
NEW METHODS OF DAMAGE AND FAILURE ANALYSIS
OF STRUCTURAL PARTS
Ostrava, 2014, VI.
124
VŠB-TECHNICKÁ UNIVERZITA OSTRAVA
Printo spol. s r.o.
150

Conference Proceedings

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