Ref. p. 6]
1 Introduction
1
1 Introduction
K. YAGI, G. MERCKLING, T.–U. KERN
1.1 General remarks
To utilize energy resources efficiently and preserve the global environment, efforts are being made to
raise the temperature at which high-temperature equipment is used at power and chemical plants. As a
result, the conditions to which the structural components of these plants are exposed have become much
more demanding. Structural components have been developed that endure these extreme conditions and
some of them are now coming into use. It is therefore necessary to ensure the effective and safe use of
these materials, to gain a full understanding of the characteristics of the new structural components, to
evaluate their strengths, and to predict their life with greater accuracy. On the other hand, many of the
world’s high-temperature plants were constructed as long ago as the 1970s and have deteriorated
markedly with age. These aged plants are sometimes used under operating conditions different from those
planned at the time of their construction. Therefore, the key issue is to be aware of the changes that take
place in materials for structural components over time and to predict their remaining life with high
accuracy.
With regard to the prediction of the life of high temperature structural materials, the evaluation of their
deterioration with time and the prediction of the remaining life of aged materials, it is essential to have a
full knowledge of the characteristics of these materials and to be aware of the existence of material data
that is a source of that knowledge. Creep characteristics are typical properties of high-temperature
structural materials. Because the creep characteristics of structural materials in high-temperature plants
are understood to be important in the design of boilers and pressure vessels, creep tests have been actively
conducted since the 1930s [1, 2]. Creep data have been systematically obtained and published in the
United States and European countries such as UK, Germany and Italy since the end of World War ll.
Large-scale facilities and operating funds are required to obtain and register creep data, and in recent
years it has become increasingly impractical for single organizations to be tasked with collecting
systematic and long-term data. For this reason, it is increasingly important to share basic data and
knowledge at the international level
With the beginning of a new Landolt-Börnstein data collection series, the present data book was
planned and compiled through the cooperation of European Creep Collaborative Committee (ECCC),
German Creep Committee (GCC) and National Institute for Materials Science, Japan (NIMS). The
purpose of the data book was to collate previously obtained creep data on major heat resistant steels and
alloys as well as knowledge concerning creep characteristics. It could then serve as a basis for
technological development to predict the life of structural materials, evaluate their deterioration with age
and predict their technically usable life, as well as act as a resource for the design of safe structural
components and safe maintenance of plants. More than four years were spent from planning to
completion of this data book. However, this is a short period of time compared to the 10 years or more
required to obtain service relevant creep data. The editors of this data book hope that it will help in the
development of new technology as well as in the design and maintenance of safe power plants and
comparable applications.
2
1 Introduction
[Ref. p. 6
1.2 Status of creep database
Research institutes, academic societies and private industries have collected and organized up to now
creep data independently or through coordination as a group. Because of the need for special facilities and
of constraints on funding and time, most long-term creep data have been collected on a national level or
through programs run by academic societies. Typical data series published internationally are outlined
below.
1.2.1 The ASTM data series
The American Society for Testing and Materials (ASTM) published a collection of data on hightemperature strength as part of the Special Technical Publication (STP) series in the 1950s. Nearly 50
volumes have been published for the Data Series (DS) [3]. The features of this series are that the editors'
analyses and the results of their evaluations are included in each volume. The description of data has not
followed a prescribed format.
1.2.2 The BSCC high-temperature strength data series
This is a data series compiled by the British Steelmakers Creep Committee (BSCC) in 1972, under the
leadership of the British Steel Corporation [4]. The data series shows the results of high-temperature
tensile testing and creep rupture testing on representative materials such as carbon steel, alloy steel and
austenitic steel and follows a stipulated format.
1.2.3 The long-term data series by the Iron and Steel Institute of Germany
The German Creep Committee with it’s secretary of Verein Deutscher Eisenhüttenleute (VDEh) compiled
creep data on heat resistant steels in 1968 [5]. Data collection was made by a joint working group of
steelmakers and equipment manufacturers. Data on carbon steel, low alloy steel, 12Cr steel and stainless
steel are shown in a fixed format. Many years later, a data series on cast steels [6] and heat resistant alloys
[7] were jointly published by Forschungsvereinigung Warmfeste Stähle (FWS) and Forschungsvereinigung Hochtemperaturwerkstoffe (FVHT) of VDEh and Forschungsvereinigung Verbrennungskraftmaschinen e.V. (FVV) in 1986 and 1987, respectively.
1.2.4 European Creep Collaborative Committee (ECCC)
The ECCC was established to jointly acquire, collate and analyze creep data on metallic materials for
high temperature plants in the European community in 1992. Actually 14 nations are members of the
ECCC, including Germany, United Kingdom, Italy, France, Sweden, Denmark, Finland, Belgium, The
Netherlands, Portugal, Austria, Switzerland, Czech Republic and Slovakia. The ECCC aims to harmonise
and encourage European creep data generation, provide creep and creep rupture strength data as well as
design relevant information to European standards, exchange information on material development, and
develop rules for data generation, exchange and assessment [8]. Target materials are carbon steel and low
alloy steel, 9-12% Cr steel, austenitic stainless steel, welded joints, bolts, and Ni base alloys. While
Landolt-Börnstein
New Series VIII/2B
Ref. p. 6]
1 Introduction
3
experimental data have not been made public, the results assessed with proceduralised methods (e.g. BS
PD 6605) and validated with the aid of innovative credibility checks [8] were presented to the public in
1999 [9].
Newer European standards, like EN 10028, EN 10216, EN 10222, contain creep strain and creep
rupture strength data assessed by ECCC and derived from all over Europe, sometimes all over world
collated experimental results.
1.2.5 Report on the mechanical properties of metals at elevated temperatures by
the Iron and Steel Institute of Japan
The High-temperature Research Committee (formerly called the Creep Committee) collected data and
published five volumes of data series covering low alloy steel, stainless steel, carbon steel and cast iron,
heat resistant alloys and welded joints [10].
1.2.6 Creep data sheet published by NIMS (formerly called NRIM)
In 1966, the National Institute for Materials Science (NIMS) launched a 100,000-hour creep rupture
strength testing project on domestically produced high-temperature metallic materials. The results of
these creep tests series were summarized in 49 kinds of NIMS (formerly called NRIM) creep data sheets
and published in 122 volumes up to 2003 [11]. Although still in progress, the project is one of the largest
in the world planned to obtain creep test data. A collection of microstructural photographs was also
published, showing the microstructure of metal using long-term crept specimens obtained from this data
sheet project [12].
1.2.7 Others
Academic societies have been producing data series limited to specific fields only. Those concerned with
welded joints [13] and chemical equipment materials [14] have already been published. Creep data on
products accumulated by industries have also been published as data series [15, 16].
1.3 Testing procedures for obtaining creep data
To obtain creep test data, a large-scale facility needs to be built. The tests are also time-consuming,
making it impossible for one research organization alone to acquire all the required data. It is therefore
important to conduct tests based on a shared method so that highly reliable test data can be obtained,
exchanged and compared. To meet this need, different countries introduced standards for creep testing
and creep rupture testing, and later their respective test standards were incorporated into ISO standards.
Because of this background, the standards currently in use around the world are ISO [17], ASTM [18] of
the United States and in the last years also the European standards EN [19]. The current ISO 204-1997
will be revised as a result of voting at ISO TC164-SC1. EN 10291:2000 [20] appears to have become the
foundation for this amendment. In the revised plan, interrupted tests are allowed and specifications for
temperature tolerance and accuracy in measurement of the cross-sectional area are modified.
Landolt-Börnstein
New Series VIII/2B
4
1 Introduction
[Ref. p. 6
NIMS creep data have been produced not only in conformity with JIS but also with ISO and ASTM. With
regard to temperature and load accuracies in particular, they are found to be better than the relevant
standards, allowing the acquisition of data by performing high-accuracy creep tests [21]. During
experiments that continue for more than 100,000 hours, there is a risk of exceeding the specified
temperature range. In the NIMS creep data sheet, information on this problem is given for each data
point. Temperature measurement during creep testing is generally made using a PR thermocouple, but PR
thermocouples deteriorate during testing and the thermo-electromotive force declines. These results are
published in the paper [22]. This type of information will be helpful in evaluating data and using it
effectively.
European creep data were produced according to several standards in the contributing Nations (DIN
50118, BS 3500, UNI 5111, etc.). A widespread overview on all European standards as well as on relevant
laboratory intern testing practices formed the basement of the data generation recommendations stated in
[19], which then further developed into EN 10291. European data collated in the ECCC programmes
according to a specified recording scheme taking account of all testing details (see [19], Volume 4), were
assessed for conformity with the minimum testing requirements as stated in Volume 3 of [19] before
introduced in general assessment. New data, generated by the ECCC joint programmes, are mandatorily
produced with testing procedures conformed to at least EN10291 or to the "high quality" testing
recommendations in [19].
1.4 Evaluation and assessment of creep data
The determination of creep strength is a process which requires a high accuracy in the application of
the pure testing technique as well as in the handling of the whole process, starting with the specimen
manufacture and including the precision in the measurement of length changes during creep as well as the
final strength computation. Therefore stringent procedures and specifications are necessary to both
guarantee reliable test results and design relevant long-term creep and creep rupture strengths.
The long term strength and creep behaviour of a material is dependent on several factors such as:
• chemical composition,
• way of manufacturing and heat treatment, and
• component size and specimen location.
The reliability of long-term properties of a particular material is fundamentally dependent on the
material pedigree and raw test data verification, on the creep strength assessment method, its application
procedure and on a critical evaluation of their credibility [23, 24, 25, 26].
The data assessment needs to rely on a sufficiently large data base, which must include a
representative number of different casts of the same material grade and – possibly for a big amount of the
casts – on several testing results with long durations in the intended application temperature range.
Testing times should be sufficiently long to avoid extremely high extrapolations in time, and therefore the
stress levels chosen for the specimens have to be well balanced to fulfil the statistical requirements.
In Europe and Japan, the extrapolation rules allow time forecasts of three times the maximum testing
time as this was originally suggested by the meanwhile withdrawn ISO 6303, e.g. for an extrapolation of
a 100,000h (11.4 years) creep rupture strength, minimum data duration of 30,000h (3.4 years) is required
in principle.
If different casts of the same material grade are merged during the assessment, the evaluation with
only statistical or mathematical tools does often not mirror the real material behaviour. The single cast
trends and behaviour have also to be considered, and the assessment of the whole data population and of
its sometimes huge scatter band needs to take this single cast information into account.
Also the pedigree information for the single materials and of each cast of the same material grade
needs to be carefully evaluated during the assessment in order to understand particular cast behaviour and
to avoid erroneous conclusions based on numerous results belonging to casts with extraordinary
surrounding properties.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 6]
1 Introduction
5
Europe developed in the last 10 years a quite rational and objective approach to creep data assessment,
which bases on some fundamental statements:
• Creep strength is considered reliable only if the available experimental data base conforms to
given rules. They require a minimum amount of tested casts, a minimum number of tests per
significant casts and a minimum number of tests with design relevant test durations in the range
of temperatures and stresses expected to be technically relevant.
• To assess creep strength, different methods for data evaluation are required to be applied
contemporaneously to the same multi-cast, big sized and desirably long term test containing data
base in order to ensure the true material behaviour from at least two distinct views.
• A credibility check of the assessment results which were derived from test data is required
before they are allowed to become strength values in order to ensure that failures and damages
do not occur during the design life of the component, taking into account the scatter of properties
in technical applications. This credibility check is codified in the ECCC Post Assessment Tests
(PAT), which include three categories of physical, numerical and statistical tests, determining
the degree of confidence in physical realism, test data description and extrapolation stability of
the computed mathematical expression applying for becoming a creep strength prediction tool.
Actually the majority of the creep strength data proposed for the new EN standards bases on this
approach.
In Japan, a Manual of Extrapolation Methods for Creep-rupture Strength Based on ISO 6303, in which
Larson-Miller, Manson-Haferd and Manson-Brown parameter methods have been introduced as
computer-aided extrapolation methods, was published on 1983 [26]. However, it is pointed out that longterm creep strength which is predicted using these current extrapolation methods is critically
overestimated for advanced ferritic heat resistant steels. In order to improve long-term life prediction for
9-12Cr ferritic creep resistant steels, a new creep life prediction method is proposed in conjunction with a
region partitioning method of stress vs. time to rupture diagrams [27].
The present book includes raw test data in the majority of the presentations. In some cases also
computed strength values are included. The latter are determined by the mentioned rules.
Additional statements on minimum data information requirements, testing techniques, minimum
acceptability criteria and sound testing rules, data assessment procedures and post assessment tests are
available in [23, 24, 28, 29].
1.5 Application of creep data
Design criteria under temperature conditions where creep properties have to be taken account of are
determined by data on creep rupture strength, creep deformation rate, creep strain, etc. In ASME Sec.
VIII-Div. 1, for example, allowable stress may be calculated from the minimum values obtained from the
following:
(1) 67% of the average value of 100,000-hour creep rupture strength
(2) 80% of the minimum value of 100,000-hour creep rupture strength
(3) 100% of the average value of stress that produces a creep rate of 0.01% per 1,000 hours
In Europe and in Turbine and Power Plant industry generally, the recent trend has been to obtain
allowable stress from 200,000-hour creep rupture strength.
If a new material is used, creep testing is conducted to obtain its creep strength, from which an
allowable stress can be defined. However, long-term creep strength is extrapolated using various methods
for life prediction, since it is impossible to obtain data by conducting long-term creep testing under every
condition. However, the prediction of creep strength is not simple and extrapolation from short-term creep
data is not always reliable. The microstructure of metal changes during creep, and affects creep
deformation and rupture life. Many life prediction methods have been proposed, but due to this problem
Landolt-Börnstein
New Series VIII/2B
6
1 Introduction
none of them has been ideal. And even if so called Post Assessment Tests have been derived to establish a
credibility criterion for data description and extrapolation functions, no mathematical relationship
overcomes the need to obtain long-term creep data.
Meanwhile, an additional challenge has appeared due to an increasing number of high-temperature
plants which are still in use despite exceeding their design life, which cannot be replaced in short times or
shut down. It has become an important task to some National Power Production Balances to keep these
plants running by carefully evaluating the remaining life of these plants with a long history of use. In
making this evaluation, an increased volume of abundant, more reliable and longer-term test data and
microstructural information than those available at the time of plant design is essential. This data should
include the changes in the microstructure of the metal during creep, the formation of creep damage and its
growth, creep deformation, creep cracking behavior resulting from defects, the strength of welded joints,
the effectiveness of multi-axiality on strength and failure, and creep rupture strength.
1.6 References
[1] W. Cross: The Code, An authorized history of the ASME boiler and pressure vessel code, ASME,
(1990), p.93
[2] H. Jungblut: Sonderstähle für den Dampfkesselbau. Mitt. VGB (1930), H.28, pp.141-146
[3] ASTM Data Series, example
DS-11-S1 (1970); Carbon steel
DS-47 (1971); Mo steel, Mn-Mo steel, Mn-Mo-Ni steel
DS-50 (1970); 0.5Cr-0.5Mo steel, 1Cr-0.5Mo steel, 1.25Cr-0.5Mo-Si steel
DS-6-S1 (1971); 2.25Cr-1Mo steel
DS-5 8(1975); 3 to 9Cr steels
DS-18 (1958); 12 to 27Cr steels
DS-5-S1 (1965); Stainless steels
DS-7-S1 (1968); Superalloys
DS-20 (1960); Al alloys, Mg alloys
[4] BSCC High Temperature Data, The Iron and Steel Institute, (1972) stahl-Eisen
[5] Ergebnisse deutscher Zeitstandversuche langer Dauer, Verlag Stahl mbH, (1969)
[6] Ergebnisse deutsche Zeitstandversuche langer Dauer an Stahlgusssorten nach DIN 17 245
- warmfester ferritischer Stahlguss -, Bericht FVW/FVV Nr.1-86, (1986)
[7] Ergebnisse deutscher Zeitstandversuche langer Dauer an den hochwarmfesten Legierungen X 40
CoCrNi 20 20 (Typ S-590) und X 12 CrCoNi 21 20 (Typ N-155), Bericht FVHT/FVV Nr.2-87,
(1987)
[8] ECCC Recommendations 2001 “Creep Data Validation and Assessment Procedures”, Publ. ERA
Technology Ltd., Leatherhead, UK, (2001)
[9] ECCC Data Sheets, Publ. ERA Technology Ltd., Leatherhead, UK, (1999)
[10] The Iron and Steel Institute of Japan: Report on the Mechanical Properties of Metals at Elevated
Temperatures
Vol. 1, Low ally steels (1972)
Vol. 2, Stainless steels (1975)
Vol. 3, Carbon steels and cast irons (1977)
Vol. 4, Superalloys (1979)
Vol. 5, Deposited metal, weld metal and welded joint (1985)
[11] NIMS (former NRIM) Creep Data Sheets, No.0 to No.48, National Institute for Materials Science
[12] National Research Institute for Metals: NRIM Creep Data Sheet, Metallographic Atlas of Long-term
Crept Materials, No.M-1, (1999)
[13] High Pressure Institute of Japan: High Temperature Strength Data Book of Welded Joint, (1967)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 6]
1 Introduction
7
[14] The Japan Petroleum Institute: High temperature creep rupture strength data of heat-resistant alloy
and heat-resistant cast steel for oil refining and petrochemical equipment, (1979)
[15] ESCHER WYSS: Zeitstandversuche an Staehlen, (1972)
[16] Sumitomo Metal Industries, Ltd.: Creep Data Sheets, Sumitomo Seamless Tubes and Pipe, (1993)
[17] ISO 204-1997, Metallic materials-Uninterrupted uniaxial creep testing in tension-Method of test
[18] ASTM E 139-00, Standard Test Methods for Conducting Creep, Creep-Rupture, and Stress-Rupture
Tests of Metallic Materials, Annual Book of ASTM Standards, Vol.03.01, (2001), pp.270-281
[19] ECCC Recommendations 2001 “Creep Data Validation and Assessment Procedures”, Publ. ERA
Technology Ltd., Leatherhead, UK,, (2001)
[20] EN 10291:2000, Metallic materials-Uniaxial creep testing in tension-Methods of test
[21] National Research Institute for Metals: NRIM Materials Strength Data Sheet Technical Document,
No.10, “Testing Plan and Testing Procedures or NRIM Creep Data Sheet Project’’, (1996)
[22] H. Itoh, M. Egashira, H. Miyazaki, Y. Monma and S. Yokoi: Tetsu-to-Hagane, 72 (1986), 1944
[23] ECCC Recommendations 2001, Volume 3, ‘Recommendations for data acceptability criteria and the
generation of creep, creep rupture, stress rupture and stress relaxation data’, Eds. Granacher
J.,Holdsworth S.R., Klenk. A., Buchmayr B. & Gariboldi E., Publ. ERA Technology Ltd.,
Leatherhead, UK, (a) Part I: Generic recommendations for creep, creep rupture, stress rupture and
stress relaxation data, (b) Part II: Creep data for welds, (c) Part III: Creep testing of PE- (ex service)
materials.
[24] ECCC Recommendations, 2001, Volume 5 ‘Guidance for the assessment of creep rupture, creep
strain and stress relaxation data’, Eds. Holdsworth S.R. & Merckling G., Publ. ERA Technology Ltd,
Leatherhead, UK, (a) Part I: Full-size datasets, (b) Part IIa: Sub-size datasets, (c) Part IIb: Weldment
datasets, (d) Part III: Datasets for PE (ex-service) materials
[25] Yokoi, S., Monma, Y.: Prediction of Long-time Creep-rupture Strength for High-temperature
Materials; Tetsu-to-Hagane 65, No.7 (1979) 831
[26] Fujita, T., Monma, Y.: Accuracy of Extrapolation for Creep-rupture Strength and Standardization of
Extrapolation Methods; Tetsu-to-Hagane, 70, No.3 (1984) 327
[27] Kimura, K., Kushima, H., Abe, F.: Improvement of Creep Life Prediction of High Cr Ferritic Creep
Resistant Steels by Region Partitioning Method of Stress vs. Time to Rupture Diagram; J. Soc. Mat.
Sci., Japan, 52, No.1 (2003) 57
[28] Holdsworth, S.R., Orr, J., Granacher, J., Merckling, G., Bullogh, C.K. on behalf of the ECCC-WG1
“Creep Data Collation and Assessment”: European Creep Collaborative Committee Activities on
Creep Data Generation and Assessment Methodologies, in : D. Coutsouradis et. al. (editors):
“Materials for Advanced Power Engineering”, 1994, Liege, 3.- 6.10.1994, Kluwer Accademic
Publishers, p. 591 - 600
[29] National Research Institute for Metals : “Testing Plan and Testing Procedures for NRIM Creep Data
Sheets Project”, NRIM Material Strength Data Sheet, Technical Document, No.10, (1996)
Landolt-Börnstein
New Series VIII/2B
2 Creep and rupture data of heat resistant steels - 2.1 Carbon steels
9
2 Creep and rupture data of heat resistant steels
2.1 Carbon steels
2.1.1 0.1C steel
2.1.1.1 Introduction
This carbon steel for boiler and heat exchanger tubes is used as water tube, smoke tube, super-heater tube,
air-preheater tube, etc. in boiler and as heat exchanger tube, condenser tube, catalyst tube, etc. in chemical
and petrolic industries. The carbon steels are used only at temperatures lower than 400 °C, because they
have not enough creep strength for higher temperatures.
2.1.1.2 Material standards, chemical and tensile requirements
Table 1. Chemical requirements of 0.1C steel tubes; JIS STB340, ASTM A, BS360 and DIN St35.8
Standards
Designation
JIS
ASTM
BS
DIN
STB340
A
360
St35.8
C
≤0.18
0.06-0.18
≤0.17
≤0.17
Chemical composition [wt%]
Si
Mn
P
0.30-0.60 ≤0.035
≤0.35
0.27-0.63 ≤0.035
0.10-0.35 0.40-0.80 ≤0.035
0.10-0.35 0.40-0.80 ≤0.040
S
≤0.035
≤0.035
≤0.035
≤0.040
Table 2. Tensile properties of 0.1C steel tubes at room temperature; JIS STB340
Tensile strength Yield strength
Elongation
[N/mm2]
[N/mm2]
[%]
d ≥20 mm
20>d ≥10 mm d<10 mm
≥340
≥175
≥35
≥30
≥27
Landolt-Börnstein
New Series VIII/2B
Std.No.
G3461
A178
3059-2
17175
10
2.1 Carbon steels
2.1.1.3 Creep properties of 0.1C steel tubes
Information of fact on creep data for 0.1C steels can be obtained from [1], [2] and [3].
STB340 (0.12C steel)
500
400
300
200
100
80
400℃
60
50
40
30
450℃
500℃
550℃
100
101
Fig. 1. Creep rupture strength
data of STB340; [1].
102
103
104
105
Time to rupture (h)
2.1.1.4 References
[1] The Iron and Steel Institute of Japan: Report on the Mechanical Properties of Metals at Elevated
Temperatures, Vol. III Carbon Steels and Cast Irons, (1977), 195-232.
[2] American Society for Testing Materials: Elevated-Temperature Properties of Carbon Steels, ASTM
Special Technical Publication No. 180, (1955), 11-23.
[3] The Iron and Steel Institute: BSCC High Temperature Data, (1972), 1-254.
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New Series VIII/2B
Ref. p. 19]
2.1.2 0.2C-0.3C steel
11
2.1.2 0.2-0.3C steel
2.1.2.1 Introduction
0.2-0.3C steels are used as tubes for heat exchangers, boilers, superheaters and feedwater heaters in power
plants, chemical and petrochemical plants. 0.2-0.3C steel plates are used for boilers and pressure vessels
in power plants, chemical and petrochemical plants.
Creep strength of the 0.2-0.3C steel is strongly influenced by small amounts of molybdenum through
the strengthening effects of Mo-C and Mo-N atomic pairs in solid solution, as will be explained later.
2.1.2.2 Material standards, chemical and tensile requirements
2.1.2.2.1 0.2-0.3C steel tubes for heat exchangers
Table 3. Chemical requirements of 0.2-0.3C steel tubes; JIS STB410, Japanese METI KA STB480,
ASTM Gr. C, ASTM Gr. A-1 and ASTM Gr. C2
Chemical composition [wt%]
Standards Designation
Std. No
C
Si
Mn
P
S
JIS
STB410
0.30-0.80
G3461
≤0.32
≤0.35
≤0.035
≤0.035
Japanese
KA STB480 ≤0.30
0.29-1.06
≥0.10
≤0.048
≤0.058
METI
ASTM
Gr. C
A178
≤0.35
≤0.80
≤0.035
≤0.035
ASTM
Gr. A-1
A210
≤0.27
≥0.10
≤0.93
≤0.035
≤0.035
ASTM
Gr. C
0.29-1.06
A210
≤0.35
≥0.10
≤0.035
≤0.035
ASTM
Gr. C2
0.29-1.06
A556
≤0.30
≥0.10
≤0.035
≤0.035
2.1.2.3 0.2 - 0.3C steel plates for boilers and pressure vessels
Table 4. Chemical requirements of 0.2-0.3C steel plates; JIS SB410, JIS SB480, JIS SGV410
Chemical composition [wt%]
Standards Designation Thickness
C
Si
Mn
P
S
Mo
[mm]
≤25
≤0.24
JIS
SB410
25 - 50
≤0.035 ≤0.040 ≤0.27 0.15 - 0.30 ≤0.90
50 - 200
≤0.30
≤25
≤0.31
JIS
SB480
25 - 50
≤0.035 ≤0.040 ≤0.33 0.15 - 0.30 ≤0.90
50 - 200
≤0.35
≤12.5
≤0.21
12.5 - 50
≤0.23
JIS
SGV410
0.15 - 0.40 0.85 - 1.20 ≤0.030 ≤0.030 50 - 100
≤0.25
100 - 200 ≤0.27
Landolt-Börnstein
New Series VIII/2B
Std. No
G3103
G3103
G3118
12
2.1 Carbon steels
Table 5. Chemical requirements of 0.2-0.3C steel plates; ASTM Gr. B, ASTM Gr. 60 and ASTM Gr. 70
Chemical composition [wt%]
DesigStd. No
Standards
Thickness
nation
C
Si
Mn
P
S
Mo
[mm]
≤25
≤0.20
25 - 50
0.45 ≤0.23
A204
0.15 - 0.40 ≤0.90
ASTM
Gr. B
≤0.035 ≤0.035
0.60
50 - 100
≤0.25
>100
≤0.27
≤25
≤0.24
25 - 50
≤0.27
ASTM
Gr. 60
A515
50 - 100
≤0.035 ≤0.035 ≤0.29 0.15 - 0.40 ≤0.90
100 - 200
≤0.31
>200
≤0.31
≤25
≤0.31
25 - 50
≤0.33
ASTM
Gr. 70
A515
50 - 100
≤0.035 ≤0.035 ≤0.35 0.15 - 0.40 ≤1.20
100 - 200
≤0.35
>200
≤0.35
0.60 - 0.90
≤12.5
≤0.21
12.5 - 50
≤0.23
50 - 100
A516
ASTM
Gr. 60
≤0.035 ≤0.035 ≤0.25 0.15 - 0.40
0.85 - 1.20
100-200
≤0.27
>200
≤0.27
≤12.5
≤0.27
12.5-50
≤0.28
50-100
A516
ASTM
Gr. 70
≤0.30 0.15 - 0.40 0.85 - 1.20 ≤0.035 ≤0.035 100-200
≤0.31
>200
≤0.31
2.1.2.3 Creep properties of 0.2-0.3C steel tubes
Information of fact on creep data for 0.2-0.3C steel tubes can be obtained from [1].
2.1.2.3.1 Creep rupture data of 0.2-0.3C steel tubes
The results of creep tests for 9 heats of JIS STB410 steel tubes are compiled in [1]. From this data sheet
the data of rupture elongation, reduction of area and microstructures of as-received materials and crept
specimens can be also obtained.
Creep rupture strength data for 9 heats of 0.2C steel tubes (JIS STB410) is shown in Fig. 2 [1]. Very
large heat-to-heat variation of creep rupture strength is observed over the whole range of creep test
conditions from short-term to long-term. Differences in creep rupture strength are caused by differences
in small amounts of molybdenum [2, 3]. Creep strength of the 0.2C steel is strongly influenced by small
amounts of molybdenum through the strengthening effects of Mo-C and Mo-N atomic pairs in solid
solution [4].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 19]
2.1.2 0.2C-0.3C steel
13
500
Stress (MPa)
300
o
100
○ 400 C
80
△ 450 C
o
o
60
□ 500 C
n = 207
40 1
10
10
2
3
10
4
10
5
10
10
6
Fig. 2. Creep rupture strength
data of 0.2C steel tubes (JIS
STB410) according to [1]. n
indicates the total number of data
points.
Time to rupture (h)
2.1.2.3.2 Creep rupture strength of 0.2-0.3C steel tubes
Creep rupture strength was analyzed applying the Larson-Miller parameter method to NRIM creep
rupture data on 0.2C steel tubes (JIS STB 410). The result is shown in Fig. 3. Sigmoidal inflection with a
large scatter band is observed.
600
500
400 °C
450 °C
500 °C
400
Stress [MPa]
300
200
100
80
60
50
40
10
Average
n = 207
Fig. 3. Master rupture curve obtained by Larson-Miller
parameter method for 0.2C steel tubes (JIS STB 410); [1].
n indicates the total number of data points.
12
14
16
18
Larson-Miller-parameter TK [( log tR +15.753) [103 ]
2.1.2.3.3 Microstructural changes
The typical initial microstructure of 0.2C steel tubes consists of ferritic and pearlitic grains. Optical
micrographs of an as-received 0.2C steel tube are shown in Fig. 4. The bright grains are ferritic and the
dark grains are pearlitic.
Optical micrographs of 0.2C steel tube specimens creep ruptured after 138,403.7 h at 450 °C and 78
MPa are shown in Fig. 5. Coarsening of carbides within pearlitic grains is observed after long-term creep
exposure at 450 °C. Changes in morphology and distribution of carbides within pearlitic grains are used
as indicator of degradation of 0.2C steel due to long-term service at elevated temperatures.
Landolt-Börnstein
New Series VIII/2B
14
2.1 Carbon steels
Fig. 4. Optical micrographs of as-received 0.2C steel tubes (etched in 4% nital); [1].
Fig. 5. Optical micrographs of 0.2C steel tube specimens (etched in picral) creep ruptured after 138,403.7 h at
450 °C and 78 MPa; [1].
2.1.2.3.4 Creep deformation behavior of 0.2C steel tubes
The creep deformation behavior of 0.2C steel tubes strongly depends on slight differences in chemical
composition, heat treatment and initial microstructure. Creep rate vs. time curves of 0.2C steel tubes at
550 °C and 69 MPa are shown in Fig. 6 [5]. Heat-to-heat variation of creep deformation behavior is
clearly observed on these 3 heats of 0.2C steel tubes.
Creep rate vs. time curves of as-received and pre-aged 0.2C steel tubes at 550 °C and 69 MPa are
shown in Fig. 7 [6]. Since creep deformation is strongly influenced by microstructural changes during
creep exposure, complex creep deformation behavior observed for un-aged steel disappeares by preageing.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 19]
2.1.2 0.2C-0.3C steel
15
-2
-2
10
10
550oC-69MPa
550oC-69MPa
-3
10
CAM
-4
10
Creep rate (h-1)
Creep rate (h-1)
-3
CAH
CAC
-5
10
-4
10
○ un-aged
● 100h aged
△ 200h aged
▲ 300h aged
□ 500h aged
■ 1,000h aged
-5
10
-6
-6
10
10
-1
10
10
0
1
10
10
2
10
3
4
10
10
-1
10
Time (h)
Fig. 6. Creep rate vs. time curves of 0.2C steel tubes
at 550 °C and 69 MPa; [5].
10
0
1
10
10
2
3
10
10
4
Time (h)
Fig. 7. Effect of pre-ageing on creep deformation behavior of 0.2C steel tube; [6].
2.1.2.3.5 Effect of molybdenum on creep rupture strength
The creep deformation behavior of 0.2C steel tubes is strongly influenced by microstructural changes
during creep exposure at elevated temperatures, as mentioned above. Creep strength decreases as a result
of microsstructural changes and it becomes an inherent creep strength, which is the creep strength of the
ferrite matrix itself, after long-term creep exposure [7, 8].
The inherent creep strength of 0.2C steel tubes is extremely improved by small amounts of
molybdenum in solid solution [2, 3]. The very large heat-to-heat variation of long-term creep rupture
strength for 0.2C steel tubes, as shown in Figs. 2 and 3, is caused by differences in the inherent creep
strength due to a wide variety of molybdenum concentrations, even at low Mo levels of less than
0.02 mass%. Inherent creep strength of 0.2C steel tubes is increased by strengthening effects of Mo-C and
Mn-C atomic pairs in solid solution [4].
Inherent creep strength is improved by small amounts of molybdenum, however, this effect is
saturated at about 0.03 mass% of molybdenum [2, 3]. Therefore, the inherent creep strength obtained by
addition of 0.03 mass% of molybdenum is the highest for 0.2C steel. It has been experimentally found
that the inherent creep strength of ferritic creep resistant steels is almost the same independent of
chemical composition, heat treatment condition and short-term creep strength [7, 8]. There is a good
correspondence between common inherent creep strength for ferritic creep resistant steels and the highest
inherent creep strength for 0.2C steel with addition of 0.03 mass% of molybdenum [2, 3].
2.1.2.3.6 Estimated long-term creep strength
The temperature dependence of 0.2% proof stress, tensile strength and creep rupture strength at 1,000 and
100,000 h for 9 heats of 0.2C steel tubes is shown in Fig. 8. The creep rupture strength curves shown in
Fig. 8 were obtained by regression analysis using the Larson-Miller parameter.
Landolt-Börnstein
New Series VIII/2B
16
2.1 Carbon steels
800
600
500
400
Stress [MPa]
300
1000 h
{
Tensile
strength
200
Fig. 8. Temperature depen-dence
of 0.2% proof stress, tensile
strength and creep rupture strength
at 1,000 and 100,000 h for 0.2C
steel tubes (JIS STB 410); [1]. The
dashed lines are the upper and
lower 95 % confidence limit (±2σ,
σ: standard deviation).
0.2% proof
stress
100
80
60
50
40
350
100000 h
400
450
500
Temperature [°C]
550
600
2.1.2.4 Creep properties of 0.2-0.3C steel plates
Information of fact on creep data for 0.2-0.3C steel plates can be obtained from [9].
2.1.2.4.1 Creep rupture data of 0.2-0.3C steel plates
The results of creep tests for 8 heats of JIS SB480 steel plates are compiled [9]. From this data sheet data
of rupture elongation, reduction of area, minimum creep rate, time to specified strain and microstructures
of as-received materials and crept specimens can be also obtained.
Creep rupture strength data for 8 heats of 0.3C steel plates (JIS SB480) is shown in Fig. 9. Very large
heat-to-heat variation of creep rupture strength is observed, especially in the lower temperature and higher
stress condition. With increase in temperature and decrease in applied stress, heat-to-heat variation of
creep rupture strength decreases.
500
Stress (MPa)
400
300
200
o
○ 400 C
o
△ 450 C
100
o
90 □ 500 C
80 n=117
70
60 0
1
10
10
Fig. 9. Creep rupture strength
data of 0.3C steel plates (JIS
SB480); [9]. n indicates the total
number of data points.
2
10
3
10
4
10
10
5
10
6
Time to ruputre (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 19]
2.1.2 0.2C-0.3C steel
17
2.1.2.4.2 Creep rupture strength of 0.2-0.3C steel plates
Creep rupture strength was analyzed applying the Orr-Sherby-Dorn parameter method to NRIM creep
rupture data on 0.3C steel plates (JIS SB 480). The result is shown in Fig. 10. A very large scatter band of
creep rupture strength is observed, especially in the higher stress condition.
1000
800
Stress [MPa]
600
500
400
300
400 °C
425 °C
450 °C
475 °C
500 °C
525 °C
550 °C
575 °C
200
100
80
Average
n = 131
60
50
-19
-17
-13
-11
-9
-15
Orr-Sherby-Dorn parameter log tR -[247364/(19.1425 × TK )]
Fig. 10. Master rupture curve by Orr-Sherby-Dorn
parameter method for 0.3C steel plates (JIS SB 480); [9].
n indicates the total number of data points.
2.1.2.4.3 Microstructural changes
The typical initial microstructure of 0.3C steel plates consists of ferritic and pearlitic grains. Optical
micrographs of as-received 0.3C steel plates are shown in Fig. 11. The bright grains are ferritic and the
dark ones are pearlitic.
Optical micrographs of 0.3C steel plate specimens creep ruptured at 400, 450 and 500 °C are shown in
Fig. 12, 13 and 14, respectively. Coarsening of carbides within pearlitic grains is observed after long-term
creep exposure. By comparing the microstructures shown in Fig. 12, 13 and 14, carbide coarsening is
more significantly observed in the specimens crept at higher testing temperatures.
Fig. 11. Optical micrographs of as-received 0.3C steel plate (etched in 2 % nital); [9].
Landolt-Börnstein
New Series VIII/2B
18
2.1 Carbon steels
Fig. 12. Optical micrographs of a 0.3C steel plate specimen creep ruptured after 155,727.0 h at 400 °C and 333 MPa
(etched in 2 % nital); [9].
Fig. 13. Optical micrographs of a 0.3C steel plate specimen creep ruptured after 122,103.2 h at 450 °C and 196 MPa
(etched in 2 % nital); [9].
Fig. 14. Optical micrographs of a 0.3C steel plate specimen creep ruptured after 85,699.2 h at 500 °C and 88 MPa
(etched in 2 % nital); [9].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 19]
2.1.2 0.2C-0.3C steel
19
2.1.2.4.4 Estimated long-term creep strength
The temperature dependence of 0.2% proof stress, tensile strength and creep rupture strength at 1,000 and
100,000 h for 8 heats of 0.3C steel plates is shown in Fig. 15 [9]. Creep rupture strength curves shown in
Fig. 15 were obtained by regression analysis using the Orr-Sherby-Dorn parameter.
1000
800
60
50
350
400
1000 h
{
100
80
{
200
100000 h
450
500
Temperature [°C]
Tensile
strength
{
300
{
Stress [MPa]
600
500
400
0.2% proof
stress
550
600
Fig. 15. Temperature dependence
of 0.2% proof stress, tensile
strength and creep rupture strength
at 1,000 and 100,000 h for 0.3C
steel plates (JIS SB 480); [9].
2.1.2.5 References
[1] NRIM Creep Data Sheet, No. 7B, (1992).
[2] Kimura, K., Kushima, H., Yagi, K., and Tanaka, C.: Proc. of JIMIS-7 on Aspects of High
Temperature Deformation and Fracture in Crystalline Materials, Hosoi, Y., et al., eds., Nagoya,
Japan, July 1993, The Japan Inst. Metals, (1993), 309-316.
[3] Kimura, K., Kushima, H., Yagi, K., and Tanaka, C.: Tetsu-to-Hagane, 81, (1995), 757-762.
[4] Onodera, H., Abe, T., Ohnuma, M., Kimura, K., Fujita, M., and Tanaka, C.: Tetsu-to-Hagane, 81,
(1995), 821-826.
[5] Kimura, K., Kushima, H., and Yagi, K.: Proc. of 10th Int. Conf. on the Strength of Materials,
Oikawa, H., et al., eds., Sendai, Japan, August 1994, The Japan Inst. Metals, (1994), 645-648.
[6] Kimura, K., Kushima, H., Abe, F., and Yagi, K.: Tetsu-to-Hagane, 82, (1996), 713-718.
[7] Kimura, K., Kushima, H., Yagi, K., and Tanaka, C.: Tetsu-to-Hagane, 77, (1991), 667-674.
[8] Kimura, K., Kushima, H., Yagi, K., and Tanaka, C.: Proc. of Inter. Conf. on Creep and Fracture of
Engineering Materials and Structures, Wilshire, B., and Evans, R.W., eds., Swansea, UK, The
Institute of Materials, 5, (1993), 555-564.
[9] NRIM Creep Data Sheet, No. 17B, (1994).
Landolt-Börnstein
New Series VIII/2B
20
2.1 Carbon steels
2.1.3 C-Mn steel
2.1.3.1 Introduction
C-Mn steels are used as tubes for boilers and heat exchangers in power plants, chemical and
petrochemical plants. C-Mn steels shall be killed. C-Mn steel tubes are heat treated at a temperature of
900 °C or higher and followed by cooling in air.
2.1.3.2 Materials standards, and chemical and tensile requirements
2.1.3.2.1 C-Mn steel tubes for heat exchangers
Table 6. Chemical requirements of C-Mn steel tubes; JIS STB 510, ASTM Gr. D.
Chemical composition [wt%]
Standards Designation
C
Si
Mn
P
S
JIS
STB510
1.00~1.50
≤0.25
≥0.35
≤0.035
≤0.035
ASTM
Gr. D
1.00~1.50
≤0.27
≥0.10
≤0.030
≤0.015
Std. No
G3461
A178
2.1.3.3 Creep properties of C-Mn steel tubes
Information on creep data for C-Mn steel tubes can be obtained from [1].
2.1.3.3.1 Creep rupture data of C-Mn steel plates
The results of creep tests for 2 heats of JIS STB510 steel tubes are compiled in [1]. From this data sheet
the data of 0.2% proof stress, tensile strength, rupture elongation, reduction of area and microstructures of
as-received materials and crept specimens can be also obtained.
Creep rupture strength data of 2 heats of the 0.2C-1.3Mn silicon killed steel tubes (JIS STB510) is
shown in Fig. 16 [1]. The slope of the stress vs. time to rupture curve at 400 °C increases with decrease in
applied stress. The creep rupture curve at 450 °C indicates a slight inflection of sigmoidal shape. On the
other hand, good linear relationship between stress and time to rupture is observed at 500 °C.
700
Stress (MPa)
500
300
○ 400oC
△ 450oC
□ 500oC
100
80
60 0
10
n = 32
10
1
2
10
10
3
4
10
10
5
Fig. 16. Creep rupture strength
data of 0.2C-1.3Mn silicon
killed steel tubes (JIS STB510);
[1]. n indicates the total number
of data points.
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 22]
2.1.3 C-Mn steel
21
2.1.3.3.2 Microstructural change
Initial microstructure of 0.2C-1.3Mn silicon killed steel tubes consists of ferritic and pearlitic grains.
Optical micrographs in the as-received condition of 0.2C-1.3Mn silicon killed steel tubes are shown in
Fig. 17. The bright grains are ferritic and the dark ones are pearlitic. Optical micrographs of the crept steel
tubes are shown in Fig. 18. Pearlitic microstructure is broken due to carbide coarsening during creep
exposure at elevated temperatures.
Fig. 17. Optical micrographs of as-received 0.2C-1.3Mn silicon-killed steel tubes (etched in a solution of ethyl
alcohol with 2 % picric acid ); [1].
Fig. 18. Optical micrographs of 0.2C-1.3Mn silicon-killed steel tubes creep ruptured after 30,530.3 h at 500 °C and
70 MPa (etched in a solution of ethyl alcohol with 2 % picric acid); [1].
2.1.3.3.3 Estimated long-term creep strength
The temperature dependence of 0.2% proof stress, tensile strength and creep rupture strength at 100 and
10,000 h for 9 heats of 0.2C-1.3Mn steel tubes is shown in Fig. 19 [1]. That of 0.2% proof stress, tensile
strength and creep rupture strength at 1,000 and 100,000 h for the same materials is shown in Fig. 20 [1].
Creep rupture strength curves shown in Fig. 19 and Fig. 20 were obtained by regression analysis using the
Orr-Sherby-Dorn parameter.
Landolt-Börnstein
New Series VIII/2B
22
2.1 Carbon steels
1000
800
Stress [MPa]
100 h
200
100
80
60
50
40
350
Tensile
strength
{
300
{
600
500
400
0.2% proof
stress
10000 h
400
450
500
Temperature [°C]
550
600
Fig. 19. Temperature dependence
of 0.2% proof stress, tensile
strength and creep rupture strength
at 100 and 10,000 h for 0.2C-1.3Mn
silicon killed steel tubes; [1]. The
dashed lines are the upper and
lower 95 % confidence limit (±2σ,
σ: standard deviation).
1000
800
300
{
Tensile
strength
200
{
Stress [MPa]
600
500
400
0.2% proof
stress
1000 h
100
80
60
50
40
350
30000 h
400
450
500
Temperature [°C]
550
600
Fig. 20. Temperature dependence
of 0.2% proof stress, tensile
strength and creep rupture strength
at 1,000 and 100,000 h for 0.2C1.3Mn silicon killed steel tubes;
[1].
2.1.3.4 Reference
[1] NRIM Creep Data Sheet, No.40A, (2000).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 25]
2.1.4 0.25C cast
23
2.1.4 0.25C cast
2.1.4.1 Introduction
The 0.25C cast material is a traditional unalloyed creep resistant cast steel. The grade is specified as
GP240GH in EN 10213-2, material-no 1.0619. Typical features of the 0.25C cast material necessary to
consider are summarized below:
•
•
•
•
•
•
Melting processes: Electric arc, induction melting
Heat treatment: Normalized, or quenched and tempered (cooling in furnace)
Typical microstructure: Ferrite and pearlite
Weldability: Easily weldable with similar weld metal
High temperature applications: Casings of steam turbines, service temperatures up to about 450 °C
Cast steel grade with similar chemical composition: ASTM A216 Grade WCA
2.1.4.2 Standard requirements
Table 7. Chemical composition
Standard
Designation
C
Si
Chemical composition [wt%]
Mn
P
S
EN 0213GP240GH
0.18 - 0.25
≤0.60
≤1.20
≤0.030
≤0.020(1)
2:1995
(1.0619)
(1) The maximum admissible sulphur content is 0.030 % if the relevant wall thickness is not in excess of
28 mm.
Table 8. Heat treatment and tensile properties at room temperature
Min. 0.2 %
Thickness
proof strength
Standard
Designation Heat treatment
[mm]
[MPa]
N:900 °C-980 °C
EN10213GP240GH
240
Q:890 °C-980 °C 100
2:1995
(1.0619)
T:600 °C-700 °C
N: Normalized, Q: Quenched, T: Tempered
Rp0.2
Min. elongation at
rupture [%]
420-600
22
Rm
min EN
600
400
500
300
Rm (MPa)
Rp0,2 (MPa)
Tensile
strength
[MPa]
200
100
400
300
200
100
0
0
0
100
200
300
Temperature (°C)
400
500
0
100
200
300
400
500
Temperature (°C)
Fig. 21. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade GP240GH tested in creep rupture tests
by the German Creep Committee; [1]. min EN: minimum values by EN 10213-2.
Landolt-Börnstein
New Series VIII/2B
24
2.1 Carbon steels
Stress (MPa)
1000
broken
100
unbroken
400°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 22. Creep rupture strength data of cast steel grade GP240GH at 400 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
450°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 23. Creep rupture strength data of cast steel grade GP240GH at 450 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
500°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 24. Creep rupture strength data of cast steel grade GP240GH at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 25]
2.1.4 0.25C cast
25
2.1.4.3 Average creep rupture strength
Table 9. Average creep rupture strength values indicated in EN 10213-2:1995
Average creep rupture strength [MPa]
Temperature
Time to rupture
[°C]
10,000 h
100,000 h
200,000 h
400
205
160
145
450
132
83
71
500
74
40
32
2.1.4.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade GP240GH, compilation of test results; Forschungsvereinigung Warmfeste
Stähle, c. o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
26
2.1 Carbon steels
2.1.5 C-Mn cast
2.1.5.1 Introduction
For steel grade GP280GH (EN 10213-2, material-no 1.0625) C-Mn cast, a manganese content of 0.80 to
1.20 % is specified in EN 10213-2:1995. The manganese content may be increased if the specified
maximum carbon content is reduced by 0.01 % for each 0.04 % Mn in excess of 1.20 % up to a maximum
manganese content of 1.40 %.
The increased manganese content promotes the creep rupture properties. Since 1989 the German
Creep Committee has performed a number of qualification tests on quenched and tempered test materials
of casts with 1.20 to 1.40 % Mn. The results of these tests are reported in this chapter. As expected the
average creep rupture strength values of the casts are much higher than those indicated in EN 102132:1995 for steel grade GP280GH with a standard manganese content of 0.80 to 1.20 %. Typical features
of GP280GH are summarized below:
•
•
•
•
•
Melting process: Electric arc, basic oxygen, argon oxygen decarburization, induction melting
Heat treatment: Normalized, or quenched in air or water and tempered
Typical microstructure: Ferrite and tempered bainite
Weldability: Excellent weldability; weld metal of type E Mo should be used to obtain sufficient creep
rupture strength of weldments
High temperature applications: Casings of compressors (especially with very cold inlet and hot
outlet), gas turbines, outer casings of steam turbines, valves, fittings; service temperatures up to
450 °C
2.1.5.2 Standard requirements
Table 10. Chemical composition
Standard
Designation
C
Si
Chemical composition [wt %]
Mn
P
S
EN 10213- GP280GH
(1)
(1)
0.80 - 1.20
0.18 - 0.25
≤0.60
≤0.030
≤0.020(2)
2:1995
(1.0625)
(1) The maximum admissible manganese content may be exceeded up to 1.40% if the maximum
admissible carbon content is reduced by 0.01% per each 0.04% Mn in excess of 1.20%.
(2) The maximum admissible sulphur content is 0.030% if the relevant wall thickness is not in excess of
28 mm.
Table 11. Heat treatment and tensile properties at room temperature
Min. 0.2 %
Thickness
proof strength
Standard
Designation Heat treatment
[mm]
[MPa]
N:900°C-980°C
EN 10213- GP280GH
280
Q:890°C-980°C 100
2:1995
(1.0625)
T:600°C-700°C
N: Normalized, Q: Quenched, T: Tempered
Tensile
strength
[MPa]
Min. elongation at rupture
[%]
480-640
22
Landolt-Börnstein
New Series VIII/2B
Ref. p. 28]
2.1.5 C-Mn cast
Rp0.2
min EN
Rm
500
800
Rm (MPa)
(MPa)
400
,2
300
Rp
27
200
100
0
600
400
200
0
0
100
200
300
400
500
0
Temperature (°C)
100
200
300
400
500
600
Temperature (°C)
Fig. 25. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade GP280GH tested in creep rupture tests
by the German Creep Committee; [1]. min EN: minimum values by EN10213-2.
Stress (MPa)
1000
broken
unbroken
400°C_EN
100
100
1000
10000
100000
Test duration (h)
Fig. 26. Creep rupture strength data of cast steel grade GP280GH at 400 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
450°C_EN
10
100
1000
10000
100000
Test duration (h)
Fig. 27. Creep rupture strength data of cast steel grade GP280GH at 450 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
28
2.1 Carbon steels
Stress (MPa)
1000
broken
100
unbroken
500°C_EN
10
100
1000
10000
100000
Test duration (h)
Fig. 28. Creep rupture strength data of cast steel grade GP280GH at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
2.1.5.3 Average creep rupture strength
Table 12. Average creep rupture strength values indicated in EN 10213-2:1995
Average creep rupture strength [MPa]
Temperature
Time to rupture
[°C]
10,000 h
100,000 h
200,000 h
400
210
165
450
135
85
500
75
42
2.1.5.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade GP280GH, compilation of test re-sults; Forschungsvereinigung Warmfeste
Stähle, c. o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 34]
2.2.1 0.5Mo steel
29
2.2 Low alloy steels
2.2.1 0.5Mo steel
2.2.1.1 Introduction
0.5Mo steels are applied for heat exchangers and piping systems in thermal power plants and are supplied
for the plate members of pressure vessels. Mo, 0.5% of mass of the steel, increases creep rupture strength
by both solid solution strengthening and carbide precipitation strengthening. Normalizing and tempering
heat treatment processes produce ferrite and pearlite phase mixture in the microstructure of 0.5Mo steels.
2.2.1.2 Material standards, chemical and tensile requirements
2.2.1.2.1 0.5Mo steel tubes for heat exchangers of boiler application
Table 13. Chemical requirements of 0.5Mo steel tubes; JIS STBA12 [1], ASTM A209 T1
T1 [3]
Chemical composition [wt%]
Standards
Designation
C
Si
Mn
P
S
Mo
0.10 0.10 0.30 0.45 ≤
≤
JIS
STBA12
0.20
0.50
0.80
0.65
0.035
0.035
0.10 0.10 0.30 0.44 ≤
≤
T1
0.20
0.50
0.80
0.65
0.025
0.025
ASTM
0.10 0.10 0.30 0.44 ≤
≤
T1
0.20
0.50
0.80
0.65
0.025
0.025
[2] and A250
Std. No.
G3462
A209
A250
2.2.1.2.2 0.5Mo steel pipes for steam conductors of boiler piping applications
Table 14. Chemical requirements of 0.5Mo steel pipes; JIS STPA12 [1] and ASTM A335 P1 [4]
Chemical composition [wt%]
Standards
Designation
Std. No.
C
Si
Mn
P
S
Mo
0.10 0.10 0.30 0.45 ≤
≤
JIS
STPA12
G3458
0.20
0.50
0.80
0.65
0.035
0.035
0.44 0.10 0.10 0.30 ≤
≤
A335
ASTM
P1
0.65
0.20
0.50
0.80
0.025
0.025
Landolt-Börnstein
New Series VIII/2B
30
2.2 Low alloy steels
2.2.1.2.3 0.5Mo steel plates for boiler and pressure vessel applications
Table 15. Chemical requirements of 0.5Mo steel plates; JIS SB450M [1] and ASTM A204M Grade A [5]
Chemical composition [wt%]
Thickness t
Standards Designation
Std. No.
[mm]
C
Si
Mn
P
S
Mo
≤0.18
≤25
≤0.21 0.15 - ≤
25<
t ≤50
0.45
≤
≤
JIS
SB450M
G3103
0.30
0.60
0.90
0.035
0.040
≤0.23
50< t ≤100
≤0.25
100< t ≤150
≤0.18
≤25
≤0.21 0.15 - ≤
0.45 - 25< t ≤50
≤
≤
ASTM
Grade A
A204M
0.40
0.60
0.90
0.035
0.030
≤0.23
50< t ≤100
100< t
≤0.25
Table 16. Chemical requirements of 0.5Mo steel plates; JIS SB480M [1] and ASTM A204M Grade B [5]
Chemical composition [wt%]
Thickness t
Standards Designation
Std. No.
[mm]
C
Si
Mn
P
S
Mo
≤0.20
≤25
≤0.23 0.15 - ≤
0.45 - 25< t ≤50
≤
≤
JIS
SB480M
G3103
0.90 0.035 0.040 0.60
≤0.25 0.30
50< t ≤100
≤0.27
100< t ≤150
≤0.20
≤25
≤0.23 0.15 - ≤
25<
t ≤50
0.45 ≤
≤
ASTM
Grade B
A204M
0.90 0.035 0.035 0.60
≤0.25 0.40
50< t ≤100
100< t
≤0.27
Table 17. Chemical requirements of 0.5Mo steel plates; JIS SBV1A [1] and ASTM A302M Grade A [6]
Chemical composition [wt%]
Thickness t
Standards Designation
Std. No.
[mm]
C
Si
Mn
P
S
Mo
≤0.20
≤25
0.15 - 0.95 - ≤
0.45 ≤
JIS
SBV1A
G3119
≤0.23
25< t ≤50
0.30
1.30
0.035 0.040 0.60
≤0.25
50< t ≤150
≤0.20
≤25
0.15 - 0.95 - ≤
0.45 ≤
ASTM
Grade A
A302M
≤0.23
25< t ≤50
0.40
1.30
0.035 0.035 0.60
50< t
≤0.25
Table 18. Chemical requirements of 0.5Mo steel plates; JIS SBV1B [1] and ASTM A302M Grade B [6]
Chemical composition [wt%]
Thickness t
Standards Designation
Std. No.
[mm]
C
Si
Mn
P
S
Mo
≤0.20
≤25
0.15 - 1.15 - ≤
0.45 ≤
≤0.23
JIS
SBV1B
G3119
25< t ≤50
0.30
1.50
0.035 0.040 0.60
≤0.25
50< t ≤150
≤0.20
≤25
0.15 - 1.15 - ≤
0.45 ≤
≤0.23
ASTM
Grade B
A302M
25< t ≤50
0.40
1.50
0.035 0.035 0.60
≤0.25
50< t
Landolt-Börnstein
New Series VIII/2B
Ref. p. 34]
2.2.1 0.5Mo steel
31
Table 19. Chemical requirements of 0.5Mo steel plates; JIS SBV2 [1] and ASTM A302M Grade C [6]
Thickness t
Chemical composition [wt%]
Std. No.
Standards Designation
[mm]
C
Si
Mn
P
S
Mo
Ni
≤0.20
≤25
0.15 - 1.15 - ≤
0.45 - 0.40 ≤
JIS
SBV2
≤0.23
25< t ≤50 G3119
0.30 1.50 0.035 0.040 0.60 0.70
≤0.25
50< t ≤150
≤0.20
≤25
0.15 - 1.15 - ≤
0.45 - 0.40 ≤
ASTM
Grade C
≤0.23
25< t ≤50 A302M
0.40 1.50 0.035 0.035 0.60 0.70
50< t
≤0.25
Table 20. Chemical requirements of 0.5Mo steel plates; JIS SBV3 [1] and ASTM A302M Grade D [6]
Chemical composition [wt%]
Thickness t
Standards Desig-nation
Std. No.
[mm]
C
Si
Mn
P
S
Mo
Ni
≤0.20
≤25
0.15 - 1.15 - ≤
0.45 - 0.70 ≤
JIS
SBV3
≤0.23
25< t ≤50 G3119
0.30 1.50 0.035 0.040 0.60 1.00
≤0.25
50< t ≤150
≤0.20
≤25
0.15 - 1.15 - ≤
0.45 - 0.70 ≤
ASTM
Grade D
≤0.23
25< t ≤50 A302M
0.40 1.50 0.035 0.035 0.60 1.00
50< t
≤0.25
2.2.1.3 Creep properties of 0.5Mo steel tubes
The database [7] contains the creep data of 0.5Mo steel tubes, namely rupture data, minimum creep rate,
rupture elongation, reduction of area and microstructures of crept specimens.
2.2.1.3.1 Creep rupture data of 0.5Mo steel tubes
Fig. 29 shows the creep rupture data of STBA12 steel tubes of 12 heats.
2
Stress [N/mm ]
103
102
450 °C
500 °C
550 °C
10
10
102
Fig. 29. Creep rupture strength
data of STBA12 according to data
from [7].
103
104
Time to rupture [h]
Landolt-Börnstein
New Series VIII/2B
105
106
32
2.2 Low alloy steels
2.2.1.3.2 Time-Temperature-Parametric prognostication of the creep rupture strength
Fig. 30 shows Orr-Sherby-Dorn parametric plots of rupture data based on [7]. Creep rupture curve
regression by a cubic expression predicts the creep rupture strength for times longer than that of the
experiment at temperatures from 450 °C to 550 °C, Fig. 31.
Stress (N/mm2)
1000
100
Fig. 30. Master rupture curve by Orr-SherbyDorn parameter method for 0.5Mo steel tubes;
[7].
450 °C
500 °C
550 °C
TK: Temperature [K]
tr: time to rupture
OSDP fitting curve
10
-25
-23
-21
-19
-17
-15
OSDP = log tr - (348583/(19.1425 TK)
Stress (N/mm2)
1000
100
450 °C
500 °C
550 °C
Fig. 31. Estimated creep rupture curves
for 0.5Mo steel tubes; [7].
10
10
10
2
10
3
10
4
10
5
10
6
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 34]
2.2.1 0.5Mo steel
33
2.2.1.4 Creep properties of 0.5Mo steel plates
The database [8] contains creep data of 0.5Mo steel plates, namely rupture data, minimum creep rate,
rupture elongation, reduction of area and microstructures of crept specimens.
2.2.1.4.1 Creep rupture data of 0.5Mo steel plates
Fig. 32 shows the creep rupture data of SBV2 steel plates of 5 heats.
Stress (N/mm2)
1000
100
450 °C
Fig. 32. Creep rupture strength
data of SBV2 according to data
from [8].
500 °C
550 °C
10
1
10
2
10
4
10
6
Time to rupture (h)
2.2.1.4.2 Time-Temperature-Parametric prognostication of the creep rupture strength
Fig. 33 shows Manson-Haferd parametric plots of rupture data based on [8]. Creep rupture curve regression by a cubic expression predicts the creep rupture strength for times longer than that of the
experiment at temperatures from 450 °C to 550 °C , Fig. 34.
Stress (N/mm2)
1000
100
450°C
500 °C
550 °C
Fig. 33. Master rupture curve by Manson-Haferd
parameter method for 0.5Mo steel plates; [8].
MHP fitting curve
10
-6
Landolt-Börnstein
New Series VIII/2B
-5
-4
-3
MHP=[(log(tr)-11.169)/(Tk-530)]
-2
34
2.2 Low alloy steels
Stress (N/mm2)
1000
100
450 °C
500 °C
550 °C
Fig. 34. Estimated creep rupture
curves for 0.5Mo steel plates;
[8].
10
10
102
3
10
4
10
5
10
Time to rupture (h)
2.2.1.5 References
[1]
[2]
[3]
[4]
[5]
[6]
[7]
[8]
JIS Handbook.
ASTM Standard: A209/A209M (2001).
ASTM Standard: A250/A250M (2001).
ASTM Standard: A335/A335M (2001).
ASTM Standard: A204/A204M (2001).
ASTM Standard: A302/A302M (2001).
National Research Institute for Metals: NRIM Creep Data Sheet, 8B (1991).
National Research Institute for Metals: NRIM Creep Data Sheet, 18B (1987).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 37]
2.2.2 High strength steel
35
2.2.2 High strength steel
2.2.2.1 Introduction
High strength steels are applied for pressure vessels. The application temperature range of the steels,
categorized as silicon-manganese steels, is up to 350 °C. Therefore, allowable stresses are determined by
the tensile stresses at service temperatures. Some steels are enhanced in tensile properties by MC type
carbide precipitation strengthening. In order to obtain enough toughness and weldability for fabrication,
the carbon, silicon and manganese contents are restricted in standards with respect to the parameters Ceq1
and PCM2 as described in JIS G3115 and in ISO 9328-4. Heat treatment manufacturing process is also
given to unify the metallurgical microstructure as in JIS G3119 for instance.
2.2.2.2 Material standards, chemical and tensile requirements
2.2.2.2.1 High strength steel plates for pressure vessels
Table 21. Chemical requirements of high strength steel plates; JIS SPV490 [1] and ISO P500TQ [2]
Chemical composition [wt%]
Thickness
Standards Designation
Std. No.
t [mm]
C
Si
Mn
P
S
others
0.15
6< t ≤50
Ceq≤0.45
≤
≤
≤
G3115
JIS
SPV490
0.18 1.60
0.030 0.030 or PCM≤0.28 50< t ≤75
0.75
Cr≤2.0
Mo≤1.0
Ni≤2.0
Cu≤1.50
Nb≤0.06
0.70 - ≤
≤
≤
≤
ISO
P500TQ
9328-4
Ti≤0.20
3≤ t ≤70
0.20 0.55 1.70
0.030 0.030
V≤0.10
Al≤0.020
B≤0.005
N≤0.020
Zr≤0.15
2.2.2.2.2 High strength steel plates for general structures
Table 22. Chemical requirements of high strength steel plates; JIS SM570[1], ASTM A678 C and A678
D [3]
Chemical composition [wt%]
Standards
Designation
Std. No.
C
Si
Mn
P
S
others
≤0.035
≤0.035
≤
≤
≤
G3106
JIS
SM570
0.18 0.55
1.60
0.20 - 1.00 - ≤0.035
≤0.04
≤
A678
C
1.60
0.22 0.50
ASTM
0.15 - 1.15 - ≤0.035
V 0.04-0.11
≤0.04
≤
A678
D
1.50
N 0.01-0.03
0.22 0.50
1
2
Ceq = C+Si/24+Mn/6+Ni/40+Cr/5+Mo/4+V/14
PCM = C+Si/30+Mn/20+Cu/20+Ni/60+Cr/20+Mo/15+V/10+5B
Landolt-Börnstein
New Series VIII/2B
36
2.2 Low alloy steels
2.2.2.3 Creep properties of high strength steel plates
The database [4] contains creep data of high strength steel plates, namely rupture data, rupture elongation,
reduction of area, microstructures of as-received materials and crept specimens.
2.2.2.3.1 Creep rupture data of high strength steel plates
Fig. 35 shows the creep rupture data of high strength steel plates of 21 heats. Several creep tests are still
continuing.
Stress (N/mm2)
1000
100
400°C
450°C
500°C
550°C
10
10
10
2
3
4
10
10
Time to rupture (h)
5
10
6
10
Fig. 35. Creep rupture strength
data of high strength steel plates;
[4].
2.2.2.3.2 Time-Temperature-Parametric prognostication of creep rupture strength
Fig. 36 shows Orr-Sherby-Dorn parametric plots of rupture data based on [4]. Creep rupture curve
regression by a cubic expression predicts the creep rupture strength for times longer than that of the
experimental data at temperatures from 400 °C to 550 °C, Fig. 37.
700
○ 400 ℃
△ 450 ℃
□ 500 ℃
▽ 550 ℃
Stress (MPa)
500
300
100
80
60
40
-22
― Average
n = 382
-20
Fig. 36. Master rupture curve by Orr-SherbyDorn parameter method for high strength steel
plates; [4].
-18
-16
-14
-12
logtR - [ 284907 / ( 19.1425 × Tk ) ]
Landolt-Börnstein
New Series VIII/2B
Ref. p. 37]
2.2.2 High strength steel
37
700
Stress ( MPa )
500
300
400 ℃
500 ℃
100
80
60
40
101
450 ℃
550 ℃
102
103
104
105
Time to rupture ( h )
106
Fig. 37. Estimated creep rupture
curves for high strength steel
plates; [4].
2.2.2.4 References
[1]
[2]
[3]
[4]
JIS Handbook.
ISO Standard: 9328-4.
ASTM Standard: A678 (2001).
National Research Institute for Metals: NRIM Creep Data Sheet, 25B (1994).
Landolt-Börnstein
New Series VIII/2B
38
2.2 Low alloy steels
2.2.3 0.5Cr-0.5Mo steel
2.2.3.1 Introduction
0.5Cr-0.5Mo steels are used as tubes for heat exchangers, as pipes for high temperature service, as plates
and forgings for pressure vessels of power plants, chemical and petrochemical plants. The creep strength
of this steel is improved by addition of 0.5% of molybdenum. Oxidation and corrosion resistance are
improved by addition of 0.5% of chromium. The creep strength is influenced by initial microstructure and
changes in microstructure during creep exposure. Sigmoidal inflection of the stress vs. time to rupture
curve is caused by decrease in creep strength and advent of inherent creep strength due to microstructural
change during creep exposure, as will be mentioned later. The creep strength is affected by initial
microstructure even for long-term creep exposure. The creep strength of steel with fully annealed ferrite
and pearlite microstructure is higher than that of steel with martensitic and bainitic microstructures, as
will be explained later.
2.2.3.2 Material standards, chemical and tensile requirements
2.2.3.2.1 0.5Cr-0.5Mo steel tubes for heat exchangers
Table 23.
Standards
JIS
ASTM
Chemical requirements of 0.5Cr-0.5Mo steel tubes; JIS STBA 20, ASTM T2
Chemical composition [wt%]
Designation
C
Si
Mn
P
S
Cr
Mo
STBA 20 0.10-0.20 0.10-0.50 0.30-0.60 ≤0.035 ≤0.035 0.50-0.80 0.40-0.65
T2
0.10-0.20 0.10-0.30 0.30-0.61 ≤0.025 ≤0.025 0.50-0.81 0.44-0.65
Std.
No
G3462
A213
2.2.3.2.2 0.5Cr-0.5Mo steel pipes for high temperature services
Table 24.
Standards
JIS
ASTM
Chemical requirements of 0.5Cr-0.5Mo steel pipes; JIS STPA 20, ASTM P2
Chemical composition [wt%]
Designation
C
Si
Mn
P
S
Cr
Mo
STPA 20 0.10-0.20 0.10-0.50 0.30-0.60 ≤0.035 ≤0.035 0.50-0.80 0.40-0.65
P2
0.10-0.20 0.10-0.30 0.30-0.61 ≤0.025 ≤0.025 0.50-0.81 0.44-0.65
Std.
No
G3458
A335
2.2.3.2.3 0.5Cr-0.5Mo steel plates for pressure vessels
Table 25.
Standards
JIS
ASTM
Chemical requirements of 0.5Cr-0.5Mo steel plates; JIS SCMV 1-2, ASTM Gr.2 cl.2
Chemical composition [wt%]
Designation
C
Si
Mn
P
S
Cr
Mo
SCMV 1-2 ≤0.21
0.55-0.80 ≤0.030 ≤0.030 0.50-0.80 0.45-0.60
≤0.40
Gr.2 cl.2
0.05-0.21 0.15-0.40 0.55-0.80 ≤0.035 ≤0.035 0.50-0.80 0.45-0.60
Std.
No
G4109
A387
Landolt-Börnstein
New Series VIII/2B
Ref. p. 44]
2.2.3 0.5Cr-0.5Mo steel
39
2.2.3.2.4 0.5Cr-0.5Mo steel forgings for pressure vessels
Table 26. Chemical requirements of 0.5Cr-0.5Mo steel forgings; JIS SFVA F2, ASTM F2
Standards
Designation
Std.
No
Chemical composition [wt%]
C
Si
Mn
JIS
SFVA F2 ≤0.20
0.30-0.80
≤0.60
ASTM F2
0.05-0.21 0.10-0.60 0.30-0.80
P
≤0.030
≤0.040
S
Cr
Mo
≤0.030 0.50-0.80 0.45-0.65 G3203
≤0.040 0.50-0.81 0.44-0.65 A182
2.2.3.3 Creep properties of 0.5Cr-0.5Mo steel tubes
Information of fact on creep data for 0.5Cr-0.5Mo steel tubes can be obtained from [1] and [2].
2.2.3.3.1 Creep rupture data of 0.5Cr-0.5Mo steel tubes
The creep rupture strength of 0.5Cr-0.5Mo steel tubes obtained from available creep data sources is
shown in Fig. 38. The results of creep test for 9 heats of JIS STBA 20 steel tubes are compiled in [1].
From this data sheet the data of rupture elongation, reduction of area, minimum creep rate, time to
specified strain and microstructures of as-received materials and crept specimens can be also obtained.
The creep rupture curve at 550 °C shows a sigmoidal shape, which is caused by changes in creep strength
due to microstructural change during creep exposure, as will be explained later.
700
Stress (MPa)
500
300
100
80
60
○
△
□
▽
40
-1
10
450oC
500oC
550oC
600oC
0
10
n=228
10
1
10
2
10
3
4
10
10
5
10
6
Fig. 38. Creep rupture strength
data of JIS STBA 20; [1].
n indicates the total number of
data points.
Time to rupture (h)
2.2.3.3.2 Creep rupture strength of 0.5Cr-0.5Mo steel tubes
Creep rupture strength was analyzed by applying the Manson-Haferd parameter method to NRIM creep
rupture data on 0.5Cr-0.5Mo steel tubes (JIS STBA 20). The result is shown in Fig. 39. It should be noted
that the heat-to-heat variation of creep rupture strength is very large under higher stress conditions,
however, the scatter band of creep rupture strength under lower stress conditions is very narrow, in
comparison with that under higher stresses. Decrease in heat-to-heat variation of creep rupture strength
with decrease in applied stress is caused by decrease in creep strength and advent of inherent creep
strength due to microstructural changes during long-term creep exposure at the elevated temperatures, as
will be explained later [3, 4].
Landolt-Börnstein
New Series VIII/2B
40
2.2 Low alloy steels
Stress [MPa]
1000
800
600
500
400
300
450 °C
500 °C
550 °C
575 °C
600 °C
625 °C
200
100
80
60
50
Average
40
n = 231
30
-2.5
-4.5
-3.0
-3.5
-4.0
Manson-Haferd parameter [( log tR -18.087)/( TK -304.0 )] [×10 2 ]
Fig. 39. Master rupture curve by Manson-Haferd
parameter method for 0.5Cr-0.5Mo steel tubes (JIS
STBA 20); [1]. n indicates the total number of data
points.
2.2.3.3.3 Microstructural change
The typical initial microstructure of 0.5Cr-0.5Mo steel tubes consists of ferritic and pearlitic grains.
Optical micrographs of 0.5Cr-0.5Mo steel tubes are shown in Fig. 40. The bright grains are ferritic grains
and the dark ones are pearlitic.
Bright field TEM images within ferritic grains of 0.5Cr-0.5Mo steel tubes after crept and creep
ruptured at 550 °C and stresses of 294, 196, 88 and 59 MPa are shown in Fig. 41 [3, 4]. In the specimen
creep ruptured after 155.8 h under stress of 294 MPa (Fig. 41a), huge amounts of dislocations with a lot
of very fine carbide particles are observed. After creep for 1,778.3 h under stress of 196 MPa (Fig. 41b),
the dislocation density is still high and carbide particles are coarsened. After creep for 23,788.3 h under
stress of 88 MPa (Fig. 41c), precipitation of many needle-like Mo2C carbides is observed and the
dislocation density is low in comparison with that in specimens creep ruptured after short-term creep
exposure. After long-term creep exposure for 112,776.4 h under stress of 59 MPa (Fig. 41d), further
coarsening of carbide is observed and the dislocation density is significantly lowered.
Fig. 40. Optical micrographs of as-received 0.5Cr-0.5Mo steel tubes (etched in 4 % natal); [1].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 44]
2.2.3 0.5Cr-0.5Mo steel
41
Fig. 41. Bright field TEM images within ferritic grains of 0.5Cr-0.5Mo steel tubes after crept and
creep ruptured at 550 °C at stresses of 294 (a), 196 (b), 88 (c) and 59 MPa (d); [3, 4].
Rupture times: a) tr = 155.8 h, b) tr = 1778.3h, c) tr = 23788.3 h, d) tr = 112776.4 h.
2.2.3.3.4 Inherent creep strength
Stress vs. time to rupture curves of 0.5Cr-0.5Mo steel tubes at 550 and 600 °C are shown in Fig. 42 [3, 4].
Sigmoidal inflection is observed for the creep rupture curve at 550 °C. The slope of the curve increases
with decrease in stress from 300 to 150 MPa, however, the curve turns to be gentle at about 100 MPa.
Inflection of the curve at about 100 MPa is observed also at 600 °C, similar to that at 550 °C.
Sigmoidal inflection of stress vs. time to rupture curves is caused by changes in creep strength due to
microstructural change during creep exposure at elevated temperatures as shown in Fig. 41. Creep
strength of 0.5Cr-0.5Mo steel tubes is strongly influenced by precipitation and coarsening of carbides and
changes in dislocation density during long-term creep exposure.
Changes in hardness within ferritic grains of 0.5Cr-0.5Mo steel tubes with increase in creep exposure
time at 550 and 600 °C are shown in Fig. 43. With increase in creep exposure time, hardness decreases at
both temperatures. However, the magnitude of the decrease in hardness decreases after about 10,000 h
and 1,000 h of creep exposure at 550 and 600 °C, respectively. The hardness of crept 0.5Cr-0.5Mo steel
tubes indicates an almost constant value from HV130 to HV150 for long-term creep exposure.
Strengthening effects obtained by fine precipitates and high dislocation density are essentially lost
during long-term creep exposure for about 10,000 h and 1,000 h at 550 and 600 °C, respectively. The
constant hardness value observed for long-term creep exposure corresponds to that of the ferrite matrix
itself, consequently, the creep strength after long-term creep exposure is the same as that of the ferrite
matrix. The slope of the stress vs. time to rupture curve increases with decrease in applied stress, since
creep strength decreases due to microstructural change, however, the creep strength decreases to that of
the ferrite matrix after long-term creep exposure and the creep rupture curve turns to be gentle.
Landolt-Börnstein
New Series VIII/2B
42
2.2 Low alloy steels
The creep strength corresponding to that of the ferrite matrix is constant and independent of creep
exposure time, and it has been proposed as a concept of “Inherent Creep Strength” [3, 4]. Long-term
creep strength of 0.5Cr-0.5Mo steel is governed by inherent creep strength.
250
500
550oC
Vickers hardness (98N)
Stress (MPa)
300
600oC
100
80
60
40 2
10
3
10
10
4
10
5
10
6
Time to rupture (h)
Fig. 42. Stress vs. time to rupture curves of 0.5Cr0.5Mo steel tubes at 550 and 600 °C; [3, 4].
550oC
200
150
600oC
100
2
10
3
10
10
4
10
5
10
6
Time to rupture (h)
Fig. 43. Changes in hardness within ferritic grains with
increase in creep exposure at 550 and 600 °C; [3, 4].
2.2.3.3.5 Influence of initial microstructure on long-term creep strength
The creep strength of 0.5Cr-0.5Mo steel is strongly influenced by initial microstructure, which results in
large scatter of short-term creep strength as shown in Fig. 39. The microstructure of the steel is affected
by heat treatment conditions. Bright field TEM images of 0.5Cr-0.5Mo steels subjected to various
conditions of heat treatment are shown in Fig. 44. Martensitic microstructure (a), tempered martensitic
microstructures (b,c), bainitic microstructure (d) and a complex of ferritic and pearlitic grains (e) are
obtained by different heat treatment conditions.
The creep rupture strengths of 0.5Cr-0.5Mo steels with different initial microstructures shown in Fig.
44 are shown in Fig. 45 [5, 6]. For creep exposure longer than 10,000 h, creep rupture strengths of the
steels with martensitic microstructures are almost the same independent of tempering heat treatment prior
to the creep test. Differences in microstructure due to tempering heat treatment essentially disappear
during creep exposure for about 10,000 h at 575 °C. Long-term creep strength in the stress range below
50 MPa and creep rupture strength of furnace cooled steel with a complex of ferritic and pearlitic grains
are higher than those of other steels with martensitic or bainitic microstructures. Although the creep
strength in the range of stresses lower than about 50 MPa is thought to be inherent creep strength of the
steel, creep strength of furnace cooled steel is clearly higher than that of other steels.
In contrast to very high dislocation densities of steels with martensitic or bainitic microstructures in
the as heat treated condition, the dislocation density of furnace cooled steels is significantly lower. Creep
deformation can be carried with dislocations generated by applied stress high enough to produce them,
even if the amount of dislocations is small in the furnace cooled condition. However, if the applied stress
is such low that it can not produce enough dislocations for creep deformation, the creep rate should be
very small.
Long-term creep strength of 0.5Cr-0.5Mo steel is governed by inherent creep strength, which is
independent of initial microstructure. However, long-term creep rupture strength of furnace cooled
microstructure with very low dislocation density is higher than of other steels containing a lot of
dislocations, since there are not enough dislocations for creep deformation.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 44]
2.2.3 0.5Cr-0.5Mo steel
43
Fig. 44. Bright field TEM images of 0.5Cr-0.5Mo steels in the as heat treated conditions of (a) quenched from
920 °C, (b) and (c) quenched from 920 °C followed by tempering at 650 °C, (d) quenched from 920 °C to 450 °C and
isothermally transformed for 1 h and (e) furnace cooling from 920 °C; [5, 6]
200
Stress (MPa)
575oC
100
90
80
70
60
50
40
30
○ Martensite
o
△ Tempered martensite (650 C/1h)
o
▽ Tempered martensite (650 C/100h)
□ Bainite
o
o
● Ferrite + Pearlite
20 1
10
10
2
10
3
10
4
10
5
6
10
Fig. 45. Creep rupture strength of 0.5Cr0.5Mo steels with different initial microstructures; [5, 6].
Time to rupture (h)
2.2.3.3.6 Estimated long-term creep strength
The temperature dependence of 0.2% proof stress, tensile strength and creep rupture strength at 100 and
10,000 h for 9 heats of 0.5Cr-0.5Mo steel tubes is shown in Fig. 46 [1]. That of 0.2% proof stress, tensile
strength and creep rupture strength at 1,000 and 100,000 h for the same materials is shown in Fig. 47 [1].
Creep rupture strength curves shown in Fig. 46 and Fig. 47 were obtained by regression analysis using the
Manson-Haferd parameter.
Landolt-Börnstein
New Series VIII/2B
44
2.2 Low alloy steels
Stress [MPa]
1000
800
600
500
400
300
Tensile
strength
0.2% proof
stress
200
100
80
60
50
40
30
400
100 h
10000 h
450
500
550
Temperature [°C]
600
Stress [MPa]
1000
800
600
500
400
300
Tensile
strength
0.2% proof
stress
200
100
80
60
50
40
30
400
650
Fig. 46. Temperature dependence
of 0.2% proof stress, tensile
strength and creep rupture strength
at 100 and 10,000 h for 0.5Cr0.5Mo steel tubes (JIS STBA 20);
[1].
1000 h
100000 h
450
500
550
Temperature [°C]
600
650
Fig. 47. Temperature dependence
of 0.2% proof stress, tensile
strength and creep rupture strength
at 1,000 and 100,000 h for 0.5Cr0.5Mo steel tubes (JIS STBA 20);
[1].
2.2.3.4 References
[1]
[2]
[3]
[4]
National Research Institute for Metals: NRIM Creep Data Sheet, No.20B, (1994).
Japan Pressure Vessel Research Committee: 0.5Mo and Cr-Mo steels Data Book, (1998).
Kimura, K., Kushima, H., Yagi, K., and Tanaka, C.: Tetsu-to-Hagane, 77, (1991), 667-674.
Kimura, K., Kushima, H., Yagi, K., and C. Tanaka: Proc. of Inter. Conf. on Creep and Fracture of
Engineering Materials and Structures, Wilshire, B., and Evans, R.W., Swansea, eds., UK, The
Institute of Materials, 5, (1993), 555-564.
[5] Kimura, K., Kushima, H., Baba, E., Shimizu, T., Asai, Y., Abe, F., and Yagi, K.: Proc. of 5th Inter.
Charles Parsons Turbine Conf. on Advanced Materials for 21st Century Turbines and Power Plant,
Strang, A., et al. eds., Cambridge, UK, The Institute of Materials, 5, (2000), 558-571.
[6] Kimura, K., Kushima, H., Baba, E., Shimizu, T., Asai, Y., Abe, F., and Yagi, K.: Tetsu-to-Hagane,
86, (2000), 542-549.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 53]
2.2.4 1Cr-0.5Mo steel
45
2.2.4 1Cr-0.5Mo steel
2.2.4.1 Introduction
1Cr-0.5Mo steels are used as tubes for heat exchangers and as plates for pressure vessels. The 1Cr-0.5Mo
steel tubes were introduced in the 1950s. As the microstructure of this steel is strongly affected by heat
treatment conditions and changes in microstructure are related with creep strength, the changes in
microstructure during creep have been investigated. For this steel, M2C carbides precipitate in ferrite and
M23C6 carbides precipitate in pearlite or in tempered bainite during creep. Studies on the effect of Al and
N on creep strength have also been done.
1Cr-0.5Mo steel plates are used as materials for pressure vessels of petroleum refinery. For this steel,
the reduction of creep ductility due to long-term service becomes a subject of investigation, and many
studies have been done for the prevention of this creep embrittlement.
2.2.4.2 Material standards, chemical and tensile requirements
2.2.4.2.1 1Cr-0.5Mo steel tubes for heat exchangers
Table 27. Chemical requirements of 1Cr-0.5Mo steel tubes; JIS STBA 22, ASTM T12, BS 620 and DIN
13CrMo44
Standards Designation
JIS
STBA 22
ASTM
T12
BS
620
DIN
13CrMo44
C
Si
≤0.15 ≤0.50
0.050.15
0.100.15
0.100.15
≤0.50
0.100.35
0.100.35
Chemical composition [wt%]
Std. No
Mn P
S
Cr
Mo Ni
Others
0.300.80- 0.45G3462
≤0.035 ≤0.035
0.60
1.25 0.65
0.300.80- 0.44A213
≤0.025 ≤0.025
0.61
1.25 0.65
0.400.70- 0.45A1≤
3606
≤0.040 ≤0.040
≤0.30
0.70
1.10 0.65
0.020
0.400.70- 0.4517175
≤0.035 ≤0.035
0.70
1.10 0.65
Table 28. Tensile properties at room temperature of 1Cr-0.5Mo steel tubes; JIS STBA22
Tensile strength
Yield strength
Elongation
[N/mm2]
[N/mm2]
[%]
d<10 mm
d≥20 mm
10≤d<20 mm
≥410
≥205
≥30
≥25
≥22
2.2.4.2.2 1Cr-0.5Mo steel plates for pressure vessels
Table 29. Chemical requirements of 1Cr-0.5Mo steel plates; JIS SCMV 2, ASTM Gr.12, EN 13CrMo4-5
Chemical composition [wt%]
Standards
Designation
Std.No
C
Si
Mn P
S
Cr
Mo Cu
Others
0.400.80- 0.45JIS
SCMV 2-2 ≤0.17 ≤0.40
G4109
≤0.030 ≤0.030
0.65
1.15 0.60
0.05- 0.15- 0.400.80- 0.45A 387M
ASTM
Gr.12-cl.2
≤0.035 ≤0.035
0.17 0.40 0.65
1.15 0.60
0.080.400.70- 0.40BS EN,
13CrMo4-5
10028-2
≤0.35
≤0.030 ≤0.025
≤0.30
0.18
1.00
1.15 0.60
DIN EN
Landolt-Börnstein
New Series VIII/2B
46
2.2 Low alloy steels
Table 30. Tensile properties at room temperature of 1Cr-0.5Mo steel plate; JIS SCMV 2-2.
Tensile strength
[N/mm2]
Yield strength
[N/mm2]
Elongation
[%]
450-590
≥275
≥22
2.2.4.3 Creep properties of 1Cr-0.5Mo steel tubes
Information of fact on creep data for 1Cr-0.5Mo steel tubes can be obtained from [1], [3], [4] and [5].
2.2.4.3.1 Creep rupture data of 1Cr-0.5Mo steel tubes
Creep rupture strength data of 1Cr-0.5Mo steel tubes is shown in Fig. 50. The results of creep tests for 11
heats of STBA 22 steel tubes are compiled in [1]. From this data sheet the data of elongation, reduction of
area and minimum creep rate, and microstructures of as-received materials and crept specimens can also
be obtained.
The relation of stress vs. time to rupture at 550 °C shows an inverse sigmoidal shape. This shape of
the creep rupture curve is due to the fact that the creep deformation curve is changed complicatedly by
microstructural evolutions during creep, as will be seen later.
2.2.4.3.2 Creep rupture strength of 1Cr-0.5Mo steel tubes
It should be noted from Fig. 50 that creep rupture strength has a large scatter. The creep rupture strength
is dependent on manufacturing conditions, chemical composition, and initial microstructure. This
information is given in [1].
Creep rupture curves were analyzed using the Manson-Haferd method to NRIM creep data. The result
is shown in Fig. 51, The creep rupture strength was estimated.
Tensile strength
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
400
300
400
300
200
200
100
100
0
0
100 200 300 400 500 600 700 800
Test temperature (℃)
0
0
100 200 300 400 500 600 700 800
Test temperature (℃)
Fig. 48. Tensile properties of 1Cr-0.5Mo steel tubes; [1].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 53]
2.2.4 1Cr-0.5Mo steel
47
Tensile strength
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
400
300
400
300
200
200
100
100
0
0
0
100 200 300 400 500 600 700 800
0
100 200 300 400 500 600 700 800
Test temperature (℃)
Test temperature (℃)
Fig. 49. Tensile properties of 1Cr-0.5Mo steel plates; [2].
500
500 °C
550 °C
600 °C
650 °C
Stress [MPa]
300
100
80
60
40
n = 309
20
10
10 2
500
400
10 5
10 6
500 °C
550 °C
570 °C
580 °C
600 °C
625 °C
640 °C
650 °C
675 °C
300
200
Stress [MPa]
10 3
10 4
Time to rupture [h]
Fig. 50. Creep rupture strength data
of STBA 22; [1]. n indicates the total
number of data points.
100
80
60
50
40
30
Average
n = 315
20
-2.0
-4.0
-3.0
-2.5
-3.5
-4.5
Manson-Haferd parameter [( log tR -13.088)/( TK -510.0 )] [×10 2 ]
Landolt-Börnstein
New Series VIII/2B
Fig. 51. Master rupture curve by Manson-Haferd parameter method for 1Cr-0.5Mo steel tubes. n indicates the
total number of data points.
48
2.2 Low alloy steels
500
300
500 °C
Stress [MPa]
550 °C
100
80
600 °C
650 °C
60
40
20
10
10 2
10 4
10 3
Time to rupture [h]
10 5
10 6
Fig. 52. Estimated creep rupture curves of STBA 22 steel
tubes.
2.2.4.3.3 Microstructural changes
The microstructure of 1Cr-0.5Mo steels changes during creep. The carbides precipitate, M2C in ferrite
portion, and M23C6 in pearlite or bainite portion. The form of these carbides changes with aging time.
The percentage of alloying elements in cementite for post-service material was tested at 823 K as a
function of creep exposure time [7]. The material which was investigated in this paper was supplied in the
form of pipe section, which had been removed from service after about 70,000 h at 838 K, and a nominal
stress of 17.2 MNm−2. The service-exposed material contained a large amount of spheroidized M3C in the
bainitic regions with M2C precipitates in the form of fine needles within the ferrite grains. It is suggested
that these changes in the substitutional solute element concentrations in cementite could provide an aid to
the estimation of effective exposure temperature for use in determining the remanent life of components.
Continuous microstructural changes lead to property degradation; the reduction of the solid-solution Mo
content and the increased interparticle spacing consequent on spherodization reduce creep resistance.
The effect of heat treatment on creep rupture strength and microstructural change was investigated for
annealed 1Cr-0.5Mo steel (Ann.) and for normalized and tempered one (NT) [8]. Based on the study on
the change in creep rupture strength and carbide distribution, for as-heat-treated materials, the annealed
steel has a larger amount of precipitates in ferrite and a higher creep strength. However, after long-term
aging, there is no difference in the amount of precipitates between both steels, and the inversion of creep
strength can be seen at longer times.
2.2.4.3.4 Creep deformation behavior and creep rupture strength
Inverse sigmoidal shaped stress vs. time to rupture curves are observed for 1Cr-0.5Mo steels. The
sigmoidal shape is caused by the transition of creep deformation behavior due to microstructural changes.
Fig. 53 shows creep deformation, strain vs. time and creep rate vs. time curves, obtained under testing
conditions where stress vs. time to rupture relations with inverse sigmoidal shape are observed. The creep
rate curve has two minimum values in one curve. The creep rate curve of material pre-aged for 500 h at
873 K has one minimum value [9].
Fig. 54 shows the comparison of experimental and predicted creep rupture life of STBA 22 [9]. The
creep deformation curve with two minimum values of creep rate was divided into two regions, Type I and
Type II. The creep rupture life was predicted from each creep deformation analysis of Type I and Type II
regions. For higher stresses and shorter times, experimental creep rupture life agrees well with the curves
predicted from Type I creep deformation behavior, and for lower stresses and longer times, experimental
Landolt-Börnstein
New Series VIII/2B
Ref. p. 53]
2.2.4 1Cr-0.5Mo steel
49
data agree well with the curves predicted from Type II. It is considered that since experimental short-term
life of pre-aged material agree with the curve predicted from Type II, the change in controlling factor
from Type I to Type II of creep rupture strength is caused by the change in controlling factor of creep
deformation behavior due to microstructural evolution during creep.
The relation of time to rupture vs. minimum creep rate of STBA 22 was investigated [10]. This
relation is well known to be linear in double logarithmic scales. For Cr-Mo ferritic heat resistant steels,
however, this relation has a large scattering, which is explained by the change in combination of
controlling factor of time to rupture and minimum creep rate.
2.2.4.3.5 Effect of Al and N on creep rupture strength
For carbon steels, the effect of Al and N on creep rupture strength has been investigated. However, for
1Cr-0.5Mo steels, this effect is more complicated because N interacts with Cr and Mo.
The effect of Al and N on creep rupture strength of 1Cr-0.5Mo steel was examined [11]. Heat
treatment conditions of this steel are 920 °C × 30 min → 720 °C × 25 min, AC. The creep rupture
strength strongly depends on the amount of Al. Smaller amounts of Al produce higher creep rupture
strength. All of nitrogen in the steel exists as CrN. This means that the detrimental effect of Al addition
cannot be ascribed to the decrease in the amount of active N. The decrease in creep rupture strength due
to Al addition is caused by grain refining effect of AlN.
Fig. 55 shows the amount of soluble nitrogen in 1Cr-0.5Mo steel [12]. There is little amount of
soluble N in low alloy steels, and strengthening due to soluble N is not expected for 1Cr-0.5Mo steel.
5
1 Cr-0.5Mo steel (JIS STBA 22)
300
3
2
Stress [MPa]
Strain e [%]
4
500
a
823 K
88 MPa
s / E = 0.509×10-3
1
0
b
100
80
873 K
60
Creep rate e [%/h]
Calculated
40
10-3
Type I
Type II
un- aged
pre-aged for 500h at 873 K
20
10
10-4
Measured
Calculated
10-5
0
{
5000
{
10000
15000
Time t [h]
Fig. 53. Creep deformation curves; [9].
Landolt-Börnstein
New Series VIII/2B
Type I
Type II
20000
25000
10 2
10 4
10 3
Time to rupture [h]
10 5
10 6
Fig. 54. Comparison of experimental and predicted
creep rupture life of STBA22; [9].
50
2.2 Low alloy steels
100
1 Cr-0.5Mo steel
2.25 Cr-1Mo steel
Extraction rate [10-6 %/°C ]
80
60
40
20
0
0
200
600
400
800
1000
Extraction temperature [°C ]
1200
Fig. 55. Extraction curves of nitrogen for as-received
and crept 1Cr-0.5Mo steel and 2.25Cr-1Mo steel; [12].
2.2.4.3.6 Estimated long-term creep rupture strength
800
600
500
400
{
200
{
Stress [MPa]
300
100
80
60
50
40
30
20
450
Tensile
strength
0.2% proof
stress
1000 h
100000 h
500
550
600
Temperature [°C]
650
700
Fig. 56. Creep rupture strength of
STBA 22 steel tubes [1]. The
relations stress vs. temperature
with the parameter time to rupture
were estimated based on the most
suitable master rupture curve,
which was obtained from NRIM
Creep Data using Manson-Haferd
method.
2.2.4.4 Creep properties of 1Cr-0.5Mo steel plates
Information of fact on creep data for 1Cr-0.5Mo steel plates can be obtained from [2], [3] and [5].
Creep rupture strength data of 1Cr-0.5Mo steel plates is shown in Fig. 57. The results of creep tests
for 8 heats of SCMV 2 NT steel plates are compiled in [2]. From this data sheet the data of elongation and
reduction of area, and microstructures of as-received materials can also be obtained.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 53]
2.2.4 1Cr-0.5Mo steel
51
500
Stress [MPa]
300
100
80
60
40
20
10 -1
450 °C
475 °C
500 °C
550 °C
600 °C
650 °C
1
10
10 3
10 2
Time to rupture [h]
10 4
10 5
10 6
Fig. 57. Creep rupture strength
data of SCMV 2 NT; [2].
2.2.4.4.1 Creep rupture strength of 1Cr-0.5Mo steel plates
It should be noted from Fig. 57 that creep rupture strength has large scatter. The creep rupture strength
depends on manufacturing conditions, chemical composition, and initial microstructure. This information
is obtained from [2].
The master rupture curve, Fig. 58, was analyzed using the Orr-Sherby-Dorn method to NRIM creep
data, which are shown in Fig. 57, and creep rupture strength was estimated, Fig. 59.
2.2.4.4.2 Creep properties of coarse grained 1Cr-0.5Mo steel
The creep ductility of 1Cr-0.5Mo steels is low, and especially, the understanding of creep ductility at
heat-affected zones (HAZ) is important for this steel.
Fig. 60 shows the creep test results of smooth specimens and notched specimens for Cr-Mo steels
[13]. The coarse grained heat affected zones of 1Cr-0.5Mo steel show notch weakening due to decrease in
creep rupture ductility. The rupture ductility increases with increasing amount of Cr, and the material
shows the notch strengthening. The influence of postweld heat treatment on creep properties has also
been investigated [14].
The multiaxial creep properties of coarse grained 1Cr-0.5Mo steel were investigated [15]. The
material is quenched and tempered, and has a microstructure comprising coarse grained tempered bainite
with a prior austenite grain size of about 190 µm. The minimum creep rate stress index is in the range of
2-3 and the ductility is about 1 %. This material is creep brittle. The multiaxial stress rupture criterion
(MSRC) of this steel varies with stress state; at high triaxiality (notch), MSRC is dependent on maximum
principle stress, and at low triaxiality (shear), MSRC is dependent on both maximum principle stress and
equivalent stress.
Landolt-Börnstein
New Series VIII/2B
52
2.2 Low alloy steels
100
500 °C
550 °C
600 °C
650 °C
300
100
80
550 °C
1Cr- 1/2 Mo
1 1/4 Cr- 1/2 Mo
2 1/4 Cr-1Mo
3Cr-1Mo
60
40
20
0
60
50
40
Average
n = 158
30
20
-26
-24
-18
-22
-20
Orr-Sherby-Dorn parameter log tR -[390498/( 19.1425×TK )]
Stress [kgf / mm 2 ]
Stress [MPa]
200
80
Reduction of area [%]
600
500
400
30
25
20
15
10
Fig. 58. Master rupture curve by Orr-Sherby-Dorn
parameter method for 1Cr-0.5Mo steel plates.
n indicates the total number of data points.
0
open: smooth
specimen
solid: notched
specimen
10 2
10 3
Time to rupture [h]
10 4
Fig. 60. Creep test result of smooth specimens and
notched specimens for Cr-Mo steels; [13].
500
Stress (MPa)
o
500 C
340
o
550 C
180
o
600 C
Fig. 59. Estimated creep rupture
curves of SCMV 2 NT steel
plates.
o
650 C
20
101
102
103
104
105
106
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 53]
2.2.4 1Cr-0.5Mo steel
53
2.2.4.4.3 Crack growth behavior in the creep range
Components and structural parts are mostly subjected to strongly alternating loads during operation
because of start-up and shut-down phases as well as load reversals. The understanding of crack growth
behavior under creep-fatigue conditions is important.
The crack growth rate for CT-specimens in tests with hold times compared with tests without hold
times of creep-prestrained 13CrMo 44 was examined [16]. Hold times at maximum load caused
additional creep damage and a further increase in the crack propagation rate.
2.2.4.5 References
[1] NRIM Creep Data Sheet, No.1B, (1996).
[2] NIMS Creep Data Sheet, No.35B, (2002).
[3] The Iron and Steel Institute of Japan: Report on the Mechanical Properties of Metals at High
Temperatures, Vol.I Low Alloy Steel, (1972).
[4] Smith,G.V.: Evaluation of the Elevated Temperature Tensile and Creep-Rupture Prosperities of
1/2Cr-1/2Mo, 1Cr-1/2Mo, and 1-1/4Cr-1/2Mo-Si Steels, ASTM Data Series Publication DS 50,
ASTM, (1973).
[5] The British Steelmakers Creep Committee: BSCC high temperature data, (1972).
[6] Krisch, A.: Proc.of Joint Intern. Conf. on Creep, (1963), Inst. Mech.Eng., p.1-81.
[7] Afrouz, A., Collins, M.J., Pilkington, R.: Metals Techonology, 10 (1983), 461-463.
[8] Yukitoshi, T., Nishida, K.: Journal of the Society of Meterials Science, Japan, 21 (1972), 204-211.
[9] Kushima, H., Kimura, K., Abe, F., Yagi, K., Irie, H., Maruyama, K.: Tetsu-to-Hagane, 86 (2000)
131-137.
[10] Kushima, H., Kimura, K., Yagi, K., Tanaka, C., Maruyama, K.: Proc. of 7th JIM Int..Symp. on
Aspects of High Temperature Deformation and Fracture in Crystalline Materials, (1993), The Japan
Insititue of Metals, 609-616.
[11] Yukitoshi, T., Nishida, K.: Trans. ISIJ, 12 (1972), 429-434.
[12] Shinya. N., Yokoi, S., Kushima, H., Imai, Y.: Tetsu-to Hagane, 67 (1981), S439.
[13] Ishiguro, T., Tsukeda, T., Murakami, T.: Tetsu-to-Hagane, 69 (1983), S673.
[14] Wu, Rui, Storesund, J., Sandstrom, R.: Materials Sccience and Techonology, 9 (1993), 773-780.
[15] Browne, R. J., Flewitt, P. E., Lonsdale, D., Shammans, M. S., Soo, J. N.: Materials Science and
Techonology, 7 (1991), 707-717.
[16] Kullik, M., Maile, K.: Nuclear Engineering and Design, 199 (1990), 215-222.
Landolt-Börnstein
New Series VIII/2B
54
2.2 Low alloy steels
2.2.5 0.5Cr-0.5Mo-0.25V steel
2.2.5.1 Introduction
0.5Cr-0.5Mo-0.25V steel (14MoV6-3, 12MoCrV6-2-2) is primarily used as a steam pipe material in the
normalized and tempered heat treatment condition. The microstructure is typically tempered bainite, the
creep resistance being derived from a fine distribution of V4C3 precipitate.
The requirements for 0.5Cr-0.5Mo-0.25V steel as 14MoV6-3 in EN 10216 replace those for BS 3604
(Grade 660) and DIN 17 175 (14MoV6-3).
2.2.5.2 Material standards, chemical composition and tensile requirements
Table 31. Chemical requirements of 0.5Cr-0.5Mo-0.25V steel, 14MoV6-3 (EN 10216)
Chemical composition [wt%]
Stan- Std.
Desigdard No.
nation
C
Si
Mn
P
S
Cr
Mo
0.15 - 0.35 - 0.70 - ≤
0.60 - 0.70 ≤
EN
10216 14MoV6-3
0.10
0.15
0.40
0.50
0.025 0.020 0.30
V
0.28 0.22
Al
≤
0.040
The material is usually supplied in the normalized and tempered condition. The recommended
austenitizing temperature range is 930 - 990 °C with tempering in the range 680 - 750 °C.
Table 32. Room temperature mechanical property requirements for 0.5Cr-0.5Mo-0.25V steel, 14MoV63 (EN 10216)
StanStd.
DesigHeat
Thickness Rp0.2
Rm
A
Kv(RT)
dard
No.
nation
treat
[mm]
Nmm-2
[J]
[Nmm-2]
[%]
610 EN
10216 14MoV6-3 N+T
320
≤40
≥20
≥40
460
610 EN
10216 14MoV6-3 N+T
>40
310
≥20
≥40
460
Table 33. Minimum 0.2 % proof strength, Rp0.2, values at elevated temperatures for 0.5Cr-0.5Mo-0.25V
steel, 14MoV6-3 (EN 10216) (wall thicknesses ≤60 mm)
StanStd.
DesigHeat
Rp0.2 [Nmm-2] at a temperature [°C] of
dard
No.
nation
treat
100 150 200 250 300 350 400 450 500 550
EN
10216 14MoV6-3 N+T
282 276 267 241 225 216 209 203 200 197
2.2.5.3 Creep rupture strength
The creep rupture strength of 0.5Cr-0.5Mo-0.25V steel is shown in Fig. 61. The analysis from which the
data in the figure are derived was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found from their published data sheets [1].
The creep rupture properties have been obtained by analysis of (i) tube material with outer diameters
in the range 162 - 521 mm and thicknesses of 19 - 89 mm, (ii) forgings with sections sizes in the range
381 - 508 mm. The test data were from 46 heats with test temperatures of 475 - 740 °C. The distribution
of test durations is shown in Table 34.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 56]
2.2.5 0.5Cr-0.5Mo-0.25V steel
55
Table 34. Distribution of test durations used to derive the stress rupture properties of 14MoV6-3
(12MoCrV6-2-2).
Number of test points at the various test durations
10,000 20,001 30,001 50,001 70,001 >100,000 h
<10,000 h
20,000 h
30,000 h
50,000 h
70,000 h
100,000 h
331 (3)
81 (5)
30 (4)
29 (4)
11 (1)
13 (1)
5 (3)
( ) denotes unbroken tests
The data were assessed using the BS PD6605 procedure and the following master equation was derived:
ln(tu*) = β0 + β1log(σ) + β2σ + β3σ2 + β4T + β5/T
where tu* is the predicted rupture time in hours, T is the absolute temperature and σ is the stress in
Nmm-2.
Table 35. The βi are constants with the following values
β0
−39.7658730
β1
−8.43513298
β2
−0.00186616660
β3
−2.91037377×10−5
β4
0.00935613085
β5
49662.4102
STRENGTH, MPa
1000
100
10,000h
100,000h
200,000h
250,000h
Fig. 61. Creep rupture strength data of
14MoV6-3 (12MoCrV6-2-2).
10
400
450
500
550
600
o
TEMPERATURE, C
Landolt-Börnstein
New Series VIII/2B
650
56
2.2 Low alloy steels
2.2.5.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 61 the 100,000 h rupture strength values for a range of temperatures are
as follows:
100,000 h rupture strengths [Nmm-2] at specified temperatures [°C]
Temperature 450
460
470
480
490
500
510
520
Stress
305
276
249
224
200
177
155
135
100,000 h rupture strengths [Nmm-2] at specified temperatures [°C]
Temperature 530
540
550
560
570
580
590
600
Stress
117
102
87
75
65
56
48
41
2.2.5.5 Reference
[1] ECCC Data Sheet for 12MoCrV6-2-2, European Creep Collaborative Committee, BRITE EURAM
Thematic Network BET2-0509 “Weld Creep”, (1999).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 61]
2.2.6 1Cr-1Mo-V steel
57
2.2.6 1Cr-1Mo-V steel
2.2.6.1 Introduction
1Cr-1Mo-V steel forgings have been used as rotors for steam turbines during the application of high
temperature steam in the 1950s. The manufacturing technologies of 1Cr-1Mo-V steel forgings have been
investigated for turbine rotors of large size. The establishment of most suitable heat treatment conditions
was important in order to obtain excellent creep properties for this steel. Because the microstructure
changes during creep, these changes and the effecting factors on creep behavior have been studied. The
performance of turbine rotors was improved by the development of manufacturing technology for high
purity forgings.
2.2.6.2 Material standards, chemical and tensile requirements
Material standard for 1Cr-1Mo-V steel forgings for turbine rotors and shafts is ASTM A470-94a, class 8.
Table 36. Chemical requirements of 1Cr-1Mo-V steel forgings; ASTM A470-94a, class 8
Chemical composition [wt%]
C
Si
Mn
P
S
Ni
Cr
Mo
V
ASTM A470-94a 0.25-0.35 0.15-0.35 ≤1.00 ≤0.012 ≤0.015 ≤0.75 1.05-1.50 1.00-1.50 0.20-0.30
Standards
Designation
Table 37. Tensile properties at room temperature of 1Cr-1Mo-V steel forgings; ASTM A470-94a, class 8.
Tensile strength
Yield strength Elongation to radial body
Reduction of area to radial body
[MPa]
[MPa]
[%]
[%]
725 - 860
≥585
≥14
≥38
Table 38. Notch toughness requirements of 1Cr-1Mo-V steel forgings; ASTM A470-94a, class 8
Transition temperature FATT50 [°C]
Room temperature impact [J]
≤121
≥16
Tensile strength
1000
800
800
Stress (MPa)
Stress (MPa)
0.2% proof stress
1000
600
400
200
0
600
400
200
0
100
200
300
400
500
600
700
Test temperature (℃)
Fig. 62 Tensile properties of 1Cr-1Mo-V steel forgings [1].
Landolt-Börnstein
New Series VIII/2B
0
0
100
200
300
400
500
Test temperature (℃)
600
700
58
2.2 Low alloy steels
2.2.6.3 Creep properties of 1Cr-1Mo-V steel forgings
Information of fact on creep data for 1Cr-1Mo-V steel forgings can be obtained from [1] to [8].
2.2.6.3.1 Creep rupture data of 1Cr-1Mo-V steel forgings
Creep rupture strength data of 1Cr-1Mo-V steel forgings is shown in Fig. 63.
500
Stress [MPa]
300
100
80
60
40
10
450 °C
500 °C
525 °C
550 °C
575 °C
600 °C
625 °C
650 °C
675 °C
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 63. Creep rupture strength
data of 1Cr-1Mo-V steel forgings;
[1].
2.2.6.3.2 Creep rupture strength of 1Cr-1Mo-V steel forgings
The creep rupture data from Fig. 63 were analyzed using the Manson-Haferd parameter method. The
master rupture curve is shown in Fig. 64 [1].
800
600
500
400
Stress [MPa]
300
200
450 °C
500 °C
525 °C
550 °C
575 °C
600 °C
625 °C
650 °C
675 °C
100
80
60
50
Average
1)
n = 229 (238)
40
30
-2.0
-3.0
-2.5
-3.5
-4.0
Manson-Haferd parameter [( log tR -17.145)/( TK -370.0 )] [×10 - 2 ]
Fig. 64. Master rupture curve by Manson-Haferd
parameter method for 1Cr-1Mo-V steel forgings; [1].
1)
The number in parenthesis includes the estimated
values.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 61]
2.2.6 1Cr-1Mo-V steel
59
2.2.6.3.3 Effect of heat-treatment conditions on creep strength of 1Cr-Mo-V steel forgings
The creep strength is an important property for turbine rotors, and toughness and thermal fatigue strength
are also requested. The balance of these properties must be kept. Heat treatment conditions have been
investigated in order to obtain the most suitable properties. Oil-quenching and tempering was at first used
as heat treatment of rotor forgings, and air-cool normalizing and tempering was later introduced. High
strength was easily obtained by oil-quenching, but quenching crack often happened. On the other hand,
high strength could not be obtained by air-cooling. The heat treatment method using water sprays of fog
was developed, and made it possible to control any cooling condition between oil-cooling and air-cooling.
The turbine rotor with sound strength and toughness was manufactured using this method [2, 3]. A lot of
research concerning effects of heat treatment on creep rupture notch strength was carried out. One case is
shown in Fig. 65 [4].
100
80
70
A
Stress [1000 psi]
60
B
22
17
A
B
50
13
14
40
Cooling Smooth Notch
rate
A 300°F/h
B
400 °F/h
Single lines are smooth-bar data
Double lines are notch-bar data
Numbers are rupture elongations [%]
Rotor
30
20
1
10
9
10 3
10 2
Rupture time [h]
16
Fig. 65. Comparison of the 1000 °F stress-rupture
properties of rotors A and B. Austenitized at 1750 °F,
cooled as shown, and tempered at 1225 °F [4] 1 psi = 6.895
kPa.
10 4
2.2.6.3.4 Microstructural changes and long-term creep rupture properties for 1Cr-1Mo-V steel
forgings
Superior creep strength of 1Cr-1Mo-V steel forgings for turbine rotors is caused by precipitation
strengthening due to finely distributed carbides. The degradation of high temperature strength properties
is due to cohesion and enlargement of carbide particles, and caused by strength loss due to formation of
voids and cracks during creep.
Microstructural change is schematically presented in Fig. 66. For this steel, the cohesion and
enlargement of grain boundary carbide particles during creep are remarkable. The formation of subgrains
due to grain boundary migration is observed in the vicinity of enlarged grain boundary carbides, and low
density precipitation areas of fine carbide particles are also observed along grain boundaries. The creep
damage of this steel seems to be due to local recovery near grain boundaries, and not due to the formation
of voids and cracks.
The creep rupture ductility of 1Cr-1Mo-V steel is affected by impurity elements such as P, S, Sn, As,
Sb and so on. The lowering of long-term creep rupture ductility is caused by formation of voids due to
enlarged carbide particles at the grain boundaries and the linking of voids. The correlation between creep
rupture ductility and numbers of grain boundary cavities is shown in Fig. 67 [6].
Landolt-Börnstein
New Series VIII/2B
60
2.2 Low alloy steels
Fig. 66. Schematic representations of microstructual change
during creep; [5].
-5
-10
-15
80
600
500
400
60
300
40
Cavities [mm 2 ]
D D/D [×10 4 ]
0
Reduction in area [%]
5
Reduction in area
Density change, D D/D [×10 4 ]
Grain boundary cavities [mm 2 ]
100
200
-20
20
-25
0
100
10 2
10 4
10 3
Time to rupture [h]
0
10 5
Fig. 67. Correlation between rupture ductility, density
change and area of grain boundary cavities in specimens
tested at 550 °C; [6].
2.2.6.3.5 Creep crack growth property of 1Cr-1Mo-V steel forgings
The understanding of creep crack growth properties of 1Cr-1Mo-V steel is important. The effect of
specimen size on creep crack growth properties is shown in Fig. 68 [7]. The creep crack growth rate is
dependent on the thickness of the specimen: The larger the thickness the faster the rate. This is due to the
restriction of deformation to the direction of thickness at the crack tip.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 61]
2.2.6 1Cr-1Mo-V steel
61
10
d a /d t [mm/ h]
1Cr-Mo-V steel
538 °C
10- 1
W [mm] B [mm]
254.0 63.5
254.0 12.7
50.8 25.4
50.8 12.7
50.8 6.35
50.8 6.35
no S.G.
10- 2
10- 3
10- 1
1
10
10 2
Fig. 68. Effect of specimen thickness, B, on the relationship between creep crack growth rate da/dt and C*
parameter; [7].
C * [kJ/m2 h ]
2.2.6.4 References
[1]
[2]
[3]
[4]
[5]
[6]
[7]
[8]
National Research Institute for Metals: NRIM Creep Data Sheet, No.9B, (1990).
Sakabe, K., Hori, K., Honma, R.: Tetsu-to-Hagane, 46 (1960), 1340-1342.
Watanabe, J., Kumada, Y., Iwasaki, T.: Tetsu-to-Hagane, 52 (1966), 687-689.
Werner, F.E., Eichelberger, T.W., Hann, E.K.: Trans.Amer.Soc.Metals, 52 (1960), 376-403.
Matsuo, T., Kisanuki, T., Tanaka, R., Komatsu, S.: Tetsu-to-Hagane, 70 (1984), 565-572.
Shin-ya, N., and Keown, S.R.: Material Science. 13 (1979), 89-93.
Tabuchi, M., Kubo, K., and Yagi, K.: Engineering Fracture Mechanics, 40 (1991), 311-321.
The Iron and Steel Institute of Japan: Report on the Mechanical Properties of Metals at High
Temperatures, Vol. I, Low Alloy Steels, (1972).
Landolt-Börnstein
New Series VIII/2B
62
2.2 Low alloy steels
2.2.7 1.25Cr-0.5Mo steel
2.2.7.1 Introduction
2.2.7.1.1 1.25Cr-0.5Mo steel plates
This material is normalized and tempered and used as plates for boilers and pressure vessels. The steel
plate shall be made of the killed steel by hot rolling process. The steel plate shall be subjected to heat
treatment either annealing or normalizing and tempering so that the normal grain size may be obtained.
When high tensile strength is required, generally, the steel plate shall be subjected to normalizing and
tempering. The maximum thickness of 1.25Cr-0.5Mo steel plates is 200 mm.
2.2.7.1.2 1.25Cr-0.5Mo steel tubes
1.25Cr-0.5Mo steel is used for boiler and heat exchanger seamless tubes. It is applied in alloy steel tubes,
used for the purpose of heat exchange at the inside and outside of the tube, such as water tubes, smoke
tubes, superheater tubes and air preheater tubes of boilers, or heat exchanger tubes, condenser tubes and
catalyst tubes in the chemical and petroleum industries. However, it is not used for tubes in heating
furnaces and heat exchangers at low temperatures. These steel tubes shall be made by the seamless
process. Generally, tubes made of 1.25Cr-0.5Mo steel shall be subjected to heat treatments with
isothermal annealed, fully annealed or normalized and tempered conditions.
2.2.7.2 Material standards, chemical compositions and tensile properties
2.2.7.2.1 1.25Cr-0.5Mo steel plates
The following information was obtained from [1].
Table 39 shows the specification (SCMV 3 NT (JIS G 4109)) for the chemical composition of 1.25Cr0.5Mo steel plates and the analysis results of chemical compositions of typical three heats used for data
treatment in this article. Moreover, Table 40 shows the heat treatment history of these three heats. Here,
any test material has been normalized and tempered.
The normalizing temperature shall be within the range of 875 to 1000 °C. After heating to the
normalizing temperature, accelerated cooling such as liquid cooling, air-blasting or other appropriate
methods may be performed to obtain the specified mechanical properties. Here, air-blasting has been
applied to the test materials.
The tempering temperature shall be above 620 °C. For the test materials, 630 °C as one-step tempering
temperature or 710 °C and 700 °C as two-step tempering temperatures have been applied.
Furthermore, Table 41 shows the specification for the tensile properties at room temperature.
Table 39. Specification for chemical composition and analysis results for 1.25Cr-0.5Mo steel plates.
Chemical composition [wt%]
Standard/Heats
C
Si
Mn
P
S
Ni
Cr
Mo
Fe
SCMV 3 NT
0.44
0.36
0.940.40Rem.
≤0.17
≤0.03
≤0.03
(JIS G 4109)
0.86
0.69
1.56
0.70
Heat 1
0.15
0.61
0.6
0.009
0.013
0.35
1.2
0.51
Rem.
Heat 2
0.16
0.66
0.56
0.015
0.015
0.05
1.27
0.51
Rem.
Heat 3
0.13
0.52
0.54
0.009
0.011
0.05
1.1
0.48
Rem.
Rem. = Remainder.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 66]
2.2.7 1.25Cr-0.5Mo steel
63
Table 40. Heat treatment history for 1.25Cr-0.5Mo steel plates.
Thermal history
Heat 1
920 °C × 1.5 h A.C. 710 °C × 1.5 h A.C. 700 °C × 2 h F.C.
Heat 2
950 °C × 1.5 h A.C. 630 °C × 2.2 h A.C.
Heat 3
930 °C × 1.5 h A.C. 630 °C × 2.2 h A.C
Table 41. Specification of tensile properties for 1.25Cr-0.5Mo steel plates.
Tensile properties
Requirement
Tensile strength
0.2 % proof stress Elongation
[MPa]
[MPa]
[%]
SCMV 3 NT
520 - 690
315 min.
22 min.
(JIS G 4109)
Reduction of area
[%]
-
800
800
700
700
y strength [MPa]
0.2 % proof
Tensile strength [MPa]
Fig. 69a - Fig. 69d show tensile strength, 0.2% proof strength, tensile elongation and reduction of area of
1.25Cr-0.5Mo steel plates from room temperature to 650 °C. The solid lines express the average level of
the respective property.
600
500
400
300
200
600
500
400
300
200
100
100
0
0
0
100
200
300
400
500
600
0
700
100
200
Temperature [°C]
Fig. 69a. Tensile strength of 1.25Cr-0.5Mo steel plates.
400
500
600
700
Fig. 69b. 0.2% proof strength of 1.25Cr-0.5Mo steel
plates.
100
100
80
80
Reduction of area [%]
Elongation [%]
300
Temperature [°C]
60
40
60
40
20
20
0
0
0
100
200
300
400
500
600
700
Temperature [°C]
Fig. 69c. Tensile elongation of 1.25Cr-0.5Mo steel plate.
Landolt-Börnstein
New Series VIII/2B
0
100
200
300
400
500
600
700
Temperature [°C]
Fig. 69d. Reduction of area of 1.25Cr-0.5Mo steel
plates.
64
2.2 Low alloy steels
Although the strong decrease of tensile strength is not indicated till about 400 °C, tensile strength
decreases slightly near 100 °C and subsequently it increases slightly near 300 to 400 °C where it reaches a
similar level as at room temperature. However, for temperatures above 500 °C, tensile strength decreases
rapidly. 0.2% proof strength decreases gradually with increasing temperature above around 200 °C. For
temperatures above 550 °C, degression becomes large. The tensile elongation curve inversely correlates
with that of tensile strength. It increases strongly from near 500 °C. This tendency is the same also for the
reduction of area. It should be noted that the scatter of these tensile properties comes from different heats.
2.2.7.2.2 1.25Cr-0.5Mo steel tubes
The following information has been obtained [1].
Table 42 shows the specification (STBA23 (JIS G 3462)) for the chemical composition of 1.25Cr-0.5Mo
steel tubes and the analysis results of chemical compositions of typical three heats used for data treatment
in this article. Moreover, Table 43 shows the heat treatment history of these three heats. Here, any test
material has been normalized and tempered. Annealing needs to be performed above 650 °C and it was
carried out above 670 °C here. Furthermore, Table 44 shows the specification for the tensile properties at
room temperature.
Table 42. Specification for the chemical composition and analysis results for 1.25Cr-0.5Mo steel tubes.
Chemical composition [wt%]
Standard/Heats
C
Si
Mn
P
S
Ni
Cr
Mo
Fe
STBA 23
0.15
0.50 0.30 0.03
0.03
1.00 0.45 Rem.
(JIS G3462)
max.
1.00
0.60
max.
max.
1.50
0.65
Heat 1
0.12
0.79
0.49
0.007
0.006
0.12
1.2
0.56
Rem.
Heat 2
0.11
0.7
0.47
0.02
0.018
0.022
1.27
0.54
Rem.
Heat 3
0.12
0.69
0.45
0.021
0.017
0.024
1.28
0.55
Rem.
Table 43. Heat treatment history for 1.25Cr-0.5Mo steel tubes.
Thermal history
Heat 1
910 °C × 10 min → 670 °C × 70 min A.C.
Heat 2
920 °C × 60 min → 740 °C × 90 min A.C.
Heat 3
930 °C × 10 min → 690 °C × 70 min A.C.
Table 44. Specification of tensile properties for 1.25Cr-0.5Mo steel tubes.
Tensile properties
Requirement
Tensile strength
0.2% proof stress
Elongation
[MPa]
[MPa]
[%]
STBA 23
≥410
≥205
≥30
(JIS G3462)
Reduction of area
[%]
-
Fig. 70a - Fig. 70d show the tensile strength, 0.2% proof strength, tensile elongation and reduction of area
of 1.25Cr-0.5Mo steel tubes from room temperature to 650 °C. The solid line expresses the average level
of the properties.
Tensile strength, 0.2% proof stress and tensile elongation show very similar behavior to 1.25Cr-0.5Mo
steel plates (see above). Although the reduction of area tends to become a little lower near 300 °C, it
increases with temperature beyond 300 °C.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 66]
2.2.7 1.25Cr-0.5Mo steel
800
800
700
y strength [MPa]
0.2 % proof
700
600
Tensile strength [MPa]
65
500
400
300
200
100
600
500
400
300
200
100
0
0
0
100
200
300
400
500
600
700
0
100
200
Temperature [°C]
300
400
500
600
700
Temperature [°C]
Fig. 70a. Tensile strength of 1.25Cr-0.5Mo steel tubes.
100
Fig. 70b. 0.2% proof strength of 1.25Cr-0.5Mo steel
tubes.
100
80
Reduction of area [%]
Elongation [%]
80
60
40
20
60
40
20
0
0
100
200
300
400
500
600
700
Temperature [°C]
0
0
100
200
300
400
500
600
700
Temperature [°C]
Fig. 70c.
tubes.
Tensile elongation of 1.25Cr-0.5Mo steel
Fig. 70d. Reduction of area of 1.25Cr-0.5Mo steel tubes.
2.2.7.3 Creep rupture properties of 1.25Cr-0.5Mo steels
2.2.7.3.1 1.25Cr-0.5Mo steel plates
The creep rupture strength of 1.25Cr-0.5Mo steel plates is shown in Fig. 71. The creep tests were carried
out in the temperature range from 500 °C to 650 °C with a 50 °C pitch and at stresses between 37 MPa
and 431 MPa. The longest time to rupture is over 180,000 h. Even if there are more tests at low
temperature, test data scatter is larger at the low temperature side.
2.2.7.3.2 1.25Cr-0.5Mo steel tubes
The creep rupture strength of 1.25Cr-0.5Mo steel tubes is shown in Fig. 72. Creep tests were carried out
in the temperature range from 500 to 650 °C with a 50 °C pitch and at stresses between 41 MPa and 373
MPa. The longest time to rupture is over 60,000 h. Even if there are more tests at low temperature, test
data scatter is larger at the low temperature side.
Landolt-Börnstein
New Series VIII/2B
66
2.2 Low alloy steels
1.25Cr0.5Mo Steel Creep Rupture (Plates)
Stress (MPa)
1000
100
T=500℃
T=550℃
T=600℃
T=650℃
10
10-1
100
Fig. 71. Creep rupture strength
data of 1.25Cr-0.5Mo steel plates.
101
102
103
104
105
106
Time to rupture (h)
1.25Cr0.5Mo Steel Creep Rupture (Tubes)
Stress (MPa)
1000
100
T=500℃
T=550℃
T=600℃
T=650℃
10
10-1
100
101
Fig. 72. Creep rupture strength
data of 1.25Cr-0.5Mo steel tubes.
102
103
104
105
106
Time to rupture (h)
2.2.7.4 References
[1] NRIM Creep Data Sheet, No. 21B, (1994).
[2] NRIM Creep Data Sheet, No. 2A, (1976).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 73]
2.2.8 2.25Cr-1Mo steel
67
2.2.8 2.25Cr-1Mo steel
2.2.8.1 Introduction
2.25Cr-1Mo steel is widely used as tubes for boilers and heat exchangers and as components for pressure
vessels. It is heat-treated in order to obtain suitable strength properties for use.
2.2.8.2 Material standards, chemical and tensile requirements
2.2.8.2.1 2.25Cr-1Mo steel tubes for boilers and heat exchangers
Table 45. Chemical requirements of 2.25Cr-1Mo steel tubes; JIS STBA24, ASTM T22, BS622 and DIN
10CrMo 910
Standards
Designation
C
STBA24
JIS
≤0.15
ASTM T22
0.05-0.15
BS
622
0.08-0.15
10CrMo 910 0.08-0.15
DIN
Si
≤0.50
≤0.50
≤0.50
≤0.50
Chemical composition [wt%]
Mn
P
S
Cr
Mo
0.30-0.60 ≤0.030 ≤0.030 1.90-2.60 0.87-1.13
0.30-0.60 ≤0.025 ≤0.025 1.90-2.60 0.87-1.13
0.40-0.70 ≤0.040 ≤0.040 2.00-2.50 0.90-1.20
0.40-0.70 ≤0.035 ≤0.035 2.00-2.50 0.90-1.20
Ni
≤0.30
-
Al
≤0.020
-
Std.No.
G3462
A213
3606
17175
Table 46. Tensile properties at room temperature of 2.25Cr-1Mo steel tubes; JIS STBA24, ASTM T22,
BS622 and DIN 10CrMo 910
Tensile strength
Yield strength
Elongation
Standards
Designation
[MPa]
[MPa]
[%]
STBA24
JIS
≥410
≥205
≥30
ASTM
T22
≥415
≥205
490-640
BS
622
≥275
10CrMo 910
450-600
DIN
≥280
Tensile strength
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
600
300
300
200
200
100
100
0
0
100
200
300
400
500
600
700
0
0
Test temperature (℃)
Fig. 73. Tensile properties of 2.25Cr-1Mo steel tubes; JIS STBA24 [1].
Landolt-Börnstein
New Series VIII/2B
100
200
300
400
Test temperature
500
(℃)
600
700
68
2.2 Low alloy steels
2.2.8.2.2 2.25Cr-1Mo steel plates for boilers and pressure vessels
Table 47. Chemical requirements of 2.25Cr-1Mo steel plates; JIS SCMV4-2, ASTM Gr.22-cl.2 and EN
10CrMo 9-10
Standards
Chemical composition [wt%]
Std.No.
C
Si
Mn
P
S
Cr
Mo
≤0.17
≤0.50 0.30-0.60 ≤0.030 ≤0.030 2.00-2.50 0.90-1.10 G4109
0.05-0.15 ≤0.50 0.30-0.60 ≤0.035 ≤0.035 2.00-2.50 0.90-1.10 A387M
Designation
JIS
SCMV 4-2
ASTM Gr.22-cl.2
BS EN
10CrMo 9-10 0.08-0.14 ≤0.50 0.40-0.80 ≤0.030 ≤0.025 2.00-2.50 0.90-1.10 10028-2
DIN EN
Table 48. Tensile properties at room temperature of 2.25Cr-1Mo steel plates; JIS SCMV4-2, ASTM
Gr.22-cl.2 and EN 10CrMo 9-10.
Tensile strength
Yield strength
Elongation
Reduction of area
Standards Designation
[MPa]
[MPa]
[%]
[%]
520-690
JIS
SCMV 4-2
≥315
≥18
≥45
515-690
ASTM
Gr.22-cl.2
≥310
BS EN
10CrMo 9-10 480-630
≥310
DIN EN
Tensile strength
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
400
300
400
300
200
200
100
100
0
0
100
200
300
400
500
600
700
0
0
100
Test temperature (℃)
200
300
400
500
600
700
Test temperature (℃)
Fig. 74. Tensile properties of 2.25Cr-1Mo steel plates; JIS SCMV 4-2NT [2].
2.2.8.2.3 Quenched–and-tempered 2.25Cr-1Mo steel plates; ASTM A542, cl.1, 2, 3, 4 and 4a
Table 49. Chemical requirement of quenched-and-tempered 2.25Cr-1Mo steel plates; ASTM A542,
TypeA.
C
≤0.18
Si
≤0.50
Chemical composition according to product analysis [wt%]
Mn
P
S
Cr
Mo
Ni
0.25-0.66 ≤0.025 ≤0.025 1.88-2.62 0.85-1.15
≤0.43
Cu
≤0.43
V
≤0.04
Landolt-Börnstein
New Series VIII/2B
Ref. p. 73]
2.2.8 2.25Cr-1Mo steel
69
Table 50. Tensile properties at room temperature of quenched-and-tempered 2.25Cr-1Mo steel plates;
ASTM A542.
Tensile strength
Yield strength
Elongation
class
[MPa]
[MPa]
[%]
725 - 860
1
≥585
≥14
795 - 930
2
≥690
≥13
655 - 795
3
≥515
≥20
585 - 760
4
≥380
≥20
585 - 760
4a
≥415
≥18
Tensile strength
800
800
700
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
400
300
400
300
200
200
100
100
0
0
0
100
200
300
400
500
600
700
Test temperature (℃)
0
100
200
300
400
500
600
700
Test temperature (℃)
Fig. 75. Tensile properties of quenched-and-tempered 2.25Cr-1Mo steel plates; ASTM A542, cl.1, 3 and 4a; [3].
2.2.8.3 Creep properties of 2.25Cr-1Mo steels
Information of fact on creep data for 2.25Cr-1Mo steels can be obtained from [1], [2], [3], [9] and [10].
2.2.8.3.1 Creep rupture data of 2.25Cr-1Mo steels
Creep rupture data of annealed 2.25Cr-1Mo steel tubes, normalized-and-tempered 2.25Cr-1Mo steel
plates and quenched-and-tempered 2.25Cr-1Mo steel plates are shown in Fig. 76, Fig. 77 and Fig. 78,
respectively.
Landolt-Börnstein
New Series VIII/2B
70
2.2 Low alloy steels
500
450 °C
475 °C
500 °C
525 °C
550 °C
600 °C
650 °C
Stress [MPa]
300
100
80
60
40
20
10
10 2
10 3
10 4
Time to rupture [h]
10 5
Fig. 76. Creep rupture strength
data of 2.25Cr-1Mo steel tubes;
JIS STBA24; [1].
10 6
500
Stress [MPa]
300
100
80
450 °C
475 °C
500 °C
525 °C
550 °C
600 °C
650 °C
60
40
20
1
10 2
10
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 77. Creep rupture strength
data of 2.25Cr-1Mo steel plates;
JIS SCMV 4-2NT; [2].
700
500
Stress [MPa]
300
100
80
60
40
1
450 °C
475 °C
500 °C
525 °C
550 °C
575 °C
600 °C
650 °C
10
10 3
10 2
Time to rupture [h]
10 4
10 5
Fig. 78. Creep rupture strength
data of quenched - and - tempered
2.25Cr-1Mo steel plates; ASTM
A542; [3].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 73]
2.2.8 2.25Cr-1Mo steel
71
2.2.8.3.2 Creep rupture strength of 2.25Cr-1Mo steels
Creep rupture data of 2.25Cr-1Mo steel tubes was analyzed using the Manson-Haferd parameter method
[1]. The master rupture curve obtained is shown in Fig. 79.
500
400
450 °C
475 °C
500 °C
525 °C
550 °C
600 °C
650 °C
300
Stress [MPa]
200
100
80
60
50
40
30
Average
n = 294 (331)
20
-2.0
-3.0
-2.5
-3.5
-4.0
Manson-Haferd parameter [( log tR -16.053)/( TK -380 )] [×10 - 2 ]
Fig. 79. Master rupture curve by Manson - Haferd
parameter method for 2.25Cr-1Mo steel tubes; [1]. n indicates the total number of data points. The number in
parenthesis includes the estimated values.
2.2.8.3.3 Effect of heat treatment conditions on long-term creep rupture strength of 2.25Cr-1Mo
steel
The long-term creep rupture strength of 2.25Cr-1Mo steels subjected to various heat treatment conditions
is shown in Fig. 80. For shorter times, the creep rupture strength of quenched-and-tempered steel (Q.T.) is
higher than that of annealed (Ann.) and of normalized-and-tempered steel (N.T.), because short-term
creep rupture strength is dependent on tensile strength. However, long-term creep rupture strength is not
related to any heat treatment condition, because differences in microstructure disappear due to changes in
microstructure during creep [4].
10 3
2.25Cr-1Mo steels
Stress [MPa]
Ann.
N.T.
Q.T.
10 2
10
16000
Fig. 80. Creep rupture strength properties of 2.25Cr-1Mo
steels; [4].
18000
20000
22000
24000
Larson-Miller parameter T [K] (20+log t r [ h])
Landolt-Börnstein
New Series VIII/2B
72
2.2 Low alloy steels
2.2.8.3.4 Microstructual changes of 2.25Cr-1Mo steels
Various precipitates such as ε-carbide, Fe3C, Mo2C, M7C3, M23C6, M6C etc. are formed in 2.25Cr-1Mo
steels by heat treatment, during long heating, and during creep [5, 6]. Fig. 81 shows the carbide stability
of 2.25Cr-1Mo steel during tempering.
800
M23C6 + Cr7C3
M23C6 + M6C
Tempering temperature [°C ]
700
Fe3C + Cr7C3 + Mo2C
600
Fe3C + Mo2C
500
Fe3C
400
Fe3C + ε carbide
300
10 -1
1
2
10
10
Tempering time [h ]
10
3
4
Fig. 81. Carbide stability diagram of 2.25Cr-1Mo steel
by Baker and Nutting; [5].
10
The creep rupture curves of 2.25Cr-1Mo steel show inverse sigmoidal bending such as those of 1Cr0.5Mo steel. Fig. 82 shows the relationship between rupture life and minimum creep rate. It shows a large
data scatter at longer rupture lifes and lower minimum creep rates. This scatter can be explained by the
change in controlling factor of creep deformation and rupture life [7].
Time to rupture t r [h]
10 6
10
5
10
4
873 K
10 3
10 2
10
10 -6
2.25 Cr-1Mo steel
Experimental data
723 K
773 K
823 K
823 K
873 K
923 K
923 K
723 K
Predicted relations
( eI :30 %)
emin I vs. tr I
emin I vs. t r II
823 K
773 K
Fig. 82. Comparison of measured results with predicted
relations of time to rupture vs. minimum creep rate of
2.25Cr-1Mo steel; [7].
emin II vs. tr II
10 -5 10 -4 10 -3 10 -2 10 -1
Minimum creep rate, emin [% /h ]
1
2.2.8.3.5 Creep crack growth property of 2.25Cr-1Mo steels
The shape of crack tip for CT specimens of 2.25Cr-1Mo steel shows tunneled creep crack growth. The
electrical potential method for measuring tunneled crack growth was applied and a relationship between
crack growth rate and C* integral was obtained as shown Fig. 83 [8].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 73]
2.2.8 2.25Cr-1Mo steel
73
Creep crack growth rate d a /d t [mm /h]
1
10 -1
2.25Cr-1Mo steel
540 °C
CT specimen
(12.7, 6.4mmB)
10 -2
10 -3
da = 9.090×10- 3( C *) 0.818
dt
(correlation
coefficient = 0.956)
10 -4
10 -1
10
1
C *- Integral [kJ / m 2h ]
without SG
corrected by
max.final
crack length
SG
by Johnson’s eq.
10 2
Fig. 83. Relation between C*-integral and creep crack
growth rate modified by the maximum final crack
length for specimens without side grooves for 2.25Cr1Mo steel at 540 °C; [8]. SG means “side groove”.
2.2.8.4 References
[1]
[2]
[3]
[4]
[5]
[6]
[7]
[8]
[9]
[10]
NRIM Creep Data Sheet, No.3B, (1986).
NRIM Creep Data Sheet, No.11B, (1997).
NRIM Creep Data Sheet, No.36A, (1991).
Kushima, H., Kimura, K., Abe, F., Yagi, K., Irie, H., Maruyama, K.: Current Advances in Materials
and Processes, 9 (1996), 1310.
Baker, R.G., and Nutting, J.: The Iron and Steel Institute of Japan, 192 (1959), 257.
Thomson, R.C., and Bhadeshia, H.K.D.H.: Materials Science and Technology, 10 (1994), 193-203.
Yagi, K., Abe, F., Kimura, K., and Kushima, H.: Proc.of IUTAM Symposium on Creep in
Structures, Kluwer Academic Publishers, (2001), 267-276.
Fuji, A., Yamatani, I., Kitagawa, M., Ohtomo, A.: Tetsu-to-Hagane, 73 (1987), 1754-1761.
The Iron and Steel Institute of Japan: Report on the Mechanical Properties of Metals at High
Temperature, Vol.1, Low Alloy Steel, (1972).
The British Steelmakers Creep Committee: BSCC high temperature data, (1972).
Landolt-Börnstein
New Series VIII/2B
74
2.2 Low alloy steels
2.2.9 2.25Cr-1.6W-V-Nb steel
2.2.9.1 Introduction
2.25Cr-1.6W-V-Nb ferritic steel (T23, P23; HCM2S) is used as water wall, superheater and reheater
tubes, and header and main steam pipe in fossile fired boilers and heat recovery boilers. The steel has
been developed for improving creep rupture strength of 2.25Cr-1Mo steel at elevated temperatures mainly
by substituting Mo by W. The microstructure of the steel consists of a bainite matrix strengthened by
M23C6 carbides located mainly along grain boundaries and finely dispersed VC carbides in the matrix. VC
is fine and stable even after long term high temperature exposure.
2.2.9.2 Material standards, chemical and tensile requirements
Tables 51 and 52 give the chemical requirements and the corresponding tensile requirements of
2.25Cr-1.6W-V-Nb steel tubes and pipes designated by the following standards: Japanese
KA-STBA24J1, KA-STPA24J1, ASTM A213 T23, A335 P23, ASME Sec.I CC 2199.
Table 51. Chemical requirements of 2.25Cr-1.6W-V-Nb steel tubes and pipes;
Japanese KA-STBA24J1, KA-STPA24J1, ASTM A213 T23, A335 P23, ASME Sec.I CC 2199.
DesigGrade
nation
Japanese KA-STBA
METI
24J1
KA-STPA
24J1
ASTM- T23
A213
ASTM- P23
A335
C
Si
0.04 -
Mn P
0.10 -
S
-
Chemical composition [wt%]
Cr
Mo W
V
Nb N
1.90 0.05 1.45 0.20 0.02 -
Al
-
Std.
No.
B
0.0005
0.10 0.50 0.60 0.030 0.010 2.60 0.30 1.75 0.30 0.08 0.030 0.030 0.006
0.04 -
0.10 -
-
1.90 0.05 1.45 0.20 0.02 -
-
0.0005 ASME
Sec I
0.10 0.50 0.60 0.030 0.010 2.60 0.30 1.75 0.30 0.08 0.030 0.030 0.006 CC
2199
Table 52. Tensile requirements of 2.25Cr-1.6W-V-Nb steel tubes and pipes;
Japanese KA-STBA24J1, KA-STPA24J1, ASTM A213 T23, A335 P23, ASME Sec. I CC 2199.
Minimal
DesigStandard No.
Grade
TS1)
YS2)
elongation
nation
KA-STBA24J1
Japanese
510 MPa
400 MPa
20 %
METI
KA-STPA24J1
ASTM
T23
ASME Sec. I
A213
510 MPa
400 MPa
20 %
CC 2199
ASTM
P23
A335
1) TS: minimal tensile strength, 2) YS: minimal yield strength as 0.2% proof stress
2.2.9.3 Tensile properties of 2.25Cr-1.6W-V-Nb steel tubes
2.2.9.3.1 Tensile properties of 2.25Cr-1.6W-V-Nb steel tubes
Fig. 84 shows tensile strength and yield stress data of 2.25Cr-1.6W-V-Nb steel tubes [1]. Their values are
higher than those of T22 steel for all temperatures up to 650 °C. The corresponding tensile elongation and
reduction of area data of 2.25Cr-1.6W-V-Nb steel tubes are available in the literature [1].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 83]
2.2.9 2.25Cr-1.6W-V-Nb steel
75
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
0
100
200
300 400 500
Temperature [°C]
600
700
Fig. 84. Tensile strength (squares) and yield stress
(circles) of 2.25Cr-1.6W-V-Nb steel tubes.
2.2.9.3.2 Tensile properties of 2.25Cr-1.6W-V-Nb steel pipe
Fig. 85 shows tensile strength and yield stress data of 2.25Cr-1.6W-V-Nb steel pipes [1]. Their values are
higher than those of P22 steel at all temperatures up to 650 °C. The corresponding tensile elongation and
reduction of area data of 2.25Cr-1.6W-V-Nb steel pipes are available in [1].
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
Fig. 85. Tensile strength (squares) and yield stress
(circles) of 2.25Cr-1.6W-V-Nb steel pipes.
100
0
0
100
200
300 400 500
Temperature [°C]
600
700
2.2.9.4 Creep rupture properties of 2.25Cr-1.6W-V-Nb steel tubes and pipes
2.2.9.4.1 Creep rupture data of 2.25Cr-1.6W-V-Nb steel tubes
Fig. 86 shows creep rupture data of 2.25Cr-1.6W-V-Nb steel tubes. The longest creep rupture time of
2.25Cr-1.6W-V-Nb steel tubes is about 40000 h at 550 °C. Their long-term creep strengths are very stable
in the temperature range between 500 and 600 °C. Fig. 87 shows the Larson-Miller parameter plot of the
creep rupture data with a master rupture curve and a 95 % confidence lower limit [1]. The best fitting was
achieved with an optimized constant of 23.37.
Landolt-Börnstein
New Series VIII/2B
76
2.2 Low alloy steels
500
400
300
Stress [MPa]
500 °C
200
550 °C
100
80
500 °C
550 °C
600 °C
650 °C
600 °C
Average curve
60
× ruptured due to substantial oxidation of specimen
40
1
×
10 2
10 3
Rupture time [h]
10
650 °C
{
10 4
10 5
Fig. 86. Creep rupture strength
data of 2.25Cr-1.6W-V-Nb steel
tubes.
500
500 °C
×10 5h
400
Stress [MPa]
300
550 °C
×10 5h
200
100
80
60
40
17
600 °C
×10 5h
500 °C
550 °C
600 °C
650 °C
average curve
minimum curve
18 19 20 21 22 23 24 25 26 27
Larson-Miller-parameter T (23.37 + log t ) [×10 -3 ]
Fig. 87. Larson-Miller parameter plot of the creep
rupture data of 2.25Cr-1.6W-V-Nb steel tubes.
2.2.9.4.2 Creep rupture data of 2.25Cr-1.6W-V-Nb steel pipes
Fig. 88 shows creep rupture data of 2.25Cr-1.6W-V-Nb steel pipes. The longest creep rupture time of
2.25Cr-1.6W-V-Nb steel pipes is about 50000 h at 600 °C. Their long-term creep strengths are very stable
at temperatures between 500 and 600 °C. Fig. 89 shows the Larson-Miller parameter plot of the creep
rupture data with a master rupture curve and a 95 % confidence lower limit [1]. The best fitting was
achieved with an optimized constant of 19.95.
2.2.9.4.3 Creep data of 2.25Cr-1.6W-V-Nb steel tubes
Fig. 90 shows minimum creep rate data of 2.25Cr-1.6W-V-Nb steel tubes measured at various stress
levels at temperatures between 600 °C and 750 °C. Fig. 91 shows the Larson-Miller parameter plot of the
minimum creep rate data of 2.25Cr-1.6W-V-Nb steel tubes with a master minimum creep rate curve. The
best fitting was achieved with an optimized constant of 25.19.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 83]
2.2.9 2.25Cr-1.6W-V-Nb steel
77
500
400
Stress [MPa]
300
500 °C
200
550 °C
500 °C
550 °C
600 °C
650 °C
100
80
60
600 °C
average curve
650 °C
40
1
10 2
10 3
Rupture time [h]
10
10 4
10 5
Fig. 88. Creep rupture strength
data of 2.25Cr-1.6W-V-Nb steel
pipes.
500
400
500 °C
×10 5h
Stress [MPa]
300
550 °C
×10 5h
200
600 °C
×10 5h
100
80
60
40
15
625 °C
×10 5h
500 °C
550 °C
600 °C
650 °C
average curve
minimum curve
16 17 18 19 20 21 22 23 24 25
Larson-Miller-parameter T (19.95 + log t ) [×10 -3 ]
Fig. 89. Larson-Miller parameter plot of the creep
rupture data of 2.25Cr-1.6W-V-Nb steel pipes.
500
400
Stress [MPa]
300
200
500 °C
550 °C
100
80
500 °C
550 °C
600 °C
650 °C
average curve
600 °C
60
650 °C
40
10 -2
Landolt-Börnstein
New Series VIII/2B
10 -1
1
10
Minimum creep rate [% / 10 3 h]
10 2
Fig. 90. Minimum creep rate data
of 2.25Cr-1.6W-V-Nb steel tubes.
10
78
2.2 Low alloy steels
500
500 °C
0.01 % /10 3h
400
Stress [MPa]
300
550 °C
0.01 % /10 3h
200
100
500 °C
550 °C
600 °C
650 °C
80
60
40
average curve
17
18 19 20 21 22 23 24 25 26
.
Larson-Miller-parameter T (25.19 - log e ) [×10 -3 ]
Fig. 91. Larson-Miller parameter plot of the minimum
creep rate data of 2.25Cr-1.6W-V-Nb steel tubes.
2.2.9.4.4 Creep data of 2.25Cr-1.6W-V-Nb steel pipes
Fig. 92 shows minimum creep rate data of 2.25Cr-1.6W-V-Nb steel pipes measured at various stress
levels at temperatures between 600 °C and 750 °C . Fig. 93 shows the Larson-Miller parameter plot of the
minimum creep rate data of 2.25Cr-1.6W-V-Nb steel pipes with a master minimum creep rate curve [1].
The best fitting was achieved with an optimized constant of 31.27.
500
400
300
Stress [MPa]
500 °C
200
550 °C
100
600 °C
500 °C
550 °C
600 °C
650 °C
average curve
80 650 °C
60
40
10 -2
10 -1
1
10
Minimum creep rate [% / 10 3 h]
10 2
10 3
Fig. 92. Minimum creep rate data
of 2.25Cr-1.6W-V-Nb steel pipes.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 83]
2.2.9 2.25Cr-1.6W-V-Nb steel
79
500
400
500 °C
0.01 % /10 3h
Stress [MPa]
300
550 °C
0.01 % /10 3h
200
600 °C
0.01 % /10 3h
100
500 °C
550 °C
600 °C
650 °C
80
60
40
21
average curve
22 23 24 25 26 27 28 29 30 31
.
Larson-Miller-parameter T (31.27 - log e ) [×10 -3 ]
Fig. 93. Larson-Miller parameter plot of the minimum
creep rate data of 2.25Cr-1.6W-V-Nb steel pipes.
2.2.9.5 Allowable stress of 2.25Cr-1.6W-V-Nb steel tubes and pipes
Figs. 94 and 95 show the allowable tensile stresses determined for 2.25Cr-1.6W-V-Nb steel tubes and
pipes (Japanese METI KA-STBA24J1 and KA-STPA24J1) according to the METI standard procedure
comparing with those for the conventional steels ASME SA213-T22 and SA335-P22 (JIS STBA24 and
STPA24).
180
Allowable tensile stress (MPa)
160
KA-STBA24J1
140
120
100
STBA24
80
60
40
20
0
0
100 200 300 400 500 600 700 800
Temperature (℃)
Landolt-Börnstein
New Series VIII/2B
Fig. 94. Allowable tensile stress designated for 2.25Cr-1.6WV-Nb steel tubes (Japanese METI KA-STBA24J1).
80
2.2 Low alloy steels
180
Allowable tensile stress (MPa)
160
KA-STPA24J1
140
120
100
STPA24
80
60
40
20
Fig. 95. Allowable tensile stress designated for 2.25Cr1.6W-V-Nb steel pipes (Japanese METI KA-STPA24J1).
0
0
100 200 300 400 500 600 700 800
Temperature (℃)
2.2.9.6 Alloying philosophy of 2.25Cr-1.6W-V-Nb steel tubes
Fig. 96 shows the alloying philosophy of 2.25Cr-1.6W-V-Nb steel tubes. The steel has been developed to
improve the creep rupture strength of 2.25Cr-1Mo steel, which is mainly achieved by the combination of
solid solution of W and (V,Nb) C dispersion hardening in fully tempered bainitic matrix. Addition of B
enhances the bainitic microstructure and is found to improve the toughness of the steel to a great extent.
Low C content has been chosen to improve the weldability of the steel and as a result no preheating and
no post welding heat treatment (PWHT) is potentially required for some applications.
Weldability
Toughness
Creep strength
max.Hv≦350
Hardenability
: B addition
Solution Strengthening
Low C-0.06mass%
Fully tempered
bainitic structure
Precipitation
Strengthening
: V, Nb, B
: High W
No preheating and
PWHT after welding*
0.06C-2.25Cr-1.6W-0.1Mo-0.25V-0.05Nb-B
Matching welding
filler used without
PWHT
* according to the limitation of the spec.
Fig. 96. Alloying philosophy of 2.25Cr1.6W-V-Nb steel tubes.
2.2.9.7 Microstructural change of 2.25Cr-1.6W-V-Nb steel tubes
Fig. 97 shows a calculated phase diagram of 2.25Cr-1.6W-V-Nb steel at 600 °C with changing Cr and C
contents. The equilibrium phase diagram suggests that the final microstructure of 2.25Cr-1.6W-V-Nb
steel consists of ferrite (α) + MX ((V,Nb)C) + M6C and may include a small amount of M23C6.
Fig. 98 shows TEM micrographs of extraction replicas of 2.25Cr-1.6W-V-Nb steel normalized and
tempered (a) and crept for 12567.6 h at 600 °C (b). In the tempered specimen, M23C6 is formed along
prior austenitic grain boundaries and MX is formed along lath boundaries and in the bainitic matrix. M7C3
is occasionally observed along lath boundaries. In the crept specimen, on the other hand, no M23C6 and
M7C3 are observed and instead blocky M6C is formed along the prior austenitic grain boundaries and fine
MX remains along lath boundaries and in the bainitic matrix as shown in Fig. 99. It is noted that the
pronounced lath structure is kept even after long-term creep deformation.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 83]
2.2.9 2.25Cr-1.6W-V-Nb steel
81
α + M 23C 6+M 6C +M X
α + M X+M 6C
Fig. 97. Calculated phase diagram of 2.25Cr-1.6W-V-Nb
steel at 600 °C using Thermo-Calc.
M23C 6
MX
MX
M7C3
1μm
1μm
(a) normalized and tempered
M6C
MX
M6C
1μm
1μm
(b) crept for 12567.6 h at 600 °C
Fig. 98. TEM micrographs of extraction replicas of 2.25Cr-1.6W-V-Nb steel normalized and tempered (a) and crept
for 12567.6 h at 600 °C (b).
Landolt-Börnstein
New Series VIII/2B
82
2.2 Low alloy steels
Formation of M6C means that the steel looses W and/or Mo in solution, which are the key elements for
solid solution strengthening. It is, however, found that W is superior to Mo to retard the formation of M6C
during long-term creep deformation. Fig. 100 shows the change in precipitated W in 2.25Cr-1.6W0.2Mo-V-Nb steel and Mo in 2.25Cr-1.0Mo-V-Nb steel during aging for up to 10000 h at temperatures
between 550 °C and 650 °C. The lines indicate fitted curves assuming that the growth rate is controlled
by the Johnson-Mehrl-Abrami theory. It is seen that in 2.25Cr-1.6W-0.2Mo-V-Nb steel the growth rate of
M6C is 10 to 100 times slower that that in 2.25Cr-1.0Mo-V-Nb steel. This is one of the major reasons for
the high creep strength of 2.25Cr-1.6W-V-Nb steel.
020
000
110
Fig. 99. TEM micrograph showing fine
dispersion of MX in 2.25Cr-1.6W-V-Nb
steel crept for 12567.6 h at 600 °C.
200
B :001
100nm
1.0
1.0
Experiment value
0.8
Fraction of W - precipitation
Fraction of Mo - precipitation
Experiment value
650 °C
0.6
600 °C
0.4
550 °C
0.2
0
1
0.8
650 °C
0.6
600 °C
0.4
0.2
550 °C
0
10
10 2
10 3
Aging time [h]
(a) 2.25Cr-1.0Mo-V-Nb steel
10 4
10 5
1
10
10 2
10 3
Aging time [h]
10 4
10 5
(b) 2.25Cr-1.6W-0.2Mo-V-Nb steel
Fig. 100. Changes in precipitated W in 2.25Cr-1.6W-0.2Mo-V-Nb steel and Mo in 2.25Cr-1.0Mo-V-Nb steel during
aging for up to 10000 h at temperatures between 550 °C and 650 °C.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 83]
2.2.9 2.25Cr-1.6W-V-Nb steel
83
2.2.9.8 Performance of service exposed tubes
Detailed performance of service exposed tubes is available in [3], [6] and [7].
2.2.9.9 References
[1] Sumitomo seamless tubes and pipe Creep Data Sheets, Sumitomo Metal Industries, (1993).
[2] Masuyama, F., Yokoyama, T., Sawaragi, Y., and Iseda, A.: Materials for Advanced Power
Engineering, Part 1, Kluwer Academic Publishers (1994), 173.
[3] Sawaragi, Y., Kan, T., Yamadera, Y., Masuyama, F., Yokoyama, T., and Komai, N.: Proc. of the 6th
International Conference on Materials for Advanced Power Engineering, Liege Forschungszentrum
Jülich GmbH, 1998, p. 61.
[4] Komai, N., Masuyama, F., Ishihara, I., Yokoyama, T., Yamadera, Y., Okada, H., Miyata, K., and
Sawaragi, Y.: Advanced Heat Resistant Steels For Power Generation, The University Press,
Cambridge (1998), 96.
[5] Sawaragi, Y., Miyata, K., Yamamoto, S., Masuyama, F., Komai, N., Yokoyama, T.: Advanced Heat
Resistant Steels For Power Generation, The University Press, Cambridge (1998), 144.
[6] Miyata, K., Igarashi, M., and Sawaragi, Y.: ISIJ International, 39 (1999), 947.
[7] Miyata, K., and Sawaragi, Y.: ISIJ International, 41 (2001), 281.
Landolt-Börnstein
New Series VIII/2B
84
2.2 Low alloy steels
2.2.10 3Cr-0.5Mo-V steel
2.2.10.1 Introduction
3Cr-0.5Mo-V steel (20CrMoV13-5) is primarily used as a material for seamless circular tubes for use
mainly in chemical plants for high pressure applications at elevated service temperatures, which may also
include simultaneous exposure to hydrogen or hydrogen containing fluids (e.g. boiler or heat exchanger
tubes in petrochemical industry).
The microstructure is typically bi-phase bainite and ferrite. Its creep resistance is due to structural
stability and saturated bainite, the carbide distribution and solid solution hardening, which balances the
lower strength of the ferrite phase.
20 CrMoV13-5 is relatively easy to cold and hot bend (pipes/tubes) and to weld by typical welding
procedures for low alloy high temperature steels.
20CrMoV13-5 is suitable for use from 200 °C.
2.2.10.2 Material standards, chemical composition and tensile requirements
Table 53. Chemical requirements of 3Cr-0.5Mo-V steel (20CrMoV13-5).
Chemical composition [wt%]
DesigStanStd. No.
nation
dard
C
Si
Mn P
S
Cr
Mo Ni
Data
20CrMoV 0.17 0.15 0.30 ≤
3.00 0.50
≤
ECCC
sheet
13-5
0.23 0.35 0.50 0.020 0.020 3.30 0.60
20CrMoV 0.17 0.15 0.30 ≤
3.00 0.50 ≤
≤
EN
10216-2
13-5-5
0.23 0.35 0.50 0.025 0.020 3.30 0.60 0.30
20CrMoV 0.17 0.15 0.30 ≤
3.00 0.50
≤
DIN
17176
13-5
0.23 0.35 0.50 0.025 0.020 3.30 0.60
V
0.45
0.55
0.45
0.55
0.45
0.55
Al
≤
0.040
≤
0.040
≤
0.040
Cu
≤
0.30
-
Note: 20CrMoV13-5-5 EN 10216-2 supersedes 20CrMoV13-5 DIN 17176 specification.
The material is usually supplied in quenched and tempered condition. The recommended quenching
temperature range is 980 - 1030 °C with tempering in the range 680 - 730 °C, according to both of the
standards and to the ECCC data sheet.
Table 54. Room temperature mechanical property requirements for 3Cr-0.5Mo-V steel (20CrMoV13-5-5).
StanDesigHeat Thickness t
Rp0.2
Rm
A
Cv (RT)*
Std. No.
dard
nation
treat [mm]
[Nmm-2] [Nmm-2] %
[J]
Data
20CrMoV
740
ECCC
Q+T 1.6 < t < 100
590
sheet
13-5
880
740
20CrMoV
16
27
Q+T ≤100
590
EN
10216-2
880
13-5-5
740
20CrMoV
17
34
Q+T ≤100
590
DIN
17176
880
13-5
*The Charpy V-notch figure shown is in transverse direction.
Table 55. Minimum 0.2% proof strength values at elevated temperatures for 3Cr-0.5Mo-V steel
(20CrMoV13-5-5).
Thickness Heat
Rp0.2 [Nmm-2] at a temperature [°C] of
Standard Designation
[mm]
treat 100 150 200 250 300 350 400 450 500
EN
10216-2
≤ 60
Q+T 575 570 560 550 510 470 420 370
Landolt-Börnstein
New Series VIII/2B
Ref. p. 86]
2.2.10 3Cr-0.5Mo-V steel
85
2.2.10.3 Creep rupture strength
The creep rupture strength of 20CrMoV13-5-5 is shown in Fig. 101. The analysis, from which the data in
the figure were derived, was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found on their published data sheets [1].
20CrMoV13-5
Stress [MPa]
1000
100
100,000h
10,000h
Fig. 101. Creep rupture strength data of
3Cr-0.5Mo-V steel (20CrMoV13-5-5).
10
400
450
500
550
600
Temperature [°C]
The creep rupture properties have been obtained by comparative analysis of strength values as reported in
DIN 17176.
2.2.10.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 101 the 100,000 h rupture strength values for a range of temperatures are
as follows:
100,000 h rupture strength [Nmm-2] at specified temperatures [°C]
Temperature
420
430
440
450
460
470
Stress
420
370
310
260
220
190
480
165
100,000 h rupture strength [Nmm-2] at specified temperatures [°C]
Temperature
490
500
510
520
530
540
Stress
145
127
114
101
87
74
550
59
The tabled values do not include any extended extrapolation in time or stress by more than a factor of 3,
which is generally accepted as safe.
Landolt-Börnstein
New Series VIII/2B
86
2.2 Low alloy steels
2.2.10.5 References
[1] ECCC Data sheet for 20CrMoV13-5-5, European Creep Collaborative Committee, BRITE EURAM
Thematic Network BET2-0509 “Weld Creep”, 1999.
[2] DIN 17176, “Seamless Circular Steel Tubes for Hydrogen Service at Elevated Temperatures and
Pressures; Technical Delivery conditions”, 1990.
[3] EN 10216-2, “Seamless tubes for pressure purposes; Technical delivery conditions. Part 2: Nonalloy and alloy steel tubes with specified elevated temperature properties”, 2002
Landolt-Börnstein
New Series VIII/2B
Ref. p. 89]
2.2.11 5Cr-0.5Mo steel
87
2.2.11 5Cr-0.5Mo steel
2.2.11.1 Introduction
5Cr-0.5Mo steels are used as water tubes, smoke tubes, super-heater tubes, air preheater tubes and so on,
in boilers, heat exchanger tubes, condenser tubes, catalyst tubes and so on, as well as in chemical and
petrolic industries. Especially, 5Cr-0.5Mo steels are widely used in petrochemical industries, primarily
because of high strength and corrosion resistance against crude oils containing hydrogen sulfide and other
corrosive agents.
2.2.11.2 Material standards, chemical and tensile requirements
Table 56. Chemical requirements of 5Cr-0.5Mo steel tubes; JIS STBA 25, ASTM T5 and BS 625
Standards
Designation
C
JIS
STBA 25 ≤0.15
ASTM T5
≤0.15
BS
625
≤0.15
Si
≤0.50
≤0.50
≤0.50
Mn
0.30-0.60
0.30-0.60
0.30-0.60
Chemical composition [wt%]
P
S
Cr
≤0.030 ≤0.030 4.00-6.00
≤0.025 ≤0.025 4.00-6.00
≤0.030 ≤0.030 4.00-6.00
Mo
Ni
Al
0.45-0.65
0.44-0.65
0.45-0.65 ≤0.30 ≤0.020
Table 57. Tensile properties at room temperature of 5Cr-0.5Mo steel tubes; JIS STBA25.
Tensile strength
Yield strength
Elongation
[N/mm2]
[N/mm2]
[%]
d <10 mm
d ≥20 mm
10≤ d <20 mm
≥410
≥205
≥30
≥25
≥22
Tensile strength
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
400
300
400
300
200
200
100
100
0
0
100
200 300 400
500 600 700
800
Test temperature (℃)
Fig. 102. Tensile properties of 5Cr-0.5Mo steel tubes [1].
Landolt-Börnstein
New Series VIII/2B
0
0
100
200 300 400 500 600 700
Test temperature (℃)
800
88
2.2 Low alloy steels
2.2.11.3 Creep properties of 5Cr-0.5Mo steel
Information of fact on creep data for 5Cr-0.5Mo steel can be obtained from [1] and [3].
2.2.11.3.1 Creep rupture data of 5Cr-0.5Mo steel tubes
Creep rupture strength data of 5Cr-0.5Mo steel tubes is shown in Fig. 103.
Stress [MPa]
300
100
80
60
40
20
10
500 °C
550 °C
600 °C
650 °C
10 2
10 3
10 4
Time to rupture [h]
Fig. 103. Creep rupture strength
data of STBA 25; [1].
10 5
10 6
2.2.11.3.2 Creep rupture strength of 5Cr-0.5Mo steel tubes
It should be noted from Fig. 103 that creep rupture strength has a large scatter at higher stresses. The
creep rupture strength is dependent on manufacturing conditions, chemical composition, and initial
microstructure. This information is obtained from [1].
Creep rupture curves were analyzed using the Orr-Sherby-Dorn parameter method to NRIM creep
data. The master rupture curve is shown in Fig. 104.
400
500 °C
550 °C
600 °C
650 °C
300
Stress [MPa]
200
100
80
60
50
40
30
Average
n = 237
20
-22
-20
-18
-16
-24
Orr-Sherby-Dorn parameter log tR -[365457/(19.1425 × TK )]
Fig. 104. Master rupture curve for 5Cr-0.5Mo steel tubes.
n indicates the total number of data points.
Note TK : absolute Temperature [K]
T: test temperature [°C], and
tR: time to rupture.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 89]
2.2.11 5Cr-0.5Mo steel
89
2.2.11.3.3 Microstructural changes
Microstructural changes of long-term service-exposed process heater tube pipes of 5Cr-0.5Mo steels have
been studied in [2]. The samples selected for this study have been used in the temperature range of 450 °C
to 500 °C for about 20 to 25 years in oil refineries. The size, shape, position, distribution, and type of
carbides in virgin steel have been found to have changed significantly due to prolonged exposure of
220,000 h in the temperature range from 450 °C to 500 °C. The sequence of carbide precipitation in
5Cr-0.5Mo steels seems to be as follows:
M2C + M3C
→
M23C6
→
M23C6
→
M23C6
or
M3C
→
M7C3 + M3C
or
M2C + M3C + M7C3
2.2.11.4 References
[1]
[2]
[3]
National Research Institute for Metals: NRIM Creep Data Sheet, No.12B, (1992).
Das, S., Joarder, A.: Metallugical and Materials Transactions A, 28A (1997), 1607-1616.
The Iron and Steel Institute of Japan: Report on the Mechanical Properties of Metals at High
Temperatures, Vol. I Low Alloy Steels, (1972), pp.191-209.
Landolt-Börnstein
New Series VIII/2B
90
2.2 Low alloy steels
2.2.12 1.2Ni-Mo steel
2.2.12.1 Introduction
1.2Ni-Mo steel (15NiCuMo5-6-4, 9NiMoCuNb5-6-4) is primarily used as a steam pipe material for
pressure purposes (e.g. in boiler feedwater lines) and for steam drums and boiler tanks (mainly bottoms)
in the normalized and tempered heat treatment condition. The microstructure is typically tempered upper
bainite. Its creep resistance is derived from the structural stability of the over saturated bainite, the carbide
distribution and solid solution hardening.
1.2Ni-Mo steel is particularly suitable for use at temperatures around 300 °C due to enhanced ageing
resistance and stands prolonged exposure in creep regime with no creep strength loss compared to other
low alloyed structural steels. Practical experience nevertheless reported a tendency to embrittlement after
long service periods (more than 100000 h), particularly related to weldment heat affected zones.
2.2.12.2 Material standards, chemical composition and tensile requirements
Table 58. Chemical requirements of 1.2Ni-Mo steel, 9NiMoCuNb5-6-4, 15NiCuMo5-6-4
Standard
Std. No.
Data
ECCC
sheet
Designation
9NiMoCuNb
5-6-4
15NiCuMoNb
5-6-4
9NiMnMoCrNb
5-4-4
EN
10216-2
ISO
9329-2
BS
3604:PT1 591
Chemical composition [wt%]
C
≤
Si
P
0.25 0.80 ≤
0.17 0.50
0.25
≤
0.17 0.50
0.25
≤
0.17 0.50
0.10 0.25
0.17 0.50
Mn
1.20
0.80
1.20
0.80
1.20
0.80
1.20
0.030
≤
0.025
≤
0.030
≤
0.030
S
≤
Cr
≤
≤
≤
≤
≤
≤
≤
0.025 0.30
0.020 0.30
0.030 0.30
0.030 0.30
Mo
Ni
Nb
0.25
0.50
0.25
0.50
0.25
0.40
0.25
0.50
1.00
1.30
1.00
1.30
1.00
1.30
1.00
1.30
0.015
0.045
0.015
0.045
0.015
0.045
0.015
0.045
Al
≤
Cu
0.50
0.050 0.80
0.50
≤
0.050 0.80
0.50
≤
0.020 0.80
0.50
≤
0.045 0.80
Note: 15NiCuMoNb5-6-4 EN 10216-2 supersedes the BS specification for grade 591.
The material is usually supplied in normalized and tempered condition. According to the EN 10216-2 and
ISO 9329-2 standards and the ECCC data sheet, the recommended austenitizing temperature range is 880
- 980 °C with tempering in the range 580 - 680 °C. According to the BS 3604:PT1(1990) standard, the
austenitizing temperature is 900 - 980 °C with tempering at 580 - 660 °C.
The material could be also supplied in quenched and tempered condition. In this case the
recommended temperatures for heat treatment by the BS 3604:PT1(1990) standard are 880 - 930 °C for
quenching and 620 - 690 °C for tempering.
Table 59. Room temperature mechanical property requirements for 1.2Ni-Mo steel, 9NiMoCuNb5-6-4,
15NiCuMo5-6-4
StanHeat Thickness t
Rp0.2
Rm
A
Cv (RT)*
Std. No.
Designation
dard
treat [mm]
[Nmm-2] [Nmm-2] %
[J]
610
Data
9NiMoCuNb
N+T 1.6 < t < 100 ≥440
ECCC
780
Sheet
5-6-4
15NiCuMoNb
610
EN
10216-2
N+T ≤80
≥440
≥19
≥27
5-6-4
780
610
9NiMnMoCrNb
N+T ≤65
ISO
9329-2
≥440
≥19
≥27
780
5-4-4
610
BS
3604:PT1 591
N+T ≥440
≥20
760
610
BS
3604:PT1 591
Q+T ≥440
≥20
760
*The Charpy V-notch figure shown is in transverse direction.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 92]
2.2.12 1.2Ni-Mo steel
91
Table 60. Minimum 0.2% proof strength values at elevated temperatures for 1.2Ni-Mo steel,
9NiMoCuNb5-6-4, 15NiCuMo5-6-4
Minimum 0.2% proof strength, Rp0.2 in Nmm-2 at a
Thickness
Heat treat temperature in °C of
Standard Designation
[mm]
100 150 200 250 300 350 400 450
15NiCuMo
EN
N+T
422 412 402 392 382 373 343 304
≤80
5-6-4
2.2.12.3 Creep rupture strength
The creep rupture strength of 15NiCuMo5-6-4 is shown in Fig. 105. The analysis, from which the data in
the figure are derived, was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found in their published data sheets [1].
Stress [MPa]
1000
100
10,000h
100,000h
Fig. 105. Creep rupture strength
data of 15NiCuMo5-6-4.
10
350
400
450
500
550
Temperature [°C]
The creep rupture properties have been obtained by comparative analysis of strength values as reported in
VdTÜV Richtlinie 377/2. The test data used for the original analyses were related to several casts tested
at temperatures of 400 - 500 °C. The available maximum test durations and distribution of casts per
temperature are shown in Table 61.
Landolt-Börnstein
New Series VIII/2B
92
2.2 Low alloy steels
Table 61. Distribution of casts from different sources and maximum test duration available to derive the
stress rupture properties of 15NiCuMo5-6-4
Temperature [°C]
400
450
500
No. of casts Duration [h] No. of casts Duration [h] No. of casts Duration [h]
2
25,000
2
25,000
2
≤13,241
3
15,000
5
2
25,000
≤25,954
5
3
45,000
3
51,000
≤77,160
Due to the comparative analysis assessment method, no master equation is available.
2.2.12.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 105 the 100,000 h rupture strength values for a range of temperatures are
as follows:
100,000 h rupture strength at specified temperatures
Temperature [°C]
400
410
420
430
440
Stress
373
349
325
300
273
450
245
100,000 h rupture strength at specified temperatures
Temperature [°C]
460
470
480
490
500
Stress [Nmm-2]
210
175
139
104
69
-
The tabled values do not include any extended extrapolation in time or stress by more than a factor of 3,
which is generally accepted as safe.
2.2.12.5 NOTE
Due to its tendency to embrittlement after prolonged exposure, 15NiCuMo5-6-4 is nowadays often
replaced by 18NiCrMo2, which shows comparable high temperature strength, weldability and ageing
resistance, but does not exhibit significant loss in rupture elongation and impact strength after service.
2.2.12.6 Reference
[1] ECCC Data sheet for 9NiMoCuNb5-6-4, European Creep Collaborative Committee, BRITE
EURAM Thematic Network BET2-0509 “Weld Creep”, (1999).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 95]
2.2.13 1.4Cr-Mo steel
93
2.2.13 1.4Cr-Mo steel
2.2.13.1 Introduction
1.4Cr-Mo steel (Durehete 900) is used in fasteners for turbines and process plant and as boiler support
rods.
The steel was developed in the 1940’s as part of the Durehete series of creep resisting alloys. The
composition was based on an existing high strength engineering steel, En20B (0.4%C, 1.2%Cr, 0.7%Mo).
It derives its creep resistance [1] from a dispersion of Fe and Cr containing carbides, principally Fe3C and
Cr7C3, and is suitable for use in applications up to 480 °C (900 °F). During exposure at the service
temperature some Mo2C precipitates form which counteract reduction in strength due to coarsening of the
Cr based carbides.
2.2.13.2 Material standards and chemical composition and tensile requirements
Table 62. Chemical requirements of 1.4Cr-Mo steel, Durehete 900; EN 10269:1999
Chemical Composition [wt%]
StanStd. No.
Designation
dard
C
Si
Mn
P
S
Cr
0.39 0.40 1.20 ≤
≤
EN
10269:99 42CrMo5-6
≤0.40
0.45
0.70
1.50
0.035
0.035
Mo
0.50 0.70
The material is usually supplied in the oil quenched and tempered condition. The recommended
austenitizing temperature range is 840 - 870 °C with tempering in the range 600 - 700 °C.
Table 63. Room temperature mechanical property requirements for 1.4Cr-Mo steel, Durehete 900; EN
10269:1999.
StanHeat Diameter d
Rp0.2
A
Z
Cv(RT)
Rm
Std. No.
Designation
dard
treat [mm]
[%]
[J]
[MPa] [MPa] [%]
860
≤100
≥700
≥16
≥50
≥50
1060
EN
10269:99 42CrMo5-6 Q+T
850
100 <d ≤150 ≥640
≥16
≥50
≥40
1000
Table 64. Minimum 0.2% proof strength values at elevated temperatures for 1.4Cr-Mo steel, Durehete
900; BS EN 10269:1999
Minimum 0.2% proof strength, Rp0.2 [Nmm-2] at a temperature
DiaStan- DesigHeat
[°C] of
meter d
dard
nation
treat
[mm]
50 100 150 200 250 300 350 400 450 500 550 600
681+ 662 639 616 601 585 570 547 516 462 362 223
≤100
EN
42CrMo5-6 Q+T 100 <
625+ 605 584 563 549 535 521 500 472 422 331 204
d ≤150
+ values calculated by linear interpolation
2.2.13.3 Creep rupture strength
The creep rupture strength of 42CrMo5-6 is shown in Fig. 106. The analysis, from which the data in the
figure are derived, was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found from their published data sheets [2].
Landolt-Börnstein
New Series VIII/2B
94
2.2 Low alloy steels
10 3
8
6
10000 h
30000 h
100000 h
200000 h
Stress [MPa]
4
2
10 2
8
6×10
400
Fig. 106. Creep rupture strength data of 42CrMo5-6.
450
550
500
Temperature [°C]
600
The creep rupture properties have been obtained by analysis of bar material with sizes of 29 - 136 mm.
The tensile properties of the material covered the range 660 - 830 Nmm-2 for Rp0.2 and 855 - 1060 Nmm-2
for Rm. The test data were available from 26 heats with test temperatures of 450 - 550 °C. The distribution
of test durations is shown in Table 65.
Table 65. Distribution of test durations used to derive the stress rupture properties of 42CrMo5-6.
Number of test points at the various test durations
10,000 20,000 30,000 <10,000 h
20,000 h
30,000 h
50,000 h
151
18
9 (11)
4 (1)
( ) Denotes unbroken tests
The data were assessed using the ISO 6303 procedure and the following master equation was derived:
P(σ) = (log(tr*)−log(ta)) / (T−Ta)r = a + b(log σ) + c(log σ)2 + d(log σ)3 + e(log σ)4
where P(σ) is the creep rupture parameter, tr* is the predicted rupture time in hours, T is the absolute
temperature and σ is the stress in Nmm-2. The remaining symbols are constants with the following values:
log (ta)
Ta r
a
b
c
d
−2.4880708599 600 −1 −136047.7656 247130.3594 −164831.9531 48477.22266
e
−5323.73877
2.2.13.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 106 the 100,000 h rupture strengths for a range of temperatures are as
follows:
100,000 h rupture strengths at a range of temperatures
Temperature [°C] 450
460
470
480
490
Stress [Nmm-2]
410* 346* 276* 219* 176
500
148
510
124
520
102
* Values which have involved extended time extrapolation
Landolt-Börnstein
New Series VIII/2B
Ref. p. 95]
2.2.13 1.4Cr-Mo steel
95
2.2.13.5 References
[1] Everson, H., Orr, J., and Dulieu, D., “Low Alloy Ferritic Bolting Steels for Steam Turbine
Applications- The Evolution of the Durehete Steels”, ASM/EPRI Conference “Advances in Material
Technology for Fossil Power Plant”, Chicago, (1987), pp.375-383.
[2] ECCC Data Sheet for 42CrMo5-6, European Creep Collaborative Committee, BRITE EURAM
Thematic Network BET2-0509 “Weld Creep”, (1999).
Landolt-Börnstein
New Series VIII/2B
96
2.2 Low alloy steels
2.2.14 Cr-Mo-V-Ti-B steel
2.2.14.1 Introduction
Cr-Mo-V-Ti-B steel (Durehete 1055) is used in fasteners for turbines and process plants and as boiler
support rods. It is part of the Durehete series of creep resisting steels and was developed originally in the
1950’s in response to the increases in operating temperatures in steam power plants. The earlier alloy
Durehete 900 was a CrMo base. Improved high temperature properties were obtained by the addition of V
resulting in the 1%Cr, 1%Mo, 0.75%V steel, Durehete 950. Further improvements were obtained by
adjusting the carbon and vanadium to near stoichiometric levels maximizing the elevated temperature
strength to give the alloy Durehete 1050. However the high strength of this material was coupled with
poor ductility which resulted in grain boundary cracking after relatively short exposures at 550 - 565 °C.
Durehete 1055 was the result of further development to improve the ductility. Additions of Ti and B were
made resulting in a finer austenite grain size. Further enhancement of the ductility, and hence a decrease
in the susceptibility to grain boundary cracking, has been produced by control of residual levels
principally phosphorus, arsenic, tin, antimony and copper.
Durehete 1055 is suitable for use at temperatures up to 570 °C (approximately 1055 °F).
2.2.14.2 Material standard, chemical composition and mechanical property requirements
Table 66. Chemical requirements for Cr-Mo-V-Ti-B steel, Durehete 1055; EN 10269:1999
Chemical composition [wt%]
StanDesignation
dard
C
Si Mn P
S
Cr Mo Ni Al B
V
Ti As Sn Cu
20CrMoVTiB 0.17 ≤
0.35 ≤
0.90 0.90 ≤
0.015 0.001 0.60 0.07 ≤
≤
≤
≤
EN
0.23 0.40 0.75 0.020 0.020 1.20 1.10 0.20 0.080 0.010 0.80 0.15 0.020 0.020 0.20
4-10
The material is usually supplied in the oil or water quenched and tempered condition. The recommended
austenitizing temperature range is 970 - 990 °C with tempering in the range 680 - 720 °C.
Table 67. Room temperature mechanical property requirements for Cr-Mo-V-Ti-B steel, Durehete 1055;
EN 10269:1999
Stan- Std.
Heat Diameter d
Rp0.2
Rm
A
Z
Cv(RT)
Designation
dard No.
treat [mm]
[MPa] [MPa] [%]
[%]
[J]
820
≥
≥
≥
≥
≤100
1000
660
15
50
40
1026 20CrMoVTiB
EN
Q+T
9:99 4-10
820
≥
≥
≥
≥
100 < d ≤ 160
1000
660
15
50
27
Table 68. Minimum 0.2% proof strength values at elevated temperatures for Cr-Mo-V-Ti-B steel,
Durehete 1055; EN 10269:1999
Minimum 0.2% proof strength, Rp0.2 [Nmm-2] at a
Heat Diameter
StanDesignation
temperature [°C] of
treat [mm]
dard
50
20CrMoVTiB
Q+T ≤160
EN
4-10
+ values calculated by linear interpolation
100 150 200 250 300 350 400 450 500 550 600
642+ 624 603 595 581 573 559 537 508 464 406 334
Landolt-Börnstein
New Series VIII/2B
Ref. p. 99]
2.2.14 Cr-Mo-V-Ti-B steel
97
2.2.14.3 Creep rupture strength
The creep rupture strength of 20CrMoVTiB4-10 is shown in Fig. 107. The analysis, from which the data
in the figure are derived, was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found from their published data sheets [1].
10 3
8
10000 h
30000 h
100000 h
200000 h
6
Stress [MPa]
4
2
10 2
8
6×10
400
450
550
500
Temperature [°C]
600
650
Fig. 107. Creep rupture strength data of 20CrMoVTiB410
The creep rupture properties have been obtained by analysis of bar material with sizes of 19 - 190 mm.
The tensile properties of the material covered the range 680 - 900 Nmm-2 for Rp0.2 and 830 - 1040 Nmm-2
for Rm. The test data were from 75 heats with test temperatures of 450 - 700 °C with the majority in the
range 500 - 600 °C. The distribution of test durations are shown in Table 69.
Table 69. Distribution of test durations used to derive the stress rupture properties of 20CrMoVTiB4-10
Number of test points at the various test durations
10,00020,00030,00050,00070,000>100,000 h
<10,000 h
20,000 h
30,000 h
50,000 h
70,000 h
100,000 h
513 (26)
81 (11)
94 (11)
44 (8)
15 (2)
1 (2)
8 (2)
( ) Denotes unbroken tests
The data were assessed using the ISO 6303 procedure and the following master equation was derived:
P(σ ) = (log(tr*)−log(ta)) / (T−Ta)r = a + b(logσ) + c(logσ)2 + d(logσ)3 + e(logσ)4
where P(σ) is the creep rupture parameter, tr* is the predicted rupture time in hours, T is the absolute
temperature and σ is the stress in Nmm-2. The remaining symbols are constants with the following values:
log (ta)
Ta r
a
b
10.523564339 590 1 −4.465617180 8.252388000
c
d
e
−5.727178097 1.762619853 −0.2033182085
2.2.14.4 Creep properties from research papers
At the sizes found in typical bolting applications after oil quenching from temperatures in the range 970 990 °C the microstructure of Durehete 1055 consists of fine grained bainite. Subsequent tempering in the
range 680 - 720 °C produces a dispersion of V4C3 precipitates. The alloy derives its creep strength from
these precipitates together with solid solution strengthening from chromium and molybdenum. The
optimum creep properties are obtained by a vanadium addition of about 0.7%, which is close to the
stoichiometric level for the carbon level of 0.2% in this alloy.
Landolt-Börnstein
New Series VIII/2B
98
2.2 Low alloy steels
However the maximum creep life is associated with a minimum in the tensile ductility, Fig. 108. This is
the result of regions close to grain boundaries which are denuded of precipitates where creep strain is
concentrated. Additions of titanium and boron are made to alleviate this problem. The titanium leads to
the formation of TiC precipitates in the proximity of the prior austenite grain boundaries although there is
evidence that these are replaced by V4C3 particles during long term exposure at 550 - 600 °C.
100
10
20
10
1
Elongation [%]
Rupture life [h]
Tested at
600 °C / 324 MPa
0
0
0.2
0.8
0.4
0.6
Vanadium - content [%]
1.0
1.2
Fig. 108. Influence of vanadium level on rupture life (solid
line) and ductility (dashed line).
Further improvements in ductility were produced by controlling the levels of residual elements
phosphorus, arsenic, tin, antimony and copper. The total residual content is described by the “R” factor
given by:
R = %P + 2.43 (%As) + 3.57 (%Sn) + 8.16 (%Sb) + 0.13 (%Cu)
Elongation [%] (20000 h / 550 °C )
40
30
20
10
0
0.05
0.10
‘R’Value
Fig. 109. Influence of residual content on rupture ductility.
0.15
The influence of the R value on the rupture ductility is illustrated in Fig. 109. There is a sharp increase in
rupture elongation as the R value decreases, particularly at levels below 0.1. Detailed metallography has
shown the presence of phosphorus and tin at grain boundaries.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 99]
2.2.14 Cr-Mo-V-Ti-B steel
99
2.2.14.5 Estimated long term creep rupture strength
Based on the data shown in Fig. 107 the 100,000 h rupture strengths for a range of temperatures are as
follows:
100,000 h rupture strengths [Nmm-2] at a range of temperatures [°C]
Temperature 450 460 470 480 490 500 510 520 530 540 550
Stress
453 423 394 365 337 307 276 241 204 169 142
560
121
570
103
2.2.14.6 References
[1] ECCC Data Sheet for 20CrMoVTiB4-10, European Creep Collaborative Committee, BRITE
EURAM Thematic Network BET2-0509 “Weld Creep”, (1999).
[2] Everson, H., Orr, J., and Dulieu, D. “Low Alloy Ferritic Bolting Steels for Steam Turbine
Applications- The Evolution of the Durehete Steels”, ASM/EPRI Conference “Advances in Material
Technology for Fossil Power Plant”, Chicago, (1987), pp.375-383.
Landolt-Börnstein
New Series VIII/2B
100
2.2 Low alloy steels
2.2.15 Cr-Mo-V steel
2.2.15.1 Introduction
Cr-Mo-V steel (21CrMoV5-7) is a frequently used creep resistant steel for fasteners and other parts in
power plants. The steel was developed in the 1970´s based on the older steel grade 21CrMoV5-11 when
long-term creep rupture tests showed that a Mo content of 0.7% is sufficient.
The Ni content was limited to max. 0.6% for 21CrMoV5-7 to avoid long term brittlement. For
dimensions >160 mm the through hardening could be improved by alloying a higher content of Ni (up to
0.8%, steel grade 21CrMoNiV4-7)
The specified upper limit of tensile strength and solution temperature should not be exceeded in order
to avoid notch weakness. Oil quenching or cooling in air after hardening is preferred to obtain appropriate
toughness properties at room temperature combined with good creep rupture strength values and
sufficient long term ductility.
Typical microstructure: Mainly tempered bainite with parts of tempered martensite for high cooling rates
at the grain boundary or parts of ferrite for low cooling rates in the grain centre.
High temperature applications: Bolts, nuts, small forgings and other parts of steam turbines, compressors,
gas turbines, valves, nozzles.
Service temperature is up to 566 °C if oxidation is negligible or with surface protection, otherwise the
limiting temperature is around 540 °C
2.2.15.2 Material standards, chemical composition and tensile requirements
Table 70. Chemical requirements of 21CrMoV5-7
Chemical composition [wt%]
Stan- Std.
Designation
dard No.
C
Si
Mn P
S
Al
Cr
Mo
21CrMoV5-7
0.17 ≤
0.40 ≤
1.20 0.55
≤
≤
EN
10269
(1.7709)
0.25 0.40 0.80 0.030 0.030 0.030 1.50 0.80
21CrMoNiV4-7 0.17 0.15 0.35 ≤
0.90 0.65
≤
(1.6981)
0.25 0.35 0.85 0.030 0.035
1.20 0.80
21CrMoV5-11
0.17 0.30 0.30 ≤
1.20 1.00
≤
≤
(1.8070)
0.25 0.60 0.60 0.035 0.035 0.030 1.50 1.20
Ni
≤
0.60
0.20
0.80
≤
0.60
V
0.20
0.35
0.25
0.35
0.25
0.35
The material is usually supplied in quenched, in air or oil and tempered condition. The recommended
austenitizing temperature range is 900 - 950 °C with tempering treatment temperature in the range 680 720 °C.
Melting process: Electric arc.
Forming process: Hot rolled or forged.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 103]
2.2.15 Cr-Mo-V steel
101
Table 71. Room temperature mechanical property requirements for 21CrMoV5-7
Heat Thickness Rp0.2
Std.
Rm
Designation
Standard
treat
[mm]
No
[Nmm-2] [Nmm-2]
700
21CrMoV5-7
Q+T ≤160
EN
10269
≥550
850
(1.7709)
700
21CrMoNiV4-7
Q+T ≤600
≥550
850
(1.6981)
21CrMoV5-11
700
Q+T ≤250
≥550
(1.8070)
850
A
[%]
16
16
16
1000
1000
800
800
Tensile strength (MPa)
0.2% proof strength (MPa)
Table 72. Minimum 0.2% proof strength values at elevated temperatures for 21CrMoV5-7
Minimum 0.2% proof strength, Rp0.2 [Nmm-2] at a
Std.
Heat
Standard
Designation
temperature [°C] of
No
treat
50 100 150 200 250 300 350 400 450 500
21CrMoV5-7
EN
10269
Q+T 542 530 515 500 480 460 435 410 380 350
(1.7709)
600
400
type 21CrMoV5-7
type 21CrMoNiV4-7
200
type 21CrMoV5-11
600
400
type 21CrMoV5-7
type 21CrMoNiV4-7
type 21CrMoV5-11
200
min value EN10269
0
min value EN10269
0
0
100
200
300
400
500
600
0
100
200
Temperature (°C)
300
400
500
600
Temperature (°C)
Fig. 110. Tensile properties, tests carried out by the German Creep Committee; [1].
2.2.15.3 Creep rupture strength
The ECCC [2] has made a data collection and assessment showing that the data of all the three steel
grades were situated within the same acceptable scatter bands. The assessment was carried out using the
DESA procedure [3].
The following master equation was derived:
Larson-Miller [4] Parameter
Model function
PLM = (lg(t) + C ) τ
τ = (δ + 273) / 1000
lg(t) = −C + B1/τ + (B2/τ)√σ + (B3/τ)(√σ)2 + (B4/τ)(√σ)3
where t = test duration in h, σ = stress in MPa, δ = temperature in °C
Values for the constants:
C
20.00
Landolt-Börnstein
New Series VIII/2B
B1
22.349564
B2
−0.195449
B3
0.002218
B4
−0.000251
102
2.2 Low alloy steels
The equation is valid in the range of 420 °C to 550 °C and 100 h to 200,000 h. The results of the
assessment were published in ECCC Data Sheets in 1999 [2]. The creep rupture strength values for
420 °C to 550 °C and for 10 000 h, 100 000 h and 200 000 h are given in EN 10269 [5].
21CrMoV5-7
Stress [MPa]
1000
100
10,000h
30,000h
100,000h
200,000h
Fig. 111. Creep rupture strength data of
21CrMoV 5-7.
10
400
450
500
550
600
Temperature [°C]
The test data were available from 33 heats with test temperatures from 400 °C to 560 °C. The distribution
of test durations is shown in Table 73.
Table 73. Distribution of test durations used to derive the stress rupture properties of 21CrMoV 5-7.
Number of test points at the various test durations
10,000 20,001 - 30,001 - 50,001 70,001 <10,000 h
>100,000 h
20,000 h 30,000 h 50,000 h 70,000 h 100,000 h
190 (6)
38 (11)
16 (5)
27 (5)
15 (5)
7 (2)
7 (3)
( ) denotes unbroken tests
Landolt-Börnstein
New Series VIII/2B
Ref. p. 103]
2.2.15 Cr-Mo-V steel
103
2.2.15.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 111 and on the ECCC data sheet [2] the 100,000 h rupture strength
values and the 1% creep strength values for a range of temperatures are as follows:
Temperature [°C]
Stress [Nmm-2]
100,000 h rupture strengths at specified temperatures
420
430
440
450
460
399
375
350
325
300
470
274
480
249
Temperature [°C]
Stress [Nmm-2]
100,000 h rupture strengths at specified temperatures
490
500
510
520
530
224
199
174
150
126
540
103
550
82
Temperature [°C]
Stress [Nmm-2]
100,000 h 1% creep strengths at specified temperatures
420
430
440
450
460
470
365
340
315
288
262
235
480
208
Temperature [°C]
Stress [Nmm-2]
100,000 h 1% creep strengths at specified temperatures
490
500
510
520
530
540
182
156
132
109
89
71
550
56
2.2.15.5 References
[1] German Creep Commitee, VDEh, Düsseldorf.
[2] European Creep Collaborative Commitee, BRITE-EURAM Thematic Network BET2-0509
"WELD-CREEP", (1999).
[3] Granacher, J., and Oehl, M.: DESA Time-Temperature-Parameter Evaluation Method, Institut fur
Werkstoffkunde, Darmstadt (1993).
[4] Larson, F.R., and Miller, J.: A Time-Temperature Relationship for Rupture and Creep Stresses,
Trans ASME 74 (1952).
[5] EN 10269, Steels and nickel alloys for fasteners with specified elevated and/or low temperature
properties, (1999).
Landolt-Börnstein
New Series VIII/2B
104
2.2 Low alloy steels
2.2.16 0.5Mo cast
2.2.16.1 Introduction
The cast steel grade G20Mo5 (EN 10213-2, material-no 1.5419) represents the oldest type of the low
alloy chromium free creep resistant cast steels. Due to the molybdenum content of 0.40 to 0.60 % the
creep rupture strength values of G20Mo5 exceed those of the unalloyed cast steel grades. Industrial
applications are reported since about 1920. The main features of G20Mo5 are summarized below:
Melting process: Electric arc, induction melting.
Heat treatment: Quenched in air or liquid medium, and tempered.
Typical microstructure: Ferrite, tempered bainite and a small amount of pearlite.
Weldability: Easily weldable with similar weld metal.
High temperature applications: Casings of steam turbines, compressors, gas turbines, valves, nozzles;
service temperatures up to about 500 °C.
Cast steel grade with similar chemical composition: ASTM A 356 Grade 2.
2.2.16.2 Standard requirements
Table 74. Chemical composition
Chemical composition [wt %]
C
Si
Mn
P
S
Mo
EN10213-2:1995 G20Mo5 (1.5419) 0.15 - 0.23 ≤0.60 0.50 - 1.00 ≤0.025 ≤0.020 (1) 0.40 - 0.60
(1) The maximum admissible sulphur content is 0.030 % if the relevant wall thickness is not in excess of
28 mm.
Standard
Designation
Table 75. Heat treatment and tensile properties at room temperature.
Min. 0.2 %
Thickness
Standard Designation Heat treatment
proof strength
[mm]
[MPa]
EN
G20Mo5
Q:920°C-980°C
100
245
10213(1.5419)
T:650°C-730°C
2:1995
Tensile
strength
[MPa]
Min. elongation
at rupture
[%]
440-590
22
Landolt-Börnstein
New Series VIII/2B
Ref. p. 106]
2.2.16 0.5Mo cast
Rp0.2
Rm
min EN
700
500
600
400
Rm (MPa)
R p0.2(MPa)
105
300
200
100
500
400
300
200
100
0
0
100
200
300
400
500
0
600
0
100
Temperature (°C)
200
300
400
500
600
Temperature (°C)
Fig. 112. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade G20Mo5 tested in creep rupture tests
by the German Creep Committee; [1]. min EN: minimum values by EN 10213-2.
Stress (MPa)
1000
broken
unbroken
450 °C_EN
100
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 113. Creep rupture strength data of cast steel grade G20Mo5 at 450 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
500 °C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 114. Creep rupture strength data of cast steel grade G20Mo5 at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
106
2.2 Low alloy steels
Stress (MPa)
1000
broken
100
unbroken
550°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 115. Creep rupture strength data of cast steel grade G20Mo5 at 550 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
2.2.16.3 Average creep rupture strength
Table 76. Average creep rupture strength values indicated in EN 10213-2:1995:
Time to rupture
Temperature
10,000 h
100,000 h
200,000 h
[°C]
Average creep rupture strength [MPa]
400
360
310
290
450
275
205
180
500
160
85
70
550
66
30
23
2.2.16.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade G20Mo5, compilation of test results; Forschungsvereinigung Warmfeste
Stähle, c. o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 109]
2.2.17 1.5Cr-0.5Mo cast
107
2.2.17 1.5Cr-0.5Mo cast
2.2.17.1 Introduction
The cast steel grade G17CrMo5-5 (EN 10213-2, material-no 1.7357) was introduced in the 1940s. It
represents one of the low alloy creep resistant cast steels with improved resistance to pressurized
hydrogen. Because of the chromium content of 1.0 to 1.5 % the creep rupture strength values and yield
strength values at elevated temperatures are higher than those of the cast steel type G20Mo5 without
chromium addition. The main features of G17CrMo5-5 are summarized below:
Melting process: Electric arc, induction melting.
Heat treatment: Quenched in air and tempered (cooling in furnace).
Typical microstructure: Tempered bainite and ferrite.
Weldability: Easily weldable with similar weld metal.
High temperature applications: Casings of steam turbines, compressors, gas turbines, valves, nozzles,
chemical reactors; service temperatures up to about 530 °C.
Cast steel grade with similar chemical composition: ASTM A 217 Grade WC6.
2.2.17.2 Standard requirements
Table 77. Chemical composition.
Chemical composition [wt%]
C
Si
Mn
P
S
Cr
Mo
0.45EN10213G17CrMo5-5
0.150.50(1) 1.00≤0.60
≤0.020 ≤0.020
1.50
0.65
2:1995
(1.7357)
0.20
1.00
(1) The maximum admissible sulphur content is 0.030 % if the relevant wall thickness is not in excess of
28 mm.
Standard
Designation
Table 78. Heat treatment and tensile properties at room temperature.
Min. 0.2 %
Thickness
proof strength
Standard Designation
Heat treatment
[mm]
[MPa]
EN10213- G17CrMo5-5 Q: 920 °C - 960 °C
100
315
2:1995
(1.7357)
T: 680 °C - 730 °C
Landolt-Börnstein
New Series VIII/2B
Tensile
strength
[MPa]
Min. elongation at rupture
[%]
490-690
20
108
2.2 Low alloy steels
Rp0.2
min EN
Rm
600
800
400
Rm (MPa)
R p0.2 (MPa)
500
300
200
100
0
600
400
200
0
0
100
200
300
400
500
600
0
100
Temperature (°C)
200
300
400
500
600
Temperature (°C)
Fig. 116. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade G17CrMo5-5 tested in creep rupture
tests by the German Creep Committee [1]. min EN: minimum values by EN 10213-2.
Stress (MPa)
1000
broken
unbroken
400°C_EN
100
100
1000
10000
100000
1000000
Test duration (h)
Fig. 117. Creep rupture strength data of cast steel grade G17CrMo5-5 at 400 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
unbroken
450°C_EN
100
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 118. Creep rupture strength data of cast steel grade G17CrMo5-5 at 450 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 109]
2.2.17 1.5Cr-0.5Mo cast
109
Stress (MPa)
1000
broken
100
unbroken
500°C_EN
10
1
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 119. Creep rupture strength data of cast steel grade G17CrMo5-5 at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
2.2.17.3 Average creep rupture strength
Table 79. Average creep rupture strength values indicated in EN 10213-2:1995.
Time to rupture
Temperature
10,000 h
100,000 h
200,000 h
Average creep rupture strength [MPa]
400
420
370
356
450
321
244
222
500
187
117
96
550
98
55
44
2.2.17.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade G17CrMo5-5, compilation of test results; Forschungsvereinigung Warmfeste
Stähle, c. o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
110
2.2 Low alloy steels
2.2.18 2.25Cr-1Mo cast
2.2.18.1 Introduction
The low alloy cast steel grade G17CrMo9-10 (EN 10213-2, material-no 1.7379) with improved resistance
to high pressure hydrogen was introduced in the 1940s. Compared with G17CrMo5-5 the creep rupture
strength values are improved above 450 °C by the higher content in molybdenum and chromium.
Moreover the higher chromium content provides better resistance to oxidation. The main features of
G17CrMo9-10 are summarized below:
Melting process: Electric arc, argon oxygen decarburization.
Heat treatment: Quenched in air, water spray or oil, and tempered (cooling in furnace).
Typical microstructure: Tempered bainite and ferrite.
Weldability: Easily weldable with similar weld metal: weldments exposed to high stresses should be
welded by use of consumables with a carbon content ≥0.10 %.
High temperature applications: Casings of steam turbines, compressors, gas turbines, valves, nozzles and
chemical reactors; service temperatures up to about 570 °C.
Cast steel grade with similar chemical composition: ASTM A217 Grade WC9.
2.2.18.2 Standard requirements
Table 80. Chemical composition.
Standard
Designation
C
Si
Chemical composition [wt %]
Mn
P
S
Cr
Mo
EN10213- G17CrMo9-10
(1)
0.13-0.20 ≤0.60 0.50-0.90 ≤0.020 ≤0.020 2.00-2.50 0.90-1.20
2:1995
(1.7379)
(1) The maximum admissible sulphur content is 0.030 % if the relevant wall thickness is not in excess of
28 mm.
Table 81. Heat treatment and tensile properties at room temperature.
Min. 0.2%
Thickness
proof strength
Standard Designation
Heat treatment
[mm]
[MPa]
EN10213- G17CrMo9-10 Q:930 °C - 970 °C
150
400
2:1995
(1.7379)
T:680 °C - 740 °C
Tensile
strength
[MPa]
Min. elongation at rupture
[%]
590-740
18
Landolt-Börnstein
New Series VIII/2B
Ref. p. 113]
2.2.18 2.25Cr-1Mo cast
Rp0.2
min EN
Rm
600
800
500
400
Rm (MPa)
R p0.2(MPa)
111
300
200
100
0
600
400
200
0
0
100
200
300
400
500
600
0
Temperature °C
100
200
300
400
500
600
Temperature (°C)
Fig. 120. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade G17CrMo9-10 tested in creep
rupture tests by the German Creep Committee; [1]. min EN: minimum values by EN10213-2.
Stress (MPa)
1000
broken
400°C_EN
100
1000
10000
100000
Time to rupture (h)
Fig. 121. Creep rupture strength data of cast steel grade G17CrMo9-10 at 400 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
unbroken
450°C_EN
100
100
1000
10000
100000
1000000
Test duration (h)
Fig. 122. Creep rupture strength data of cast steel grade G17CrMo9-10 at 450 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
112
2.2 Low alloy steels
Stress (MPa)
1000
broken
100
unbroken
500°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 123. Creep rupture strength data of cast steel grade G17CrMo9-10 at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
unbroken
100
550°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 124. Creep rupture strength data of cast steel grade G17CrMo9-10 at 550 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
600°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 125. Creep rupture strength data of cast steel grade G17CrMo9-10 at 600 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 113]
2.2.18 2.25Cr-1Mo cast
113
2.2.18.3 Average creep rupture strength
Table 82. Average creep rupture strength values indicated in EN 10213-2:1995.l.
Time to rupture
Temperature
10,000 h
100,000 h
200,000 h
[°C]
Average creep rupture strength [MPa]
400
404
324
304
450
282
218
200
500
188
136
120
550
106
66
52
600
56
28
22
2.2.18.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade G17CrMo9-10, compilation of test results; Forschungsvereinigung
Warmfeste Stähle, c. o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
114
2.2 Low alloy steels
2.2.19 1Cr-1Mo-V cast
2.2.19.1 Introduction
G17CrMoV5-10 (EN 10213-2, material-no 1.7706) is the creep resistant cast steel grade which has been
used for turbine and valve casings most frequently since 1960. Creep rupture strength and creep strength
are high due to the vanadium content and the quenched and tempered microstructure, consisting mainly of
tempered bainite. The specified upper limit of tensile strength and solution temperature should not be
exceeded in order to avoid notch weakness. Oil quenching is preferred to obtain appropriate toughness
properties at room temperature combined with excellent creep rupture strength values and sufficient long
term ductility. The main features of G17CrMoV5-10 are summarized below:
Melting process: Electric arc, argon oxygen decarburization, induction melting.
Heat treatment: Quenched in air or liquid medium, and tempered (cooling in furnace).
Typical microstructure: Mainly tempered bainite with a small amount of ferrite.
Weldability: Weldable with similar weld metal; weldments exposed to high stresses (i.e. valves) should
be welded by use of similar consumables with a carbon content ≥0.12 %; higher preheating and longer
stress relief annealing (≥5 h) are necessary in comparison with other low alloy cast steel grades in order to
ensure sufficient toughness properties which otherwise may be impaired due to the vanadium content;
weldments exposed to low or moderate stresses may be welded with consumables of the type 2CrMo;
weld metal of the type 2CrMo is also recommended when root layers are required to show high toughness
properties.
High temperature applications: Casings of steam turbines, compressors, gas turbines, valves, nozzles;
service temperatures are up to about 570 °C if oxidation is negligible, otherwise the limiting temperature
is around 530 °C.
Cast steel grade with similar chemical composition: ASTM A 356 Grade 9.
2.2.19.2 Standard requirements
Table 83. Chemical composition.
Standard
Designation
C
EN10213- G17CrMoV5-10 0.152:1995
(1.7706)
0.20
Si
Mn
0.50≤0.60
0.90
Chemical composition [wt%]
P
S
Cr
Mo
1.20- 0.90≤0.020 ≤0.015
1.50
1.10
Table 84. Heat treatment and tensile properties at room temperature.
Min. 0.2 %
Thickness
proof strength
Standard Designation
Heat treatment
[mm]
[MPa]
EN10213- G17CrMoV5-10 Q:920 °C - 960 °C
150
440
2:1995
(1.7706)
T:680 °C - 740 °C
V
0.200.30
Other
Sn:
≤0.025
Tensile Min. elongastrength tion at rupture
[MPa]
[%]
590-780 15
Landolt-Börnstein
New Series VIII/2B
Ref. p. 117]
2.2.19 1Cr-1Mo-V cast
Rp0.2
min_EN
Rm
800
1000
800
600
Rm (MPa)
R p0.2 (MPa)
115
400
200
0
600
400
200
0
0
100
200
300
400
500
600
700
0
100
Temperature (°C)
200
300
400
500
600
700
Temperature (°C)
Fig. 126. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade G17CrMoV5-10 tested in creep
rupture tests by the German Creep Committee [1]. min EN: minimum values by EN 10213-2.
Stress (MPa)
1000
broken
unbroken
400°C_EN
100
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 127. Creep rupture strength data of cast steel grade G17CrMoV5-10 at 400 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
unbroken
450°C_EN
100
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 128. Creep rupture strength data of cast steel grade G17CrMoV5-10 at 450 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
116
2.2 Low alloy steels
Stress (MPa)
1000
broken
unbroken
500°C_EN
100
100
1000
10000
100000
1000000
Test duration (h)
Fig. 129. Creep rupture strength data of cast steel grade G17CrMoV5-10 at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
unbroken
100
550°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 130. Creep rupture strength data of cast steel grade G17CrMoV5-10 at 550 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
600°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 131. Creep rupture strength data of cast steel grade G17CrMoV5-10 at 600 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 117]
2.2.19 1Cr-1Mo-V cast
117
2.2.19.3 Average creep rupture strength
Table 85. Average creep rupture strength values indicated in EN 10213-2:1995.
Time to rupture
Temperature
10,000 h
100,000 h
200,000 h
[°C]
Average creep rupture strength [MPa]
400
463
419
395
450
340
275
254
500
229
171
157
550
151
96
83
600
80
28
19
2.2.19.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade G17CrMoV5-10, compilation of test results; Forschungsvereinigung
Warmfeste Stähle, c.o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
118
2.3 High Cr steels
2.3 High Cr steels
2.3.1 9Cr-1Mo steel
2.3.1.1 Introduction
Since its development for the oil industry in the 1930’s, 9Cr-1Mo steel (STBA 26, X11CrMo9-1+l1, +l2,
X11CrMo9-1+NT) has become an internationally accepted material as reflected by the several national
and international standards that exist to specify composition, heat treatment, minimum room temperature
tensile and stress rupture properties [1]. In particular, 9Cr-1Mo steel is now established as a structural
material for steam generator units by the UK nuclear industry following its successful service in the
evaporators and parts of the superheaters of the Advanced Gas Cooled Reactor. Furthermore, the steel is
also being used for the replacement superheater and reheater tube bundles now completing construction
for the UK Prototype Fast Reactor.
2.3.1.2 Material standards, chemical and tensile requirements
2.3.1.2.1 9Cr-1Mo steel tubes for boilers and heat exchangers
Table 86. Chemical requirements for 9Cr-1Mo steel tubes; JIS STBA 26, ASTM T9 and BS 629-470.
Standard
Designation
JIS
STBA 26
ASTM
T9
ASTM
T9
BS
629-470
C
Si
<0.15 0.25 1.00
<0.15 0.25 1.00
<0.15 0.25 1.00
<0.15 <1.00
Chemical composition [wt%]
Mn
P
S
Cr
0.3 0.6
0.3 0.6
0.3 0.6
0.3 0.6
<0.030 <0.030 8.00 10.00
<0.025 <0.025 8.00 10.00
<0.025 <0.025 8.00 10.00
<0.030 <0.030 8.00 10.00
Yield
Others strength
[MPa]
0.90 >205
1.10
0.90 >170
1.10
0.90 >205
1.10
0.90 - Al:
>185
1.10 <0.02
Mo
Tensile Std.
strength No.
[MPa]
>410
G3462
>415
A199
>415
A213
470 620
3059-2
2.3.1.2.2 9Cr-1Mo steel pipes
Table 87. Chemical requirements for 9Cr-1Mo steel pipes; JIS STPA 26, ASTM P9 and BS 629-470.
Standard
Designation
JIS
STPA 26
ASTM
P9
BS
629 - 470
C
Si
<0.15 0.25 1.00
<0.15 0.25 1.00
<0.15 0.25 1.00
Chemical composition [wt%]
Mn
P
S
Cr
0.3 0.6
0.3 0.6
0.3 0.6
Yield
Others strength
[MPa]
<0.030 <0.030 8.00 - 0.90 >205
10.00 1.10
<0.025 <0.025 8.00 - 0.90 >205
10.00 1.10
<0.030 <0.030 8.00 - 0.90 - Al:
>185
10.00 1.10 <0.02
Mo
Tensile Std.
strength No.
[MPa]
>410
G3458
>415
A335
470 620
3604-1
2.3.1.3 Data sources for 9Cr-1Mo steel
Information of fact on data for 9Cr-1Mo steel tubes can be obtained from [1], [2] and [3].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 125]
2.3.1 9Cr-1Mo steel
119
2.3.1.4 Creep rupture data for 9Cr-1Mo steel tubes for boiler and heat exchangers, JIS STBA 26
2.3.1.4.1 Creep rupture data for 9Cr-1Mo steel tubes, JIS STBA 26
The complete set of creep and creep rupture data, such as creep rupture time, total elongation, reduction
of area and optical micrographs of as-received and crept specimens, has been published for 11 heats of
9Cr-1Mo steel, JIS STBA 26, in [2]. The details of tube production, processing, thermal history, austenite
grain size number, Rockwell hardness and volume fraction of non-metallic inclusions before creep test,
the chemical compositions, the 0.2% proof stress and ultimate tensile strength data at high temperature
are also available for the 11 heats in [2].
The details of tube production and the chemical compositions are also given in [2]. The tensile and
creep specimens, having a geometry of 6 mm in diameter and 30 mm in gauge length, were taken
longitudinally from the middle of wall thickness of the tubes. Fig. 132 shows the 0.2% proof stress and
tensile strength obtained by short-time tensile tests between room temperature and 700 °C.
Tensile strength
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
600
300
300
200
200
100
100
0
0
100 200 300 400 500 600
700
800
0
0
Test temperature (℃)
100
200
300
400
500
600
700
800
Test temperature (℃)
Fig. 132. Short-time tensile properties of 9Cr-1Mo steel, JIS STBA 26.
Fig. 133 shows stress vs. time to rupture data for 11 heats of 9Cr-1Mo steel, JIS STBA 26, at
temperatures between 550 and 700 °C. Fig. 133 exhibits a large heat-to-heat variation in time to rupture,
which becomes more significant with increasing time and temperature. At 550 °C and 78 MPa, the time
to rupture of the strongest heat is 1.3 × 105 hours, while that of the weakest heat is only 2.7 × 104 hours. It
should be noted, however, that the observed heat-to-heat variation in time to rupture is not caused by data
scattering, because each heat has its distinct stress dependence of time to rupture as shown in Fig. 134.
The heat-to-heat variation in time to rupture comes from many factors, such as differences in production
history, chemical composition and initial microstructure.
2.3.1.4.2 Estimated long-term creep rupture strength for 9Cr-1Mo steel tubes, JIS STBA 26
The creep rupture data shown in Fig. 133 were analyzed for each heat using the Larson-Miller parameter
method. Fig. 135 shows the results for a heat (MEH) with an intermediate strength level among the 11
heats. The 105 h creep rupture strength was also estimated for the 11 heats. This is shown in Fig. 136 as a
function of temperature, together with 0.2% proof stress, ultimate tensile strength and 103 h creep rupture
strength. The regression equations for tensile and creep rupture strength are described in [1].
Landolt-Börnstein
New Series VIII/2B
120
2.3 High Cr steels
300
550 °C
600 °C
650 °C
700 °C
Stress [MPa]
200
100
80
60
50
40
30
Fig. 133. Creep rupture strength
data for 9Cr- 1Mo steel,
JIS STBA 26. n indicates the
total number of data points.
20
n = 270
10
10
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
200
9Cr-1M o
o
Stress ( MPa )
550 C
ME B
ME C
ME D
M EE
ME F
ME G
ME H
ME J
ME L
M EM
ME N
100
90
80
70
Fig. 134. Creep rupture data for each heat of
9Cr-1Mo steel, STBA 26, at 550 °C.
60
10
2
10
3
10
4
10
5
T ime to ru pture ( h )
2 00
o
55 0 C
Stress ( MPa )
1 00
o
600 C
50
o
65 0 C
o
700 C
20
10
2
10
Fig. 135. Estimated creep rupture
curves of 9Cr-1Mo steel STBA 26.
10
3
10
4
10
5
10
6
T ime to ruptu re ( h )
Landolt-Börnstein
New Series VIII/2B
Ref. p. 125]
2.3.1 9Cr-1Mo steel
121
800
600
500
400
300
Stress [MPa]
200
Tensile
{ strength
{ 0.2%stressproof
100
80
60
50
40
30
100000 h
20
10
500
550
600
650
Temperature [°C]
1000 h
700
750
Fig. 136. Estimated 105 h creep
rupture strength for the 11 heats of
9Cr-1Mo steel, JIS STBA 26, as a
function of temperature, together
with 0.2% proof stress, tensile
strength and 103 h creep rupture
strength.
2.3.1.5 Creep behavior and microstructure of 9Cr-1Mo steel
2.3.1.5.1 Effect of secondary precipitation on the creep strength of normalized and tempered
9Cr-1Mo steel
In order to demonstrate the importance of secondary precipitation on the creep and stress rupture
properties and the difficulties in extrapolation of stress rupture data for design purposes, detailed
microstructural examination has been made for laboratory creep tested specimens of 9Cr-1Mo steel at
temperatures between 475 and 550 °C [5-7]. The material was received as 25 mm bar in the normalized
(1000 °C, 1 h) and tempered (750 °C, 2 h) condition. The chemical composition was 8.6Cr-1.04Mo0.12C-0.69Si-0.48Mn-0.008S- 0.022P-0.22Ni- 0.05Cu (wt%).
The stress-rupture data are presented in Fig. 137, together with the estimated ISO values (ISO 1971)
for the rupture life for normalized and tempered 9Cr-1Mo steel. The composition and heat treatment
combination for this steel provides appreciably greater strength than is predicted from the ISO data. At
475 °C, the curve is concave downwards, distinct inflexion points are obtained at 500 and 525 °C,
whereas at 550 °C the curve is concave upwards.
450
400
350
475 °C
500 °C
Stress [N /mm 2 ]
300
525 °C
250
550 °C
I.S.
O. 5
0
0°
200
C
I.S.
O. 5
2
150
5°
C
I.S
.O.
100
650 °C
10
Landolt-Börnstein
New Series VIII/2B
10 2
55
0°
C
600 °C
10 3
Rupture time [h]
10 4
10 5
Fig. 137. Stress vs. time to rupture data for normalized
and tempered 9Cr-1Mo steel bars compared with ISO
data.
122
2.3 High Cr steels
Fig. 138 shows the isochronal strain data up to 2 % at 475 °C. The log ε versus log σ lines generally
diverge at 500, 525 and 550 °C with increasing strain and stress, but the opposite behavior occurs at
475 °C up to 1000 h testing. Divergence is often obtained with this steel, in which no period of steady
state rate elongation occurs. Instead, after the initial primary extension, the creep rate continuously
processes to failure. This can be attributed to recovery processes increasing at a greater rate than the strain
hardening, i.e. microstructural degeneration is occurring such that the dislocation structure is not
stabilized. At 475 °C, the divergence is not apparent until after 1000 h testing. At this temperature the
strain-time curves generally show a period of steady state creep, more extensive than at higher
temperatures, prior to rapid elongation to failure. Hence it appears that after 1000 h at 475 °C,
degeneration of the microstructure becomes important and the isochronal log ε versus log σ lines diverge
as at the higher temperatures.
450
Stress [N/mm 2 ]
400
0.5 % e
0.3 % e
Temperature 475 °C
350
2.0 % e
300
1.0 % e
0.5 % e
0.3 % e
10 2
10
10 3
Time [h]
10 4
10 5
Fig. 138. Stress vs. time relationship for normalized and
tempered 9Cr-1Mo steel bars at 475 °C.
However, several creep strain-time curves have been obtained which show perturbations, a marked
example of which is shown in Fig. 139. In this case, a steady state creep stage was established at low
strain, followed by a period of accelerating creep and then by a further period of steady state creep. Other
cases, where the perturbation occurs soon after primary extension, and at higher temperatures, have been
obtained, but not at 550 °C.
10
Temperature 500 °C
Stress 232 N/mm 2
Duplicate tests
repeat test
terminated
test
continuing
Strain [%]
1.0
0.1
1
10
10 2
10 3
Time [h]
10 4
10 5
Fig. 139.
Multi-stage secondary creep observed in
normalized and tempered 9Cr-1Mo steel bars at 500 °C.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 125]
2.3.1 9Cr-1Mo steel
123
The microstructure analysis concentrated mainly on the creep tested specimens at above and below the
inflexion point at 500 °C shown on the stress rupture plot of Fig. 137. Some specimens tested at 475 and
550 °C were also examined.
The microstructure of the gauge region of creep specimens above the inflexion (2000 h at 500 °C)
shows that normal recovery creep events have commenced locally. The recovery processes have led to a
local reduction in the intrinsic dislocation density, together with the onset of subcell formation and some
regions of secondary precipitate. Below the inflexion point (3×104 h at 500 °C), recovery is again
inhomogeneous, but additionally in those areas where recovery has occurred, significant secondary
precipitation on or near dislocations has taken place. The secondary precipitate is up to 50 nm in diameter
and precipitates separately or as a stack-like sequence of thin plates either on or near dislocations. The
secondary precipitate is identified as M2X type carbonitrides with high Cr content, together with some
substitutional Fe and Mo content. No secondary precipitate similar to that found in the gauge is visible,
and hence the precipitation in the gauge occurs or is enhanced by the presence of creep strain.
Examination of a 1.5×104 h test at 475 °C and 302 MPa shows a fine general precipitate of similar, or
higher density, than that in the 500 °C, 3×104 h specimen’s gauge. However, the structure of a 550 °C,
132 MPa, 104 h specimen shows little or no such precipitate, and is very similar to the grip region of the
500 °C, 3×104 h specimen.
It is concluded that the onset of secondary precipitation leads to modest increases in creep and rupture
strength of normalized and tempered 9Cr-1Mo steel. Normal microstructural degeneration processes are
perturbed under these circumstances giving rise to sigmoidal shaped creep and stress rupture curves.
2.3.1.5.2 Effect of heat treatment on the tensile strength of normalized and tempered 9Cr-1Mo steel
The potential of 9Cr-1Mo steel for use in thick sections has been assessed by using simulation heat
treatments [4]. This work involved the laboratory-scale cooling of bar samples to simulate waterquenching rates in cylindrical section up to 720 mm diameter (equivalent to 500 mm thick plate). The
material, available as 20 mm diameter bars, originated from two commercial casts of 9Cr-1Mo steel
manufactured by British Steel Corporation, one of the batches were from material produced by electroslag
refining, the other by air melting. Bundles of 20 mm diameter bars were cooled from the austenitizing
temperatures at rates simulating those at the centers of 375 and 500 mm thick plates during water
quenching. Tables 88 and 89 give the chemical compositions and the heat treatment conditions,
respectively, of 9Cr-1Mo steel examined.
The experimental work was carried out in two separate phases : Phase 1 - an investigation of the effect
of a 375 mm thick section simulation on tensile strength using relatively rapid furnace heating rates and
short soak periods at austenitizing and tempering temperatures (Cast RM348). Phase 2 - as for phase 1,
but extended to include the effects of a 500 mm thick plate simulation with similar heating rates and soak
periods to those used commercially for large section forgings (Cast D9893). Duplicate tensile tests were
carried out at room temperature and 525 °C.
Table 88. Chemical compositions of 9Cr-1Mo steel for tensile tests.
Identity
RM348 (1)
D9893 (2)
Landolt-Börnstein
New Series VIII/2B
C
0.09
0.10
Si
0.84
0.66
Mn
0.46
0.50
Chemical composition [wt%]
P
S
Cr
Mo
Ni
Cu
0.018 0.007 9.8
1.00 0.1
0.1
0.030 0.008 8.85 0.95 0.21 0.15
Co
0.03
0.15
Al
Sn
0.009 0.02
0.009 0.02
124
2.3 High Cr steels
Table 89. Heat treatment conditions of 9Cr-1Mo steel for tensile tests.
Austenititzing
Phase Cast
Identity Temp. Heating Cooling rate
Temp.
times
rate
times
[°C/h]
1
RM348 920 °C 2000
Simulated 375 mm 700 °C, 2 h
1h
thick WQ
780 °C, 2 h
980 °C
1h
2000
Simulated 375 mm
thick WQ
980 °C,
1h
980 °C
15 h
980 °C
15 h
980 °C
15 h
980 °C
15 h
WQ: Water quenched
2000
Simulated 375 mm
thick WQ
Simulated 375 mm
thick WQ
Simulated 375 mm
thick WQ
Simulated 375 mm
thick WQ
Simulated 375 mm
thick WQ
2
D9893
30
30
30
30
Tempering
Heating Cooling rate
rate
[°C/h]
1500
Air cooled as 20 mm
diameter
1500
Air cooled as 20 mm
diameter
700 °C, 2 h 1500
Air cooled as 20 mm
diameter
700 °C, 2 h 1500
Air cooled as 20 mm
diameter
770 °C
1500
Air cooled as 20 mm
2h
diameter
770 °C
30
Air cooled as 20 mm
16 h
diameter
770 °C
30
Furnace cooled
16 h
770 °C
30
Air cooled as 20 mm
16 h
diameter
770 °C
30
Furnace cooled
16 h
The results on 0.2% proof stress and tensile strength are summarized in Fig. 140 and 141, showing the
effects of variations in heat treatment temperatures and section sizes, respectively. The information in
Fig. 140 relates to the 0.84% Si steel (Cast RM348) and it is seen that austenitizing at 980 °C results in
higher strength at room temperature, but lower values at 525 °C relative to those produced after
austenitizing at 920 °C. The material tempered at 700 °C had a small strength advantage over material
tempered at 780 °C for room temperature tests, but this trend was reversed for tests at 525 °C, presumably
as a consequence of secondary precipitation. The results from Phases 1 and 2 when taken together as in
Fig. 141, demonstrate that differences in cooling rate equivalent to water-quenching sections up to
500 mm thickness have little effect on the tensile properties. However, the results show significant
strengthening effects of Si and further demonstrate that the strength is obtained relative to that from the
more rapid laboratory-type treatments.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 125]
2.3.1 9Cr-1Mo steel
800
700
125
20 AC
550 (375) WQ
550 (375) WQ
720 (500) WQ
LAB
980
0.84
LAB LAB
980 980
0.84 0.66
COM COM
980 980
0.66 0.66
COM COM
980 980
0.66 0.66
Bar dia.
(plate thick) mm
Tensile strength [MPa]
Tensile strength [MPa]
700
600
500
400
300
920 920
700 780
980 980
700 780
600
500
400
300
aust.temp. [°C]
temp.temp.[°C]
heating rate
St temp. [°C]
silicon [%]
0.2 % Proof stress [MPa]
600
500
400
RT min.props.
(BS3059)
20 °C
525 °C
value not
obtained-test
fault
0.2 % Proof stress [MPa]
200
500
400
RT min.props.
(BS3059)
20 °C
525 °C
300
200
AC
AC AC
AC FC
AC FC
cool rate from
tempering
300
200
Fig. 140. Effect of heat treatment temperatures on
9Cr-1Mo steel simulated cooled as 375 mm thick
WQ (Cast RM348).
Fig. 141. Effect of heat treatment, composition and simulated
bar size on tensile properties of 9Cr-1Mo steel.
abbreviations in Fig. 140 and 141: LAB = laboratory-type rapid heating rate; COM = commercial-type slow heating
rate; RT min. prop. = minimum value of 0.2% proof stress and tensile strength at room temperature.
2.3.1.6 References
[1]
[2]
[3]
[4]
NRIM Creep Data Sheet, No. 19B, (1997).
ASTM Data Series Publication DS50, (1973)
The British Steelmakers Creep Committee (BSCC) High Temperature Data (1972)
Orr, J., and Sanderson, S. J.: Proceedings of Topical Conference on Ferritic Alloys for Use in
Nuclear Energy Technologies, Snowbird, Utah, June 19-23, (1983), 261-267.
[5] Williams, K. R., Fidler, R. S. and Askins, M.C.: Proceedings of International Conference on Creep
and Fracture of Engineering Materials and Structures, University College, Swansea, UK, (1981),
pp.475 - 487.
[6] Sanderson, S. J.: Metal Science 11 (1977), 490 - 492.
[7] Sanderson, S. J.: Metal Science 12 (1978), 220 - 222.
Landolt-Börnstein
New Series VIII/2B
126
2.3 High Cr steels
2.3.2 9Cr-1Mo-V-Nb steel
2.3.2.1 Introduction
9Cr-1Mo-V-Nb steel was developed in the early 1980s [1]. It is improved in creep strength by
precipitation strengthening of fine MX carbonitride with addition of vanadium and niobium. 9Cr-1Mo-VNb steel has already been widely used as tubes for heat exchangers, pipes for high temperature service,
plates for pressure vessels, forged, rolled and cast steels for high temperature services.
2.3.2.2 Materials standards, and chemical and tensile requirements
2.3.2.2.1 9Cr-1Mo-V-Nb steel tubes for heat exchangers
Table 90. Chemical requirements of 9Cr-1Mo-V-Nb steel tubes; Japanese METI KA STBA28 and
ASTM T91
Chemical composition [wt%]
Std.
Standards
Designation
No
C
Si
Mn
P
S
Cr
Japanese
KA STBA 28 0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50
METI
ASTM
T91
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50 A-213
Standards
Japanese
METI
ASTM
Designation
Mo
V
Chemical composition [wt%]
Nb
Ni
N
sol. Al
KA STBA 28
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
T91
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
Std. No
A-213
2.3.2.2.2 9Cr-1Mo-V-Nb steel pipes for high temperature services
Table 91. Chemical requirements of 9Cr-1Mo-V-Nb steel pipes; Japanese METI KA STPA28 and
ASTM P91
Chemical composition [wt%]
Standards
Designation
Std. No
C
Si
Mn
P
S
Cr
Japanese
KA STPA 28
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50
METI
ASTM
P91
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50 A-335
Standards
Japanese
METI
ASTM
Designation
Mo
V
Chemical composition [wt%]
Nb
Ni
N
sol. Al
KA STPA 28
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
P91
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
Std. No
A-335
Landolt-Börnstein
New Series VIII/2B
Ref. p. 132]
2.3.2 9Cr-1Mo-V-Nb steel
127
2.3.2.2.3 9Cr-1Mo-V-Nb steel plates for pressure vessels
Table 92. Chemical requirements of 9Cr-1Mo-V-Nb steel plates; Japanese METI KA SCMV28 and
ASTM Gr.91
Chemical composition [wt%]
Standards
Designation
Std. No
C
Si
Mn
P
S
Cr
Japanese
KA SCMV 28 0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50
METI
ASTM
Gr.91
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50 A-387
Standards
Japanese
METI
ASTM
Designation
Mo
V
Chemical composition [wt%]
Nb
Ni
N
sol. Al
KA SCMV 28
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
Gr.91
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
Std. No
A-387
2.3.2.2.4 9Cr-1Mo-V-Nb forged or rolled steel for high temperature services
Table 93. Chemical requirements of 9Cr-1Mo-V-Nb forged or rolled steels; Japanese METI KA
SFVAF28 and ASTM F91
Chemical composition [wt%]
Standards
Designation
Std. No
C
Si
Mn
P
S
Cr
Japanese
KASFVAF28 0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50
METI
ASTM
F91
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50 A-182
Standards
Japanese
METI
ASTM
Designation
Mo
V
Chemical composition [wt%]
Nb
Ni
N
sol. Al
KA SFVAF28
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
F91
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
Std. No
A-182
2.3.2.2.5 9Cr-1Mo-V-Nb steel castings for high temperature services
Table 94. Chemical requirement of 9Cr-1Mo-V-Nb steel castings; Japanese METI KA SCPH91 and
ASTM C12A
Chemical composition [wt%]
Standards
Designation
Std. No
C
Si
Mn
P
S
Cr
Japanese
KA SCPH91
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50
METI
ASTM
C12A
0.08-0.12 0.20-0.50 0.30-0.60 ≤0.020 ≤0.010 8.00-9.50 A-217
Standards
Japanese
METI
ASTM
Landolt-Börnstein
New Series VIII/2B
Designation
Mo
V
Chemical composition [wt%]
Nb
Ni
N
sol. Al
KA SCPH91
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.03-0.07 ≤0.04
C12A
0.85-1.05 0.18-0.25
0.06-0.10
≤0.40 0.030-0.070 ≤0.04
Std. No
A-217
128
2.3 High Cr steels
2.3.2.3 Creep properties of 9Cr-1Mo-V-Nb steel
Information of fact on creep data for 9Cr-1Mo-V-Nb steels can be obtained from [2 - 4].
2.3.2.3.1 Creep rupture data of 9Cr-1Mo-V-Nb steel
Results of creep tests for 3 heats of Japanese METI KA STBA28 steel tubes are given in [2]. From [2],
data on tensile properties, rupture elongation, reduction of area, minimum creep rate, and microstructure
of as-received materials can also be obtained.
Creep rupture strength data of Japanese METI KA STBA28 steel tubes over a temperature range from
500 to 700 °C is shown in Fig. 142 [2]. The slope of stress vs. time to rupture curves increases with
decrease in applied stress. This inflection of creep rupture curves is due to a change of the degradation
mechanism from homogeneous recovery during short-term exposure to inhomogeneous recovery during
long-term exposure, as will be explained later.
Results of creep tests for 9Cr-1Mo-V-Nb steel tubes, pipes and plates are given in [3]. From [3] data
on product form, size, chemical composition, heat treatment condition, tensile properties, rupture
elongation and reduction of area can also be obtained.
500
○
△
□
×
▽
Stress (MPa)
300
500 oC
550 oC
600 oC
650 oC
700 oC
100
80
60
Fig. 142.
Creep rupture
strength data of Japanese
METI KA STBA 28; [2]. n
indicates the total number of
data points.
40
n=59
20
1
10
2
10
3
10
4
10
5
10
6
10
Time to rupture (h)
Creep rupture strength of 9Cr-1Mo-V-Nb steel tubes, pipes and plates at 600 and 650 °C are shown in
Fig. 143 and 144, respectively, for tempering temperatures between 760 °C and 790 °C [3]. Slightly
higher creep rupture strength is observed for steels with lower tempering temperature, at both
temperatures of 600 and 650 °C.
Creep rupture ductility data of Japanese METI KA STBA28 steel tubes over a temperature range from
500 to 700 °C is shown in Fig. 145 [2]. 9Cr-1Mo-V-Nb steel has a good rupture ductility. In some cases,
however, a decreasing tendency in rupture ductility is observed during long-term exposure. This effect is
thought to be caused by inhomogeneous recovery in the vicinity of prior austenite grain boundaries,
similar to inflection of creep rupture curves, as will be explained later.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 132]
2.3.2 9Cr-1Mo-V-Nb steel
129
300
Tempering
○
△
□
×
o
Stress (MPa)
600 C
200
temperature
790 oC
780 oC
765 oC
760 oC
Fig. 143.
Creep rupture
strength of 9Cr-1Mo-V-Nb
steels at 600 °C; [3].
100
1
10
2
10
10
3
10
4
10
5
Time to rupture (h)
Stress (MPa)
200
Tempering
○
△
□
×
100
90
80
70
60
temperature
790 oC
780 oC
765 oC
760 oC
o
650 C
50
40
1
10
2
10
10
3
4
10
10
5
Fig. 144.
Creep rupture
strength data of 9Cr-1Mo-VNb steels at 650 °C; [3].
Time to rupture (h)
Rupture elongation and
reduction of area (%)
100
80
60
Open:
Rupture elongation
Solid:
Reduction of area
40
20
0
1
10
Fig. 145. Creep rupture ductility of 9Cr-1Mo-V-Nb
steels; [2].
10
2
10
3
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
10
4
5
10
130
2.3 High Cr steels
2.3.2.3.2 Creep deformation behavior and creep rupture strength
Creep rate vs. strain curves of 9Cr-1Mo-V-Nb steel tubes at 600 °C and over the range of stresses from
100 to 200 MPa are shown in Fig. 146. Creep deformation consists of transient and accelerating creep
stages and no obvious steady state creep stage is observed. With decrease in applied stress, the onset
strain of the accelerating creep stage decreases from about 0.03 at 200 MPa to less than 0.01.
2.3.2.3.3 Microstructural change
-1
Creep rate (h )
Heat treatment condition of 9Cr-1Mo-V-Nb steel is normalizing and tempering. Tempered martensitic
microstructure is typical for the as heat treated condition of the steel.
The influence of microstructures on mechanical properties and changes in microstructures during
creep exposure of 9Cr-1Mo-V-Nb steel have been widely investigated [5-29]. It is well known that fine
MX carbonitride of Nb(C, N) and V(C, N) is very stable at elevated temperatures and plays an important
role in improving creep strength.
10
-1
10
-2
10
-3
10
-4
10
-5
10
-6
10
-7
200MPa
160MPa
140MPa
120MPa
110MPa
600oC
Fig. 146. Creep rate vs. strain curves at 600oC of
Japanese METI KA STBA 28; [5].
100MPa
10
-4
10
-3
10
-2
10
-1
10
0
True strain
Bright field TEM images of the steel in the as heat treated condition and the creep ruptured specimen are
shown in Figure 147 [5]. During creep exposure at elevated temperatures, recovery of tempered
martensitic microstructure, such as decrease in dislocation density and increases in lath width and
subgrain size, takes place. Homogeneous progress in recovery is observed in the specimens creep
ruptured during short-term exposure, as shown in Fig. 147(b) and (c). On the other hand, inhomogeneous
progress in recovery is observed in the specimen creep ruptured after 34,141.0 h at 600 °C and 100 MPa,
as shown in Fig. 147(d). The microstructure within grains is very fine, in contrast to significantly
recovered regions in the vicinity of prior austenite grain boundaries. Although the creep exposure time of
34,141.0 h at 600 °C and 100 MPa is about three times longer than that of 12,858.6 h at 600 °C and 120
MPa, the subgrain size within grains of the former specimen is smaller than that of the latter one. Bimodal
distribution of subgrain size is clearly observed only in the specimen creep ruptured at 600 °C and 100
MPa.
Inhomogeneous recovery preferentially taking place in the vicinity of prior austenite grain boundaries
is regarded to be a degradation mechanism during long-term creep exposure, in contrast to homogeneous
recovery during short-term exposure. Inflection of creep rupture curves mentioned in section 2.3.2.3.1 is
thought to be caused by such differences in recovery phenomena during short-term and long-term
exposure. It seems that decrease in rupture ductility during long-term exposure (Fig. 145) and decrease in
onset strain of accelerating creep stage with decrease in stress (Fig. 146) is also derived from
inhomogeneous recovery preferentially taking place in the vicinity of prior austenite grain boundaries [5].
There are three types of precipitates in the as heat treated condition: M23C6 carbide, NbX (niobium
carbonitride) and VX (vanadium carbonitride) [26]. Precipitation of Laves phase (Fe2Mo) takes place
after short-term exposure at temperatures lower than 600 °C. The role of precipitation of Laves phase on
creep strength has not yet been clearly understood. It has been reported that Laves phase improves creep
Landolt-Börnstein
New Series VIII/2B
Ref. p. 132]
2.3.2 9Cr-1Mo-V-Nb steel
131
strength through a precipitation strengthening effect [29]. On the other hand, the possibility to decrease
creep strength by reducing the solid solution strengthening effect of molybdenum has been also pointed
out.
Precipitation of modified Z-phase, a complex nitride of chromium, niobium and vanadium in a form
of Cr(Nb,V)N [30], takes place after several thousand hours of creep exposure at elevated temperatures
[26].
Changes in mean diameter of precipitates with increase in creep exposure time at 600 and 650 °C in
9Cr-1Mo-V-Nb steel tubes are shown in Fig. 148 [26]. The size of MX carbonitride is about 30 nm in the
as heat treated condition and less than 100 nm even after long-term creep exposure. The coarsening rate
of MX carbonitride is very small and MX carbonitride is finer than that of M23C6 carbide. Precipitation of
modified Z-phase takes place after several thousand hours of creep exposure and its growth rate after
nucleation is significantly large, in comparison with those of MX carbonitride and M23C6 carbide. Since
modified Z-phase grows at the expense of fine MX carbonitride, its nucleation and rapid growth results in
a decrease in number density of fine MX carbonitride and reduces creep strength during long-term
exposure.
Fig. 147. Bright field TEM images of 9Cr-1Mo-V-Nb steel (a) in the as tempered condition and creep exposed at
600 °C for (b) 971.2 h at 160 MPa, (c) 12,858.6 h at 120 MPa and (d) 34,141.0 h at 100 MPa [5].
～
～
・
Solid line: 650oC
Dashed line: 600oC
Laves
103
Z phase
M23C6
102
・
・
～
～ ～
～
Mean diameter (nm)
104
Fig. 148.
Changes in mean diameter of the
precipitates in 9Cr-1Mo-V-Nb steel with increase in
creep exposure time at 600 and 650 °C [26].
MX
・
～
～
101
101
102
103
104
105
Time to rupture (h)
2.3.2.3.4 Long-term creep strength prediction
Inflection of stress vs. time to rupture curves of 9Cr-1Mo-V-Nb steel makes it difficult to predict longterm creep strength from the short-term data. Creep rupture data of 9Cr-1Mo-V-Nb steel tubes in the
temperature range from 500 to 700 °C is shown in Fig. 149. Creep rupture life predicted from short-term
data up to 20,000 h using the Larson-Miller parameter with a best fit parameter constant of 38 are also
indicated in the figure. Since the curve indicates inflection, predicted long-term creep rupture strength is
extremely overestimated, especially at 600 and 650 °C.
Landolt-Börnstein
New Series VIII/2B
132
2.3 High Cr steels
From the above point of view, new life prediction methods have been investigated on high chromium
ferritic creep resistant steels. One of the proposed new life prediction methods is demonstrated in Fig. 150
on the same creep rupture data as shown in Fig. 149 [28]. Here, creep rupture data is divided into two
groups of short-term and long-term exposure with a boundary condition of half of 0.2% proof stress at
each temperature. Creep rupture data where the stresses are lower than half of 0.2% proof stress should be
used for life prediction using the Larson-Miller parameter with a constant of 20. Proper results of longterm creep strength prediction are obtained with this proposed prediction method.
500
o
500 C
o
550 C
Stress (MPa)
300
o
100
80
600 C
o
650 C
60
Fig. 149. Stress vs. time to rupture curves of 9Cr1Mo-V-Nb steel tubes and predicted long-term creep
rupture life using the Larson-Miller parameter and a
best fit parameter constant of 38.
o
700 C
40
20
1
10
Predicted
(C=38)
10
2
3
10
4
5
10
6
10
10
Time to rupture (h)
500
Stress (MPa)
300
o
500 C
o
550 C
100
80
o
600 C
60
40
20
1
10
o
650 C
0.2% proof stress
2
2
10
o
700 C
3
10
4
10
10
5
6
Fig. 150. Stress vs. time to rupture curves of 9Cr1Mo-V-Nb steel tubes and predicted long-term creep
rupture life by considering half of 0.2% proof stress
[28].
10
Time to rupture (h)
2.3.2.4 References
[1] Sikka, V. K.: Proc. of Topical Conf. on Ferritic Alloys for Use in Nuclear Energy Technologies,
Davis, J. W., and Michel, D. J., eds., TMS-AIME, Warrendale, Pennsylvania, USA, (1984), 317327.
[2] National Research Institute for Metals: NRIM Creep Data Sheet, No.43, (1996).
[3] The Iron and Steel Institute of Japan: Atlas of Stress-Strain Curves and Rupture Data at High
Temperatures, (1994).
[4] Japan Pressure Vessel Research Committee: 0.5Mo and Cr-Mo steels Data Book, (1998).
[5] Kimura, K., Kushima, H., and Abe, F.: Key Engineering Materials, 171-174 (2000), 483-490.
[6] Vitek, J.M., and Klueh, R.L.: Metall. Trans. A, 14A (1983), 1047-1055.
[7] Jemian, P.R., Weertman, J.R., Long, G.G., and Spal, R.D.: Acta Metall. Mater., 39 (1991), 24772487.
[8] Jones, W.B., Hills, C.R., and Polonis, D.H.: Metall. Trans. A, 22A (1991), 1049-1058.
[9] Iseda, A., Sawaragi, Y., and Yoshikawa, K.: Tetsu-to-Hagane, 77 (1991), 582-589.
[10] Brinkman, C.R., Gieseke, B., and Maziasz, P.J.: Proc. of Microstructures and Mechanical Properties
of Aging Materials, Liaw, P.K. et al. eds., TMS, (1993), 107-115.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 132]
[11]
[12]
[13]
[14]
[15]
[16]
[17]
[18]
[19]
[20]
[21]
[22]
[23]
[24]
[25]
[26]
[27]
[28]
[29]
[30]
2.3.2 9Cr-1Mo-V-Nb steel
133
Hamada, K., Tokuno, K., and Takeda, T.: Nuc. Eng. Des., 139 (1993), 277.
Ruggles, M.B., Cheng, S., and Krempl, E.: Mater. Sci. Eng., A186 (1994), 15-21.
Tsuchida, Y., Takeda, T., and Tokunaga, Y.: Tetsu-to-Hagane, 80 (1994), 723-728.
Ogata, T.: Soc.,J., Mat. Sci., Japan, 46 (1997), 25-31.
Spigarelli, S., Kloc, L. and Bontempi, P.: Scripta Mater., 37 (1997), 399-404.
Cadek, J., Sustek, V., and Pahutova, M.: Mater. Sci. Eng., A225 (1997), 22-28.
Kloc, L., and Sklenicka, V.: Mater. Sci. Eng., A234-236 (1997), 962-965.
Sawada, K., Maruyama, K., Komine, R., and Nagae, Y.: Tetsu-to-Hagane, 83 (1997), 466-471.
Sawada, K., Takeda, M., Maruyama, K., Komine, R., and Nagae, Y.: Tetsu-to-Hagane, 84 (1998),
580-585.
Orlova, A., Bursik, J., Kucharova, K., and Sklenicka, V.: Mater. Sci. Eng., A245 (1998), 39-48.
Cerri, E., Evangelista, E., Spigarelli, S., and Boanchi, P.: Mater. Sci. Eng., A245 (1998), 285-292.
Park, K.S., Masuyama, F., and Endo, T.: Tetsu-to-Hagane, 84 (1998), 526-533.
Park, K.S., Masuyama, F., and Endo, T.: Tetsu-to-Hagane, 84 (1998), 553-558.
Kushima, H., Kimura, K. and Abe, F.: Tetsu-to-Hagane, 85 (1999), 841-847.
Spigarelli, S., Cerri, E., Bianchi, P., and Evangelista, E.: Mater. Sci. Technol., 15 (1999), 14331440.
Suzuki, K., Kumai, S., Kushima, H., Kimura, K., and Abe, F.: Tetsu-to-Hagane, 86 (2000), 550-557.
Kimura, K., Suzuki, K., Toda, Y., Kushima, H., and Abe, F.: Proc. of 7th Liege Conference on
Materials for Advanced Power Engineering 2002, J. Lecomte-Beckers et al. eds.,
Forschungszentrum Jülich GmbH, Liege, Belgium, September 2002, 21 (2002), 1171-1180.
Kushima, H., Kimura, K., and Abe, F.: Proc. of 7th Liege Conference on Materials for Advanced
Power Engineering 2002, J. Lecomte-Beckers et al. eds., Forschungszentrum Jülich GmbH, Liege,
Belgium, September 2002, 21 (2002), 1581-1590.
Hald, J., and Korcakova, L.: ISIJ Inter., 43 (2003), 420-427.
Strang, A., and Vodarek, V.: Mater. Sci. Technol., 12 (1996), 552.
Landolt-Börnstein
New Series VIII/2B
134
2.3 High Cr steels
2.3.3 9Cr-2Mo steel
2.3.3.1 Introduction
The low carbon-9Cr-2Mo steel (STBA 27, SFVAF27) has been used as reheater and superheater tubes at
high-temperature boilers in Japan [1]. In order to obtain higher creep rupture strength and higher
oxidation resistance at around 600 °C than T22 (2 1/4 Cr-1Mo steel), a low carbon-9Cr-2Mo steel has
been developed through the considerations shown in Fig. 151. At a servicing temperature of 600 °C, 7 to
9 % Cr is needed to prevent oxidation. From the point that a steel for high-temperature use must have
enough oxidation resistance and must contain polygonal ferrite, 9 % Cr was chosen. The strengthening of
9 % Cr steel should be achieved by precipitation strengthening or solution-hardening which is effective
for the stabilization of high-temperature strength, but the strengthening by precipitation is more
remarkable than that by solid solution at high temperature. Next, heat treatment condition was considered.
Commercial steels are often used in a fully annealed condition. The annealing treatment is advantageous
for low Cr steels, but unfavorable for high Cr steels. Normalizing and tempering (NT) is usually favorable
for high Cr steels. From the microstructural viewpoint, the decrease or removal of martensite is desirable.
Alpha single phase material could be strengthened at high temperature, but coarse grain size leads to poor
toughness and decreases such practical properties as flattening, flaring and weldability. These
considerations led to the ideas of a duplex microstructure. The existence of polygonal ferrite is
advantageous for both the stabilization of high-temperature strength and the decrease of tensile strength at
room temperature. Further strengthening at high temperature was then considered. The decrease of carbon
is favorable for both obtaining (α + γ) duplex phase at the normalizing temperature and for the
improvement of weldability. Strong precipitation hardening elements, such as V, Ti and Nb, are strong
ferrite-forming elements and have a tendency to become brittle during tempering. Increase in these kinds
of elements accelerates weld cracking and makes it difficult to control ferrite content. From these
considerations, 9Cr-Mo steel with simply higher Mo content, which is easy to fabricate, was proposed.
Effect of chemical
composition
Servicing temperature:
600oC
Oxidation resistivity
High temperature
strengthening
Ti, Nb, V
Mo (W)
(Poor weldability)
(No deleterious
effect)
9%Cr
Normalizing and temper
(Excessive strength and
poor weldability)
(Formation of proeutectoid
Ferrite is depressed)
Annealing
(Ferrite + Coarse carbide)
1) Modification of strength and
Decrease of high temperature
weldability
strength and poor weldability
2) Increase and stabilization
of high temperature strength
Low C-9Cr-2Mo
α+γ at Normalizing
temperature
α single phase at
Normalizing temperature
Fig. 151. Alloy design philosophy for development of low
carbon-9Cr-2Mo steel.
(Poor toughness and weldability)
2.3.3.2 Material standards, chemical and tensile requirements
2.3.3.2.1 9Cr-2Mo steel tube for boilers and heat exchangers
Table 95. Chemical requirements for 9Cr-2Mo steel tube (JIS STBA 27)
Chemical composition [wt%]
Stan- Desigdard nation
C
Si
Mn
P
S
Cr
Mo
1)
STBA 27 2) <0.08 <0.50 0.3-0.7 <0.030 <0.030 8.00-10.00 1.80-2.20
Yield strength
[MPa]
>295
1) The Ministerial Ordinance for Providing the Technical Standard for Thermal Power Generating Facilities-1997 in Japan
2) Designation in the Ministerial Ordinance for Providing the Technical Standard for Thermal Power Generating Facilities-1997 in Japan.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 139]
2.3.3 9Cr-2Mo steel
135
Table 96. Chemical requirements for 9Cr-2Mo steel plate (JIS SFVAF 27)
Chemical composition [wt%]
Stan- Designation
dard
C
Si
Mn
P
S
Cr
Mo
1)
SFVAF 27 2) <0.08 <0.50 0.3-0.7 <0.030 <0.030 8.00-10.00 1.80-2.20
Yield strength
[MPa]
>295
1) The Ministerial Ordinance for Providing the Technical Standard for Thermal Power Generating Facilities-1997 in Japan
2) Designation in the Ministerial Ordinance for Providing the Technical Standard for Thermal Power Generating Facilities-1997 in Japan
2.3.3.3 Data sources for 9Cr-2Mo steel tubes (JIS STBA 27) and plates (JIS SFVAF 27)
Information of fact on data for 9Cr-2Mo steel tubes and plates can be obtained from NRIM Creep Data
Sheet, No.46 (1997) [2].
2.3.3.4 Creep rupture data for 9Cr-2Mo steel tubes for power boilers and 9Cr-2Mo steel plates for
power plants
2.3.3.4.1 Creep rupture data for 9Cr-2Mo steel tubes (STBA27) and plates (SFVAF27)
Creep rupture data and optical micrographs of as-received specimens have been obtained for 1 heat for
tubes and 1 heat for plates [2]. The details of steel tube and plate production, processing, thermal history,
austenite grain size number, Rockwell hardness and volume fraction of non-metallic inclusions before
creep test, the chemical compositions, the 0.2% proof stress and ultimate tensile strength data at high
temperature are also available in [2]. The tensile and creep specimens had a geometry of 6 mm in gauge
diameter and 32 mm in gauge length for the 9Cr-2Mo steel tubes, STBA 27, and 10 mm in gauge
diameter and 52 mm in gauge length for the 9Cr-2Mo steel plates, SFVAF 27.
Fig. 152 and 153 show the 0.2% proof stress and tensile strength obtained by short-time tensile tests
for the 9Cr-2Mo steels STBA 27 and SFVAF 27, respectively, between room temperature and 750 °C.
Fig. 154 and 155 show creep rupture data for 9Cr-2Mo steel tubes and plates, respectively. The
estimation of 105 h creep rupture strength is not made in [2], because the test duration was too short.
Tensile strength
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
400
300
400
300
200
200
100
100
0
0
100
200
300 400 500
600
700
800
Test temperature (℃)
Fig. 152. Short-time tensile properties of 9Cr-2Mo steel, STBA 27
Landolt-Börnstein
New Series VIII/2B
0
0
100
200
300 400 500
600
Test temperature (℃)
700
800
136
2.3 High Cr steels
Tensile strength
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
600
300
300
200
200
100
100
0
0
100
200
300 400 500
600
700
0
800
0
100
200
Test temperature (℃)
300 400 500
600
700
800
Test temperature (℃)
Fig. 153. Short-time tensile properties of 9Cr-2Mo steel, SFVAF 27
400
o
500 C
Stress ( MPa )
o
550 C
o
600 C
100
o
650 C
o
700 C
10
1
10
10
2
10
3
10
4
10
5
10
Fig. 154. Creep rupture strength
data for 9Cr-2Mo steel tubes,
STBA 27.
6
T ime t o rup ture ( h )
4 00
o
45 0 C
o
Stress ( MPa )
50 0 C
o
55 0 C
o
60 0 C
1 00
o
65 0 C
10
0
10
1
10
10
2
10
3
10
4
10
5
10
6
Fig. 155. Creep rupture strength
data for 9Cr-2Mo steel plates,
SFVAF 27.
T ime to rupture ( h )
Landolt-Börnstein
New Series VIII/2B
Ref. p. 139]
2.3.3 9Cr-2Mo steel
137
2.3.3.5 Creep behavior and microstructure of 9Cr-2Mo steel
2.3.3.5.1 Creep and creep rupture data for 9Cr-2Mo steel tubes and plates: Sumitomo data
Fig. 156 shows stress vs. time to rupture data at 450, 500, 550, 600, 650, 675 and 700 °C for 5 heats of
9Cr-2Mo steel tubes and plates, satisfying the chemical requirements for STBA 27 and SFVAF 27 [3].
The total elongation of creep ruptured specimens is more than 20 %, which is enough ductility for
practical use.
50
40
30
450 °C
500 °C
Stress [kgf / mm 2 ]
20
550 °C
10
8
6
600 °C
4
650 °C
675 °C
2
Fig. 156. Creep rupture strength
data for 9Cr-2Mo steel tubes and
plates.
700 °C
1
10
10 2
10 3
Rupture time [h]
10 4
10 5
Fig. 157 compares stress vs. time to rupture data for 9Cr-2Mo steel at 600 °C with those for 2.25Cr-1Mo
(T22) and 9Cr-1Mo (T9) steels. The 2.25Cr-1Mo steel exhibits a gentle slope at short times, but at long
times the slope becomes steep. In comparison, the stress versus time to rupture curves of the 9Cr-1Mo
and 9Cr-2Mo steels are nearly straight. Fig. 158 shows creep curves of 9Cr-2Mo steel plates at 550 °C.
The 9Cr-2Mo steel exhibits larger primary creep strain but lower creep rate in the secondary or steady
state creep region than those in 2.25Cr-1Mo steel.
30
9Cr-2Mo
2 1/4 Cr-1Mo
2 1/4 Cr-1 Mo
9 Cr-2 Mo
20
Strain [%]
Stress [kgf / mm 2 ]
20
9 Cr-1 Mo
10
8
15.0 kg /mm
2
22.0 kg /mm 2
20.0 kg /mm 2
18.0 kg /mm 2
10
16.0 kg /mm 2
12.0 kg /mm 2
6
creep ruptured at 600 °C (1112 °F)
4
10 2
10 3
Time to rupture [h]
10 4
Fig. 157. Comparison of stress versus time to rupture
curves of 2.25Cr-1Mo, 9Cr-1Mo and 9Cr-2Mo steels at
600 °C.
Landolt-Börnstein
New Series VIII/2B
0
1000
Time [h]
2000
Fig. 158. Creep curves of 9Cr-2Mo steel plates at
550 °C (solid lines), comparing with those of 2.25Cr1Mo steel (dotted lines).
138
2.3 High Cr steels
Based on the creep curves in Fig. 158, the time to reach 1 % creep strain and the time to onset of tertiary
or acceleration creep were evaluated. This is shown in Fig. 159. The time to reach tertiary creep after
reaching 1 % creep strain is much longer for 9Cr-2Mo steel than for 2.25Cr-1Mo and 18Cr-8Ni-Ti steels.
Time to 1 % creep strain [h]
10 4
9Cr-2Mo
2 1/4 Cr-1Mo
18-8-Ti
10 3
10 2
10
10 2
10 4
10 3
Time to onset of tertiary creep [h ]
Fig. 159. Relationship between time to reach 1 % creep
strain and time to onset of tertiary or acceleration creep
for 9Cr-2Mo steel at 550 °C, comparing with that for
2.25Cr-1Mo and 18Cr-8Ni-Ti steels.
The allowable stress based on ASME Code Section III, Case Interpretation 1592, was calculated from the
high-temperature tensile strength and creep rupture strength. Fig. 160 shows the ASME allowable stress
intensity value, defined as Smt value at 105 h, for 9Cr-2Mo steel together with those for various steels.
The allowable stress of 9Cr-2Mo steel is 25 - 50 % higher than that of 2.25Cr-1Mo steel (T22) and 30 60 % higher than that of 9Cr-1Mo steel (T9) at temperatures higher than 550 °C.
Stress intensity value [kg /mm2 ]
20
15
5
Smt [10 h]
10
21/4 Cr-1 Mo
9 Cr-2 Mo
Type 304
Type 316
Alloy 800 H
5
0
300
400
500
600
Temperature [°C ]
700
800
Fig. 160. Comparison of allowable stress intensity value
for various steels based on ASME case 1592.
2.3.3.5.2 Microstructure and strengthening factors of 9Cr-2Mo steel
The microstructure of 9Cr-2Mo steel consists of tempered martensite and polygonal ferrite with fine
M23C6 carbides [1]. Spheroidization of needle-like carbides and the formation of net-like carbides occur
in tempered martensite during creep at 600 °C. The precipitation behavior of carbides can be detected by
the measurement of Cr or Mo as carbides. Analysis of electrolytically extracted residues revealed that the
amount of Mo as carbides increases more slowly than that in normalized and tempered 2.25Cr-1Mo steel
Landolt-Börnstein
New Series VIII/2B
Ref. p. 139]
2.3.3 9Cr-2Mo steel
139
(T22) and annealed 9Cr-1Mo steel (T9). Half of the total Mo remains in solution in the 9Cr-2Mo steel.
From microstructural observations, the strengthening factors of the 9Cr-2Mo steel are summarized as
follows. Positive strengthening factors are: the precipitation of needle-like carbides within ferrite, the
formation of net-like carbides and slow growth of globular carbides within tempered martensite. Factors
that contribute to the stabilization of a microstructure are: the existence of polygonal ferrite, 1 % of Mo in
solution. The strengthening of the 9Cr-2Mo steel is not ascribed to only one strengthening factor, but to
the summing up of the aforementioned factors.
2.3.3.6 References
[1] Yukitoshi, T., Nishida, K., Oda, T. and Daikoku, T.: Journal of Pressure Vessel Technology,
Transactions of the ASME, (1976), 173 - 178.
[2] National Research Institute for Metals (NRIM) Creep Data Sheet, No.46, (1998).
[3] Yukitoshi, T., Yoshikawa, K., Tokimasa, K., Shida, Y. and Inaba, Y.: Tetsu-To-Hagane, 65 (1979),
876 - 885.
Landolt-Börnstein
New Series VIII/2B
140
2.3 High Cr steels
2.3.4 9Cr-0.5Mo-1.8W-V-Nb-B steel
2.3.4.1 Introduction
9Cr-0.5Mo-1.8W-V-Nb-B steel is used for heat exchangers and piping systems in Ultra Super Critical
steam conditioned thermal power plants [1]. In order to achieve higher creep rupture strength than that of
Mod. 9Cr steel, Gr. 91 of ASME standard, tungsten was replaced by molybdenum improving the creep
rupture strength of the ferritic heat resistant steel designated as KA-STBA29 for METI and as Gr. 92 for
ASME. Other alloying elements, niobium, vanadium and nitrogen, have also been optimized with respect
to the creep properties through precipitation strengthening by fine carbonitride distribution in the lath
interior [2]. Boron stabilizes the tempered martensitic sub-grain structure during creep deformation by
segregation at the lath boundaries and the surface of M23C6 type carbides [3].
2.3.4.2 Material standards, chemical and tensile requirements
2.3.4.2.1 9Cr-0.5Mo-1.8W-V-Nb-B steel tubes for heat exchanger or boiler applications
Table 97. Chemical composition of 9Cr-0.5Mo-1.8W-V-Nb steel tubes; KA-STBA29 and
213 T92 [4]
Chemical composition [wt%]
Standards
Designation
C
Si
Mn
P
S
Cr
Mo
0.07
0.30
8.50
0.30
≤
≤
≤
METI
KA-STBA29
0.13
0.60
0.60
0.50
0.020 0.010 9.50
8.50
0.30
0.30
0.07
≤
≤
≤
ASTM
Gr. T92
0.60
0.60
0.13
0.020 0.010 9.50
0.50
Standards
Designation
METI
KA-STBA29
ASTM
Gr. T92
W
1.50
2.00
1.50
2.00
Ni
≤
0.40
≤
0.40
Chemical composition [wt %]
V
Nb
N
Al
0.15
0.04
0.030 ≤
0.25
0.09
0.070 0.04
0.15
0.04
0.03
≤
0.25
0.09
0.07
0.04
ASTM SAStd. No.
SA-213
Std. No.
B
0.001
0.006
0.001
0.006
SA-335
2.3.4.2.2 9Cr-0.5Mo-1.8W-V-Nb-B steel pipes for steam conductors or boiler piping application
Table 98. Chemical composition of 9Cr-0.5Mo-1.8W-V-Nb steel pipes; KA-STPA29 and
335 P92 [5]
Chemical composition [wt%]
Standards
Designation
C
Si
Mn
P
S
Cr
Mo
8.50
0.30
0.30
0.07
≤
≤
≤
METI
KA-STPA29
0.60
0.60
0.13
0.020 0.010 9.50
0.50
8.50
0.30
0.30
0.07
≤
≤
≤
ASTM
Gr. P92
0.60
0.60
0.13
0.020 0.010 9.50
0.50
Standards
Designation
METI
KA-STPA29
ASTM
Gr. P92
W
1.50
2.00
1.50
2.00
Ni
≤
0.40
≤
0.40
Chemical composition [wt%]
V
Nb
N
Al
0.15
0.04
0.030 ≤
0.25
0.09
0.070 0.04
0.15
0.04
0.03
≤
0.25
0.09
0.07
0.04
ASTM SAStd. No.
SA-335
Std. No.
B
0.001
0.006
0.001
0.006
SA-335
Landolt-Börnstein
New Series VIII/2B
Ref. p. 143]
2.3.4 9Cr-0.5Mo-1.8W-V-Nb-B steel
141
2.3.4.3 Creep properties of 9Cr-0.5Mo-1.8W-V-Nb steel tubes
[6] contains creep data of 9Cr-0.5Mo-1.8W-V-Nb-B steel tubes, namely rupture data, minimum creep
rate, rupture elongation and reduction of area of cross section.
2.3.4.3.1 Creep rupture data of 9Cr-0.5Mo-1.8W-V-Nb steel tubes
Fig. 161 shows the creep rupture data of 9Cr-0.5Mo-1.8W-V-Nb-B steel tubes of 7 heats [6]. Creep test
duration is still continuing and is at about 55,000 h. All the rupture data satisfy the ASME Gr. 92 standard
requirement.
1000
T ube
Stress (MPa)
550 ℃
600 ℃
100
650 ℃
750 ℃
700 ℃
Fig. 161. Creep rupture strength
data of KA-STBA29; [6].
10
1
10
100
1000
10000
100000
T ime (h)
2.3.4.3.2 Time-Temperature-Parametric prognostication of creep rupture strength
Fig. 162 shows the Larson Miller Parametric plots of the rupture data based on [6]. Creep rupture curve
regression by a cubic expressionpredicts the creep rupture strength for times longer than that of the
experiment between 550 °C and 750 °C, Fig. 163.
1000
1000
Stress (MPa)
Stress (MPa)
550 ﾟC
100
100
600 ﾟC
650 ﾟC
750 ﾟC
10
----- Average
10
25000
27000
29000
700 ﾟC
1
31000
33000
35000
37000
39000
10
100
1000
10000
100000
1000000
Rupture T ime (h)
Larson-Miller-parameter TK (32.896118 + log tr )
Fig. 162. Master rupture curve by Larson-Miller
parameter method for 9Cr-0.5Mo-1.8W-V-Nb-B steel
tubes.
Landolt-Börnstein
New Series VIII/2B
Fig. 163. Estimated creep rupture curves of 9Cr-0.5Mo1.8W-V-Nb-B steel tubes.
142
2.3 High Cr steels
2.3.4.4 Creep properties of 9Cr-0.5Mo-1.8W-V-Nb steel pipes
[6] contains creep data of 9Cr-0.5Mo-1.8W-V-Nb-B steel pipes, namely rupture data, minimum creep
rate, rupture elongation and reduction of area of cross section.
2.3.4.4.1 Creep rupture data of 9Cr-0.5Mo-1.8W-V-Nb steel pipes
Fig. 164 shows creep rupture data of 9Cr-0.5Mo-1.8W-V-Nb-B steel pipes of 9 heats [6]. Creep test
duration is still continuing and is at about 70,000 h. All the rupture data satisfy the ASME Gr. 92 standard
requirement.
1000
Pipe
Stress (MPa)
550 ℃
600 ℃
100
650 ℃
700 ℃
750 ℃
Fig. 164.
Creep rupture
strength data of KA-STPA29;
[6].
10
1
10
100
1000
10000
100000
T ime (h)
2.3.4.4.2 Time-Temperature-Parametric prognostication of the creep rupture strength
Fig. 165 shows the Larson Miller Parametric plots of the rupture data based on [6]. Creep rupture curve
regression by a cubic expression predicts the creep rupture strength for times longer than that of the
experiment between 550 °C and 750 °C [6], Fig. 166.
1000
550 ﾟC
Stress (MPa)
Stress (MPa)
1000
100
100
600 ﾟC
650 ﾟC
----10
25000
750 ﾟC
30000
35000
40000
45000
Larson-Miller-parameter TK (35.277449 + log tr )
Fig. 165. Master rupture curve by Larson-Miller
parameter method for 9Cr-0.5Mo-1.8W-V-Nb steel
pipes.
700 ﾟC
10
1
10
100
1000
10000
100000
1000000
Rupture T ime (h)
Fig. 166.
Estimated creep rupture curves of
9Cr-0.5Mo-1.8W-V-Nb steel pipes.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 143]
2.3.4 9Cr-0.5Mo-1.8W-V-Nb-B steel
143
2.3.4.5 Effect of tungsten content on creep properties of 9Cr-0.5Mo-1.8W-V-Nb steel
Precipitated W [mass %]
Precipitated W [mass %]
Precipitated W [mass %]
Optimized tungsten content in 9Cr-0.5Mo-1.8W-V-Nb steel increases the creep rupture strength at high
temperatures [7]. Tungsten delays the microstructure evolution by solid solution dragging in the same
manner as molybdenum. An excess of tungsten in thermodynamic equilibrium at service temperature
precipitates as Fe2W type intermetallic compound, Laves phase. Fig. 167 shows the tungsten precipitation
during aging [8]. After saturation of precipitation in amount, Laves phase ripens at the sub-grain
boundary, and possibly affects the microstructure evolution delaying the dynamic recrystallization.
1.5
1.26 %
600 °C
1.2
0.9
0.6
0.3
0
1.5
650 °C
1.2
0.99 %
0.9
0.6
0.3
0
1.5
700 °C
1.2
0.9
0.58 %
0.6
Fig. 167.
time.
0.3
0
1
10
10 3
10 2
Aging time [h]
10 4
Tungsten precipitation behavior with aging
10 5
2.3.4.6 References
[1] Fujita, T.: COST-EPRI Workshop, Shaffhausen (1996).
[2] Masumoto, H., Sakakibara, M., Takahashi, T., and Fujita, T.: EPRI, 1st. Conf. Improved Coal-Fired
Power Plants, Palo Alto 5 (1986) 205.
[3] Naoi, H., Mimura, H., Ohgami M., Morimoto, H., Tanaka, T., Yagaki, Y., and Fujita, T.: Proc.
EPRI/National Power Conf., New Steel for Advanced Plant up to 620°C, London, May 11 (1995), 1.
[4] ASTM Standard: A213/A213M-99a (2001).
[5] ASTM Standard: A335/A335M (2001).
[6] Muraki, T.: Private communication “Nippon Steel Creep Database” (2000)
[7] Hasegawa, Y., Muraki, T., and Ohgami, M.: 123rd committee heat resistant materials and alloys, 39,
No. 3 (1998) 275.
[8] Hasegawa, Y., Muraki, T., Ohgami, M., and Mimura, H.: Proceedings of the 8th International
Conference on Creep and Fracture of Engineering Materials and Structures (1999) 427.
Landolt-Börnstein
New Series VIII/2B
144
2.3 High Cr steels
2.3.5 9Cr-1Mo-1W-V-Nb-N steel
2.3.5.1 Introduction
9Cr-1Mo-1W-V-Nb-N steel (X11CrMoWVNb9-1-1) is used for tubing, headers and piping, but also for
large forgings. The maximum service temperature for tubes and pipes is limited to 625 °C. The steel was
developed in Europe in parallel to the grade P92 and is also well known under the designation E911.
The material shows a pure martensitic microstructure after tempering. The steel has a good weldability.
Preheating up to 150 °C at least is recommended, during welding the temperature should not exceed
350 °C. After welding it is essential for the steel to cool down to a temperature below 100 °C, in order to
allow the complete transformation into martensite. The above mentioned parameters depend on the type
and thickness of the component to be welded. Thin structures may be welded below 150 °C and also
allowed to cool down at room temperature. Post heat treatment should be done at 740 - 770 °C. The
holdtime at PWHT temperature depends on the thickness of the material.
2.3.5.2 Material standards, chemical and tensile requirements
2.3.5.2.1 X11CrMoWVNb9-1-1 for seamless tubes
Table 99. Heat treatment of X11CrMoWVNb9-1-1 seamless tubes for pressure purposes; VdTÜV
Werkstoffblatt 522:2001
Quenching
1040 - 1080 °C / Air
VdTÜV
Werkstoffblatt 522/2
Tempering
750 - 780 °C/ Air
Table 100. Chemical requirements of X11CrMoWVNb9-1-1; VdTÜV Werkstoffblatt 522/2:2001
Chemical composition [wt%]
Designation
steel number C
Si
Mn P
S
Cr Ni
Mo W
Nb V
Al
N
B
8.50 0.10 0.90 0.90 0.06 0.18 0.050 0.0005
≤
X11CrMoW 0.09 0.10 0.30 ≤
0.13 0.50 0.60 0.020 0.010 9.50 0.40 1.10 1.10 0.10 0.25 0.040 0.900 0.005
VNb9-1-1
1.4905
Table 101. Tensile requirements of X11CrMoWVNb9-1-1 at room temperature; VdTÜV Werkstoffblatt
522/2:2001
Tensile requirements
Standard
Designation/steel number
Yield
Tensile
Elongation
Impact energy
strength strength
after fracture
Charpy-V-testpiece
T
Q
T
Q
[MPa]
[MPa]
[%]
[%]
[J]
[J]
VdTÜV
X11CrMoWVNb9-1-1
620 - 850 ≥19
≥450
≥17
≥68
≥41
Werkstoffblatt 1.4905
522/2
T longitudinal, Q transverse direction
2.3.5.2.2 X11CrMoWVNb9-1-1 for forgings and rolled or forged bars for pressure purposes
Table 102. Heat treatment of X11CrMoWVNb9-1-1 forgings and rolled or forged bars for pressure
purposes; VdTÜV Werkstoffblatt 522/3:2001
Quenching
1040 - 1080 °C / Oil
VdTÜV
Werkstoffblatt 522/2
Tempering
750 - 780 °C/ Air
Landolt-Börnstein
New Series VIII/2B
Ref. p. 149]
2.3.5 9Cr-1Mo-1W-V-Nb-N steel
145
Table 103. Chemical requirements of X11CrMoWVNb9-1-1; VdTÜV Werkstoffblatt 522/3:2001
Chemical composition [wt%]
Designation
steel number C
Si
Mn P
S
Cr Ni
Mo W
Nb V
Al
N
B
8.50 0.10 0.90 0.90 0.06 0.18 0.050 0.0005
X11CrMoW 0.09 0.10 0.30 0.13 0.50 0.60 0.020 0.010 9.50 0.40 1.10 1.10 0.10 0.25 0.040 0.900 0.005
VNb9-1-1
1.4905
Table 104. Tensile requirements of X11CrMoWVNb9-1-1 at room temperature; VdTÜV Werkstoffblatt
522/3:2001
Standard
Designation/
steel number
VdTÜV
X11CrMoWVNb9-1-1
Werkstoffblatt 522/2 1.4905
Tensile requirements
Range of
Yield
Tensile Elongation
significant strength strength after fracture
dimensions
T
Q
[%] [%]
[mm]
[MPa]
[MPa]
620 ≤500
≥450
≥19 ≥17
850
Impact energy
Charpy-V-testpiece
T
Q
[J]
[J]
≥68
≥41
T longitudinal, Q transverse direction
Fig. 168. Microstructure of
X11CrMoWVNb9-1-1 (as received state).
2.3.5.3 Mechanical properties of X11CrMoWVNb9-1-1 seamless tubes, forgings
1000
minimum values tensile strength VdtÜV 522
minimum values yield strength VdtÜV 522
Stress σ (MPa)
800
600
400
200
0
0
100
200
300
400
Temperature T (°C)
Landolt-Börnstein
New Series VIII/2B
500
600
700
Fig. 169. High temperature tensile strength
and yield strength of X11CrMoWVNb9-1-1
according to VdTÜV 522: 2001 [1]
146
2.3 High Cr steels
800
E911 UTS [2]
E911 YS [2]
E911 UTS [8]
E911 YS [8]
700
Stress σ (MPa)
600
500
400
300
200
100
0
0
100
200
300
400
500
600
700
800
Fig. 170. High temperature tensile strength
(UTS) and yield strength (YS) of
X11CrMoWVNb9-1-1 (E911) [2], [8].
Temperature T (°C)
2.3.5.4 Creep properties of X11CrMoWVNb9-1-1
2.3.5.4.1 Creep strength of X11CrMoWVNb9-1-1 seamless tubes and forgings
1000
Stress σ (MPa)
Tubes
Welded pipes
Pipes
100
10
100
1000
10000
100000
Fig. 171. Creep rupture strength data
of E911 (X11CrMoWVNb9-1-1) at
625 °C [2]. The data of tubes are
slightly above the pipes. Short term
tests with cross weld specimens of
pipes are at the lower band of the
parent material in the short term test
duration.
Time to rupture tR (h)
1000
Stress σ (MPa)
P91
E911
100
Fig. 172. Creep rupture strength data
of E911 (X11CrMoWVNb9-1-1) and
P91 (both pipes) at 650 °C [2]. E911
material shows a factor 3 to 5 longer
rupture time at the same constant
stress in comparison to P91 steel.
10
10
100
1000
10000
100000
Time to rupture tR (MPa)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 149]
2.3.5 9Cr-1Mo-1W-V-Nb-N steel
147
Stress σ (MPa)
1000
610 °C, heat A, broken
640 °C, heat A, broken
550 °C, heat B, broken
600 °C, heat B, broken
600 °C, heat B, unbroken
650 °C, heat B, broken
650 °C, heat B, unbroken
550 °C, VdTÜV 522
600°C, VdTÜV 522
610 °C, VdTÜV 522
640 °C, VdTÜV 522
650 °C, VdTÜV 522
550 °C, broken [8]
550 °C, unbroken [8]
600 °C, broken [8]
600 °C, unbroken [8]
650 °C, broken [8]
650 °C, unbroken [8]
100
10
10
100
1000
10000
100000
1000000
Fig. 173. Creep rupture strength
data of steel grade E911
(X11CrMoWVNb9-1-1) in dependence of temperature (3 heats,
tubes) [3], [8] and average creep
rupture strength values indicated
in VdTÜV 522/2.
Stress σ (MPa)
Test duration t (h)
Fig. 174. Creep rupture strength
data
of
steel
grade
E911
(X11CrMoWVNb9-1-1) of parent
material and cross weld specimens
at 600 °C [3], [8] and average creep
rupture strength values indicated in
VdTÜV
522/2.
Cross
weld
specimens show lower creep
strength in comparison to the parent
material, caused by failure in the
heat affected zone.
100
Crossweld, MAW, broken
Crossweld, MAW, unbroken
Parent material, broken
Parent material, unbroken
Crossweld, SAW, broken
600 °C, VdTÜV, parent material
Parent material, broken [8]
Parent material, unbroken [8]
10
100
1000
10000
100000
Stress σ (MPa)
Test duration t (h)
100
Crossweld, MAW, broken
Crossweld, MAW, unbroken
Parent material, broken
Parent material, unbroken
Crossweld, SAW, broken
650 °C, VdTÜV, parent material
Weld material, broken [8]
Crossweld, broken [8]
10
100
1000
10000
Test duration t (h)
Landolt-Börnstein
New Series VIII/2B
100000
Fig. 175. Creep rupture strength
data of the steel grade E911
(X11CrMoWVNb9-1-1) of parent
material and cross weld specimens at
650 °C [3], [8] and average creep
rupture strength values indicated in
VdTÜV
522/2.
Cross
weld
specimens show lower creep
strength in comparison to the parent
material, caused by failure in the
heat affected zone.
148
2.3 High Cr steels
2.3.5.4.2 Fact creep data of X11CrMoWVNb9-1-1 tubes
Information of fact on creep data for X11CrMoWVNb9-1-1 tubes and forgings is given in [4, 6, 7].
2.3.5.4.3 Microstructural change
The typical microstructure of X11CrMoWVNb9-1-1 is a martensitic structure. A high dislocation density
and relatively fine carbide precipitates at the martensite lath boundaries can be observed. In the initial
state Nb and NbV-rich MX particles as well as M23C6 carbides can be observed. The nucleation of Laves
phase and Z phase can be observed after creep at 600 °C. In this context a change in the chemical
composition of both M23C6 carbides as well as MX carbides could be observed [5].
2.3.5.4.4 Creep crack growth
Creep crack growth may occur at defects in components operated in the high temperature regime. The
assessment of crack growth at defects is based on fracture mechanics and the relevant data. The crack
growth rate was measured using fracture mechanic specimens (CT, Cs -side grooved CT, D) with
different ratios of crack start length a0 to specimen width W. The stress situation at crack tip was
calculated using the creep fracture mechanics parameter C*.
1
10 -7
a [mm/h]
10 -2
10 -3
230
(1035)
282
282 (1799)
(1271)
178
(304)
220 175
(674)
(305)
10 -4
10 -5
10 -6
10 -5
sn0 = 204 MPa
-8
(K l0 = 688 N/mm 3/2 ) 10
320
(1443)
10 -9
178
(901)
10 -10
320
282
176 (951)
179 (862) (709)
(679)
140
(454)
Cs25 Cs50 CT100
a0 / W 0.55 0.55 0.55
10 -11
D30 D60
D15
a0 / W 0.25 0.45 0.2 0.1 0.2 10 -12
10 -4
10 -3
10 -2
C * [N/mmh]
10 -1
1
10
a [m/s]
10
-1
Fig. 176. Crack growth rate in
dependence on the parameter C*
(based
on
the
load
line
displacement rate) for different
specimens types and sizes for
E911 (forgings). Cs25, Cs50,
CT100 – CT specimens with
thickness 25 and 50 mm (side
grooved), 100 mm (not side
grooved), D15, D30, D60– double
edge notched specimens, thickness
15 mm; σn0 nominal stress, KI0
stress intensity factor at the
beginning of the test [6].
2.3.5.4.5 Estimated long term creep rupture strength
Table 105. Average values of creep rupture strength of X11CrMoWVNb9-1-1 for seamless tubes. Long
term data is partly based on extended time and stress extrapolations; [1].
Time to rupture
Temperature
[°C]
[104 h]
[105 h]
Average creep rupture strength [MPa]
480
322
288
490
305
271
500
288
255
510
271
239
520
255
223
Landolt-Börnstein
New Series VIII/2B
Ref. p. 149]
Temperature
[°C]
530
540
550
560
570
580
590
600
610
620
630
640
650
2.3.5 9Cr-1Mo-1W-V-Nb-N steel
149
Time to rupture
[104 h]
[105 h]
Average creep rupture strength [MPa]
239
208
224
193
21
182
197
166
182
150
167
135
154
121
140
108
128
95
115
83
104
72
93
62
82
53
2.3.5.5 References
[1] VdTÜV-Werkstoffblatt 522 09 (2001): Warmfester Stahl X11CrMoWVNb9-1-1. TÜV Verlag
GmbH, Postfach 9030 60, 51123 Köln.
[2] Gianfrancesco, A. Di., and Merckling, G.: The Italian effort to the development of 9 to 12 %
Chromium steels for power plant applications. International Colloquium on the occasion of the 50th
anniversary of the German Creep Committee, November 25 (1999) Düsseldorf. Verein Deutscher
Eisenhüttenleute VDEh.
[3] Klenk, A., Maile, K., Theofel, H., and Husemann, R.-A.: Long term behaviour of weldments of
modern power plant steels. Creep 7, Proceedings of the 7th International Conference on Creep and
Fatigue at elevated Temperatures, Japan Society of Mechanical Engineers Tokyo, Japan (2001) 8791.
[4] Husemann, R.U.: Abschlussbericht und Zusammenfassung der Ergebnisse des FDBR-/VGBForschungsvorhabens „Qualifizierung von Werkstoffen zum Einsatz in Dampferzeugeranlagen mit
erhöhten Temperaturen“, AVIF-Vorhaben A77 (1999). Wirtschaftsverband Stahlbau und
Energietechnik e. V., Düsseldorf.
[5] Maile, K., Klenk, A., and Zies, G.: Determination of Microstructural Parameters for 9 % Cr Steels.
Proceedings of the 8th Japanese-German Joint Seminar on Structural Integrity and NDE in Power
Engineering, Tokyo (2001) 511-518.
[6] Berger, C., Granacher,J., Kostenko,J., Roos, E., Maile, K., and Schellenberg, G.: Kriechrissverhalten
ausgewählter Kraftwerksbaustähle in erweitertem, praxisnahem Parameterbereich, Abschlussbericht
des AVIF-Vorhabens Nr. A78 des IfW der TU-Darmstadt und der MPA der Universität Stuttgart
(1999), Forschungskuratorium Maschinenbau e. V., Frankfurt.
[7] Klenk, A., Maile, K.: Nachweis der Langzeiteigenschaften von Schweißverbindungen moderner
Stähle für den Einsatz in Dampferzeugern im Bereich bis 620 °C. Abschlussbericht AVIF Vorhaben
A129 der MPA Stuttgart, 2002. Wirtschaftsverband Stahlbau und Energietechnik e.V., Düsseldorf.
[8] Bendick, W.: Neue Entwicklungen für warmfeste Rohre im Kraftwerksbau. 3R international (40),
Heft 5/2001 264-268.
Landolt-Börnstein
New Series VIII/2B
150
2.3 High Cr steels
2.3.6 12Cr steel
2.3.6.1 Introduction
12Cr stainless steels are used as wrought bars, plates, sheets, strips, wire rods, billets and forgings. It is
hardenable corrosion and heat resistant martensitic chromium steel. Microstructure and stability at
elevated temperatures are strongly influenced by heat treatment condition. Changes in microstructure
during creep exposure and creep strength properties have been investigated, in conjunction with creep
deformation behavior.
2.3.6.2 Materials standards, chemical and tensile requirements
2.3.6.2.1 12Cr steel
Table 106. Chemical requirements of 12Cr steel; JIS SUS403 and ASTM 403
Chemical composition [wt%]
StanDesigdards
nation
C
Si
Mn
P
S
Cr
11.50JIS
SUS403-B ≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
13.00
SUS40311.50JIS
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
HP
13.00
SUS40311.50JIS
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
CP
13.00
SUS40311.50JIS
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
WR
13.00
11.50ASTM 403
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
13.00
11.50ASTM 403
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
13.00
11.50ASTM 403
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
13.00
11.50ASTM 403
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
13.00
11.50ASTM 403
≤0.15
≤0.50
≤1.00
≤0.040 ≤0.030
13.00
Std. No
Ni
≤0.60
G4303
≤0.60
G4304
≤0.60
G4305
≤0.60
G4308
≤0.60
A176
-
A276
-
A314
-
A473
-
A479
Table 107. Product forms of 12Cr steel; JIS SUS403 and ASTM 403
Standards
Designation
Std. No
Product form
JIS
SUS403-B
G4303
bar
JIS
SUS403-HP
G4304
plate, sheet and strip
JIS
SUS403-CP
G4305
plate, sheet and strip
JIS
SUS403-WR
G4308
wire rod
ASTM
403
A176
plate, sheet and strip
ASTM
403
A276
bar and shape
ASTM
403
A314
billet and bar for forging
ASTM
403
A473
forging
ASTM
403
A479
bar and shape
Landolt-Börnstein
New Series VIII/2B
Ref. p. 155]
2.3.6 12Cr steel
151
2.3.6.3 Creep properties of 12Cr steels
Information of fact on creep data for 12Cr stainless steels can be obtained [1] and [2].
2.3.6.3.1 Creep rupture data of 12Cr steel bars
The results of creep tests for 9 heats of JIS SUS 403-B steel bars are given in [1]. Here, data on rupture
elongation, reduction of area, minimum creep rate, time to specified strain and microstructures of asreceived materials and crept specimens are aloso given.
Creep rupture strength of 9 heats of 12Cr stainless steel bars (JIS SUS 403-B) are shown in Fig. 177.
It should be noted that the creep rupture strength of the steels shows a large scatter over the range of
tested temperatures from 450 to 600 °C. Heat-to-heat variation is caused by differences in initial
microstructure. The martensitic microstructure is strongly influenced by manufacturing condition,
especially heat treatment condition. The initial microstructure of the 12Cr stainless steel influences its
stability during creep exposure at elevated temperatures and results in large differences in creep rupture
strength, as will be described later.
500
Stress (MPa)
300
100
80
○
●
△
▲
60
40
10
450 oC
500 oC
550 oC
600 oC
n=226
-1
0
10
10
1
10
2
10
3
10
4
10
5
10
6
Fig. 177.
Creep rupture
strength data of 9 heats of
SUS 403-B steel bars; [1]. n
indicates the total number of
data points.
Time to rupture (h)
2.3.6.3.2 Creep rupture strength of 12Cr stainless steel bars
Creep rupture strength data of 9 heats of 12Cr stainless steel bars are shown in Fig. 178. It should be
noted that creep rupture strength of 12Cr stainless steel bars shows a large heat-to-heat variation.
Microstructure is strongly dependent on heat treatment conditions and influences the creep strength of
the steel. Heat treatment conditions and creep rupture strength of 3 heats of 12Cr stainless steel bars are
shown in Table 108 and Fig. 179, respectively. Differences in creep rupture strength of these 3 heats of
12Cr stainless steel bars are attributed to differences in heat treatment condition as shown in Table 108.
The lowest creep rupture strength of heat C is caused by an extremely high tempering temperature of
750 °C. Rapid decrease in creep rupture strength with increase in creep exposure of heat B is due to low
temperature before quenching, as will be explained in conjunction with microstructures later.
Landolt-Börnstein
New Series VIII/2B
2.3 High Cr steels
800
600
500
400
450 °C
475 °C
500 °C
525 °C
550 °C
575 °C
600 °C
Stress [MPa]
300
200
100
80
60
50
40
30
500
400
300
Stress (MPa)
152
200
100
90
80
70
60
50
40
30
16000
Average
n = 231
20
-26
-16
-24
-20
-18
-22
Orr-Sherby-Dorn parameter log tR -[354987/(19.1425 × TK )]
○ heat A
△ heat B
□ heat C
17000
18000
19000
20000
21000
22000
23000
Larson-Miller parameter (T(K)(21+Log tR(h)))
Fig. 179. Creep rupture strength of 3 heats of 12Cr
stainless steel bars; [3].
Fig. 178. Master rupture curve by Orr-Sherby-Dorn
parameter method for 12Cr stainless steel bars; [1].
TK = T + 273.15, T: test temperature [°C] and tR: time to
rupture [h]. n indicates the total number of data points.
Table 108. Heat treatment conditions of 3 heats of 12Cr stainless steel bars; [3].
Standard & Designation
heats
Heat treatment conditions
Forged
980 °C / 0.5 h / Oil quenching
heat A
640 °C / 2 h / Air cooling
630 °C / 2 h / Air cooling
Forged
JIS G4303 SUS403-B
heat B
950 °C / 1 h / Oil quenching
650 °C / 2 h / Air cooling
Hot rolled
heat C
970 °C / 0.5 h / Oil quenching
750 °C / 1 h / Water quenching
2.3.6.3.3 Microstructural change
The typical initial microstructure of 12Cr stainless steel is tempered martensite. Optical micrographs of a
12Cr stainless steel bar are shown in Fig. 180.
Bright field TEM images of 3 heats of 12Cr stainless steel bars in the as-received and crept conditions
at 600 °C and 61 MPa are shown in Fig. 181 [3]. A very fine tempered martensitic microstructure is
observed on heats A and B in the as-received condition. On the other hand there is a significantly
recovered microstructure even in the as-received condition on heat C, since the tempering temperature of
750 °C is extremely high, in contrast to 630 and 650 °C of heats A and B, respectively.
The tempered martensitic microstructure of heat A is very stable during creep exposure up to about
2000 h at 600 °C and 61 MPa. With increase in creep exposure time, however, recovery of tempered
martensitic microstructure proceeds very rapidly in heat B. In contrast to the homogeneous tempered
martensitic microstructure of heat A in the as-received condition, that of heat B is inhomogeneous and
contains small amounts of ferritic grains, as can be seen in Fig. 181 (f). The inhomogeneous
microstructure of heat B in the as-received condition is caused by a slightly lower temperature of 950 °C
before quenching, comparing to 980 °C of heat A.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 155]
2.3.6 12Cr steel
153
Fig. 180. Optical micrographs of as-received 12Cr stainless steel bars (etched in 4% natal); [1].
Differences in heat treatment condition strongly influence the initial microstructure and stability during
creep exposure at elevated temperatures. Different changes in microstructure during creep exposure of 3
heats of 12Cr stainless steel bars are clearly observed as changes in hardness, as shown in Fig. 182 [3].
Hardness of heat B is almost the same as that of heat A in the as-received condition, however, it decreases
rapidly with increase in creep exposure time, corresponding to changes in microstructures.
Fig. 181. Bright field TEM images of 3 heats of 12Cr stainless steel bars in the as-received and crept conditions at
600 °C and 61 MPa [3]. ti: creep exposed time before interrupting, tr: time to rupture
Landolt-Börnstein
New Series VIII/2B
154
2.3 High Cr steels
～
～～
～
200
～
～
600oC-61MPa
240
heat A
heat B
heat C
160
120
Asreceived
Fig. 182. Changes in hardness of the 3 heats of 12Cr
stainless steel bars with increase in creep exposure
time at 600 °C and 61 MPa; [3].
Ruptured
～
～
Vickers hardness (98N)
280
2
3
4
10
10
5
10
10
Time (h)
2.3.6.3.4 Creep deformation behavior and creep rupture strength
Differences in initial microstructure and its stability strongly influence creep deformation behavior. Creep
rate vs. time curves of 3 heats of 12Cr stainless steel bars at 500 °C - 177 MPa and 600 °C - 61 MPa are
shown in Fig. 183 and 184, respectively. Under the creep test condition of 500 °C - 177 MPa, differences
in creep strength of the 3 heats are clearly observed. Differences in creep strength of these 3 heats
correspond to differences in initial microstructure, as shown in Fig. 181. Creep rupture life of heat B is
about five times longer than that of heat C and that of heat A is about five times longer that that of heat B.
On the other hand, creep deformation behavior of heat B under the creep test condition of 600 °C 61 MPa is affected by rapid progress in recovery of tempered martensitic microstructure during creep
exposure, as shown in Fig. 181. Creep rate of heat B is smaller than that of heat C and almost the same as
that of heat A in the beginning of creep deformation up to about 100 h. However, heat B shows minimum
creep rate and onset of accelerating creep stage after about 200 - 300 h, which is significantly earlier than
for heat A (about 3,000 h). Consequently, difference in creep rupture life of heats B and C is very small in
comparison with the difference observed under the creep test condition of 500 °C-177 MPa. Such rapid
acceleration of creep rate for heat B is caused by rapid progress in recovery of tempered martensitic
microstructure, as shown in Fig. 181.
Differences in initial microstructure strongly influence short-term creep strength. The stability of
microstructure during creep exposure at elevated temperatures significantly affects long-term creep
strength. Very large heat-to-heat variations of creep rupture strength of 12Cr stainless steels, as shown in
Fig. 177 and 178, are caused by differences in initial microstructure and its stability at elevated
temperatures, which is strongly influenced by heat treatment condition.
-3
-3
10
10
heat C
-4
Creep rate (h )
10
10
-1
-1
Creep rate (h )
heat C
heat B
-4
-5
10
-6
10
500 C-177MPa
0
1
10
-6
o
600 C-61MPa
-7
10
2
10
3
10
4
10
heat A
10
heat A
o
10
heat B
-5
10
5
10
Time (h)
Fig. 183. Creep rate vs. time curves of 3 heats of 12Cr
stainless steel bars at 500 °C and 177 MPa; [3].
-7
10
0
10
10
1
10
2
3
10
4
10
10
5
Time (h)
Fig. 184. Creep rate vs. time curves of 3 heats of 12Cr
stainless steel bars at 600 °C and 61 MPa; [3].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 155]
2.3.6 12Cr steel
155
2.3.6.3.5 Estimated long-term creep rupture strength
The temperature dependence of 0.2% proof stress, tensile strength and creep rupture strength at 100 and
10,000 h for 9 heats of 12Cr stainless steel bars are shown in Fig. 185 [1]. Those of 0.2% proof stress,
tensile strength and creep rupture strength at 1,000 and 100,000 h for the same materials are shown in Fig.
186 [1]. Creep rupture strength curves shown in Fig. 185 and 186 were obtained by regression analysis
using the Orr-Sherby-Dorn parameter.
{{
100
80
60
50
40
30
20
400
450
500
550
Temperature [°C]
0.2% proof
stress
100 h
{
200
Tensile
strength
{
Stress [MPa]
1000
800
600
500
400
300
10000 h
600
650
Fig.
185.
Temperature
dependence of 0.2% proof stress,
tensile strength and creep rupture
strength at 100 and 10,000 h for
12Cr stainless steel bars; [1].
1000
800
200
{
Stress [MPa]
300
100
80
60
50
400
450
500
550
Temperature [°C]
{
{{
600
500
400
Tensile
strength
0.2% proof
stress
1000 h
100000 h
600
650
Fig.
186.
Temperature
dependence of 0.2% proof stress,
tensile strength and creep-rupture
strength at 1,000 and 100,000 h
for 12Cr stainless steel bars; [1].
2.3.6.4 References
[1] National Research Institute for Metals: NRIM Creep Data Sheet, No.13B, (1994).
[2] ASM International: Atlas of Creep and Stress-Rupture Curves, (1988).
[3] Kushima, H., Kimura, K., Yagi, K., and Tanaka, C.: Tetsu-to-Hagane, 81 (1995), 214-219.
Landolt-Börnstein
New Series VIII/2B
156
2.3 High Cr steels
2.3.7 12Cr-0.6Mo-0.3V-0.4Nb-N steel
2.3.7.1 Introduction
12Cr-0.6Mo-0.3V-0.4Nb-N steel (H46) is produced as forging quality bars and billets (blading bars) and
bars (rounds, squares, hexagons and flats) for use in rotating blades of steam and combustion turbines,
disks, rotor shafts and bolts. H46 was developed by William Jessop, UK in the 1950s for hightemperature applications up to 650 °C through modification of H40 (3Cr-0.5Mo-0.5W-0.8V steel) by
increasing the Cr content to 11 - 12 % and adding Nb to substitute for W in H40. The V content was also
reduced to 0.3 % to avoid the agglomeration of V4C3 in the temperature range of 600 - 650 °C.
2.3.7.2 Materials standards, and chemical and tensile requirements
Table 109 and Table 110 show the chemical requirements and tensile requirements of 12Cr-0.6Mo-0.3V0.4Nb-N steel (H46, JIS G4311 SUH600), respectively.
Table 111 shows the chemical compositions and heat treatment conditions of the H46 steels tested, which
were produced in the UK and Japan in the 1950s. Table 112 lists the tensile properties of three heats of
H46 steels. Fig. 187 shows the elevated temperature tensile properties of H46 steel [1].
Table 109. Chemical requirements of 12Cr-0.6Mo-0.3V-0.4Nb-N steel (H46, JIS G4311 SUH600)
Chemical composition [wt%]
Stan- DesigStd.
dard nation
No.
C
Si
Mn P
S
Cr
Ni
Mo V
Nb
N
JIS SUH600 0.15∼ ≤0.50 0.50∼ ≤0.040 10.00∼ ≤0.60 ≤0.60 0.30∼ 0.10∼ 0.20∼ 0.05∼ G4311
0.20
1.00
13.00
0.90 0.40 0.60 0.10
Table 110. Tensile requirements of 12Cr-0.6Mo-0.3V-0.4Nb-N steel (H46, JIS G4311 SUH600)
Standard Designation 0.2% yield Tensile
Elongation Reduction Hardness* Std. No.
strength
strength
[%]
of area [%] [HB]
[MPa]
[MPa]
JIS
SUH600
G4311
≥685
≥830
≥15
≥30
≤321
* ≤269 for as-annealed
Table 111. Chemical compositions of H46 steels tested [1], [2]
Chemical composition [wt%]
Steel Heat treatment
C
Si
Mn P
S
Cr
Ni
Mo
A
1150 °C AC+ 0.16 0.3 0.7 NA NA 11.6 NA 0.6
670~690 °C
AC
B
1040 °C OC+ 0.15 0.4 0.6 NA NA 11.5 NA 0.45
620 °C WC
C
1150 °C AC+ 0.13 0.15 0.79 0.008 0.028 11.47 NA 0.58
680 °C AC
D
1150 °C AC+
650 °C AC
0.16
Producer
V
0.3
Nb N
0.25 NA
Jessop
0.30 0.25 NA
Jessop
0.35 0.28 0.046 Jessop
0.18 0.66 0.012 0.014 11.90 0.12 0.17 0.30 0.26 NA
Nippon
Special
Steel
NA: Not available
Landolt-Börnstein
New Series VIII/2B
Ref. p. 160]
2.3.7 12Cr-0.6Mo-0.3V-0.4Nb-N steel
157
Table 112. Tensile properties of H46 steels at room temperature [1], [2]
Steel
Yield strength
Tensile strength
Elongation
(0.2% offset)
[MPa]
[MPa]
[%]
A
693*
919
17
C
866
1040
18
D
853
955
21
* 0.1% offset
Reduction of area
[%]
60
51
59
1100
Tensile
strength
900
800
700
600
0.1 % offset
yield strength
500
80
400
60
300 Reduction of area
40
200
100
20
Elongation
0
50
150
250 350 450
Temperature [°C]
550
650
0
750
Elongation, reduction of area [%]
Yield strength, tensile strength [MPa]
1000
Fig. 187. Elevated temperature tensile properties of H46
steel (A); [1].
2.3.7.3 Creep properties
2.3.7.3.1 Creep rupture data [1], [2]
Figs. 188 [1], 189 [2] and 190 show the creep rupture stress vs. time to rupture diagrams for three heats of
H46. H46 shows stable creep rupture strength at temperatures from 400 to 550 °C. However, strengths at
temperatures of 600 °C and above drop rapidly with longer time, except heat D shown in Fig. 190.
Heat D was normalized at a relatively low temperature followed by rapid cooling, and was also tempered
at a lower temperature, which is different from the other heats. Figs. 191 and 192 respectively show the
creep rupture elongation and reduction of area for heat D of H46 steel.
3000
2000
400℃
500℃
550℃
600℃
650℃
700℃
1000
700
Stress (MPa)
500
300
200
100
70
50
30
20
10
101
Landolt-Börnstein
New Series VIII/2B
Fig. 188. Creep rupture strength
data of H46 steel (A); [1].
102
103
104
Time to rupture (h)
105
106
158
2.3 High Cr steels
700
400℃
450℃
500
500℃
8.3
Stress (MPa)
6.9
5.0
300
550℃
8.6
200
11.8
9.8
6.1
3.6
8.8
6.8
100
8.4
70
50
Values designate
rupture elongation in %.
102
650℃
600℃
700℃
103
Time to rupture (h)
104
105
Fig. 189. Creep rupture strength data
of H46 steel (B); [2].
1000
700
550℃
600℃
650℃
700℃
500
300
Stress (MPa)
200
100
70
50
30
20
Fig. 190. Creep rupture strength data
of H46 steel (D).
10
101
102
103
Time to rupture (h)
30
20
80
Reduction of area [%]
550 °C
600 °C
650 °C
700 °C
40
Elongation [%]
105
100
50
60
40
550 °C
600 °C
650 °C
700 °C
20
10
0
10
104
10 2
10 3
10 5
10 4
Time to rupture [h]
Fig. 191. Creep rupture elongation of H46 steel (D).
0
10
10 2
10 3
10 5
10 4
Time to rupture [h]
Fig. 192. Creep rupture reduction of area of H46 steel (D).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 160]
2.3.7 12Cr-0.6Mo-0.3V-0.4Nb-N steel
159
2.3.7.3.2 Creep deformation behavior
Fig. 193 [2] shows stress vs. time to 0.1% creep strain of H46 steel in the temperature range from 500 °C
to 650 °C. Fig. 194 shows creep deformation curves, strain vs. time at 600 °C, for the stress range from
108 MPa to 294 MPa for heat D of H46 steel. Fig. 195 shows Larson-Miller plots of the average creep
strain rate obtained from time to given strain (0.1%, 0.2% and 0.3%) creep test data at temperatures from
550 °C to 700 °C, also for heat D of H46 steels. In the medium stress range of around 100 MPa, averaged
strain rates scatter more widely than in the lower stress range, meaning that the creep curves have greater
curvature at around 100 MPa than at other stresses.
300
Stress (MPa)
500℃
200
550℃
100
600℃
650℃
0
101
Fig. 193. Stress vs time to 0.1%
creep strain of H46 steel (C); [2].
103
102
104
1000
800
600
108 MPa
115 MPa
127 MPa
5
147 MPa
6
260 MPa
186 MPa
Time (h)
400
300
2
1
0
Stress [MPa]
294 MPa
3
200
219 MPa
Strain [%]
4
0.1 %
0.2 %
0.3 %
100
80
60
40
30
20
2000
6000
4000
Time [h]
8000
Fig. 194. Creep curves of H46 steel (D).
Landolt-Börnstein
New Series VIII/2B
10000
10
9
10
11
12
13
. 14
Larson-Miller-parameter T (10 - log e ) [×10 -3 ]
15
Fig. 195. Larson-Miller plots of average creep strain rate
obtained from the time to given creep strain against stress
for H46 steel (A).
160
2.3 High Cr steels
2.3.7.3.3 Creep mechanism [3]
The minimum creep rate ε&m of heat D of H46 steel is plotted against creep stress (normalized with
Young’s modulus) σ/E and reciprocal temperature in Figs. 196 (a) and (b), respectively. The stress
exponent changes at σ/E=8.5×10−4, and this value appears to be independent of temperature. The stress
exponents are 4.7 and 8.4 over the low and high stress ranges, respectively, and are temperature
independent. In Fig. 196 (a) the two points at 600 °C under σ/E=4.1×10−4 and 4.7×10−4 deviate from what
is expected from other data points. The creep curves under these conditions are completely different in
shape from those under the other conditions. Since the testing temperatures are close to the Curie
temperature (720 °C for 11.8Cr steel) the Arrhenius plot in Fig. 196 (b) is not straight as is the case with
diffusivity. An average activation energy for ε&m over the range studied is 508 kJ/mol, being substantially
greater than that for self-diffusion (360 kJ/mol).
Temperature T (℃)
10-6
700℃ ( )
650℃
600℃
700
650
10-3
10-7
σ/E=1×10-3
10
.
εm (s-1)
.
εm (h-1)
.
εm (s-1)
10-5
10-9
10-5
10-9
0.5×10-3
10-6
10
( )
( )
10-4
-8
.
10
-8
-10
550
10-3
10-4
550℃
600
10-6
10
-10
10-7
10-11
0.3
0.5
εm (h-1)
10-7
10-6
0.7 1.0
σ/E (10- 3）
(a）
2.0
3.0
10-7
10-11
1.00
1.05
1.10
1.15
T-1 (10- 3K- 1）
(b）
1.20
1.25
Fig. 196. (a) stress and (b) temperature dependence of minimum creep rate ε&m of H46 steel (D) (Creep stress s is
normalized with Young’s modulus E) [3].
2.3.7.4 References
[1] Briggs, J.Z., and Parker, T.D. (eds): The Super 12%Cr Steels, Climax Molybdenum CD. Michigan,
(1982).
[2] William Jessop, UK: Data Sheets, Jessop H46, (1956).
[3] Maruyama, K., Harada, C., and Oikawa, H.: Trans. Iron and Steel Inst. Japan, 26 (1986), 212.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 169]
2.3.8 12Cr-1Mo-1W-0.3V steel
161
2.3.8 12Cr-1Mo-1W-0.3V steel
2.3.8.1 Introduction
Steels of the 12Cr-1Mo-1W-0.3V type, specified as JIS SUH 616-B and ASTM S 42200, are used as
turbine blades at around 500 °C in modern steam power plants. Because of high operating temperatures of
such components, creep and creep rupture data are required by designer.
2.3.8.2 Materials standards and chemical requirements
Table 113. Chemical requirements for 12Cr-1Mo-1W-0.3V steel bars; JIS SUH 616-B and ASTM S
42200.
Chemical composition [wt%]
StanDesigStd.
dard
nation C
No.
Si
Mn
P
S
Ni
Cr
Mo
W
V
JIS
SUH
0.20- <0.50 0.50 - <0.040 <0.030 0.50 - 11.00 - 0.75 - 0.75 - 0.20 - G4311
616-B 0.25
1.00
1.00 13.00 1.25 1.25 0.30
ASTM S 42200 0.20 - <0.75 0.75 - <0.040 <0.030 0.50 - 11.50 - 0.75 - 0.75 - 0.15 - A565
0.25
1.25
1.00 13.50 1.25 1.25 0.30
2.3.8.3 Data sources for 12Cr-1Mo-1W-0.3V steel
Information of fact on data for 12Cr-1Mo-1W-0.3V steel bars can be obtained from [1], [2], [3].
2.3.8.4 Creep and creep rupture data for 12Cr-1Mo-1W-0.3V steel bars for turbine blades, JIS
SUH 616-B
2.3.8.4.1 Creep rupture data for 12Cr-1Mo-1W-0.3V steel bars, JIS SUH 616-B
The complete set of creep and creep rupture data, such as creep rupture time, total elongation, reduction
of area, minimum creep rate and optical micrographs of as-received and crept specimens, has been
obtained for 9 heats of 12Cr-1Mo-1W-0.3V steel, JIS SUH 616-B, in [1]. The details of steel bar
production, processing, thermal history, austenite grain size number, Rockwell hardness and volume
fraction of non-metallic inclusions before creep test, the chemical compositions, the 0.2% proof stress and
ultimate tensile strength data at high temperature are additionally available for the 9 heats in [1].
Fig. 197 shows the 0.2% proof stress and tensile strength, of the 9 heats of 12Cr-1Mo-1W-0.3V steel
bars, JIS SUH 616-B, in [1]. The tensile and creep specimens, having a geometry of 10 mm in diameter
and 50 mm in gauge length, were taken longitudinally from square bars of 50 mm by 50 mm side,
obtained by short-time tensile tests between room temperature and 650 °C.
Fig. 198 shows stress vs. time to rupture data for the 9 heats of 12Cr-1Mo-1W-0.3V steel, JIS SUH
616-B, at temperatures between 500 and 650 °C. Fig. 198 exhibits a large heat-to-heat variation in time to
rupture, which becomes more significant with increasing time and temperature. At 550 °C and 157 MPa,
the time to rupture of the strongest heat is 105 h, while that of the weakest one is only 2×104 h. It should
be noted, however, that the observed heat-to-heat variation in time to rupture is not caused by data
scattering, because each heat exhibits its distinct stress dependence of time to rupture as shown in Fig.
199. The heat-to-heat variation in time to rupture comes from many factors, as will be described later.
Landolt-Börnstein
New Series VIII/2B
162
2.3 High Cr steels
Tensile strength
1200
1000
1000
800
800
Stress (MPa)
Stress (MPa)
0.2% proof stress
1200
600
400
200
0
600
400
200
0
100
200
300
400
500
600
0
700
0
100
Test temperature (℃)
200
300
400
500
600
700
Test temperature (℃)
Fig. 197. Short-time tensile properties of 12Cr-1Mo-1W-0.3V steel bars, JIS SUH 616-B.
800
Stress [MPa]
600
500
400
300
500 °C
550 °C
600 °C
650 °C
200
100
80
60
50
40
30
n = 258
20
10
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 198. Creep rupture strength
data for 12Cr-1Mo-1W-0.3V
steel, JIS SUH 616-B. n indicates
the total number of data points.
500
12Cr-1Mo-1W -0.3V
400
Stress ( MPa )
o
550 C
300
RAA
RAB
RAC
RAD
RAE
RAF
RAG
RAH
RAJ
200
100
10
2
Fig. 199. Creep rupture strength data for each heat of
12Cr-1Mo-1W-0.3V steel, JIS SUH 616-B, at 550 °C.
10
3
10
4
10
5
Tim e to rupture ( h )
Landolt-Börnstein
New Series VIII/2B
Ref. p. 169]
2.3.8 12Cr-1Mo-1W-0.3V steel
163
2.3.8.4.2 Estimated long-term creep rupture strength for 12Cr-1Mo-1W-0.3V steel, JIS SUH 616-B
The creep rupture data shown in Fig. 198 were analyzed for each heat using the Manson-Haferd
parameter method. The solid curves in Fig. 200 are based on the Manson-Haferd parameter for the RAF
heat which showed an intermediate strength level among the 9 heats. 105 h creep rupture strength was
also estimated for the 9 heats. This is shown in Fig. 201 as a function of temperature, together with 0.2%
proof stress, ultimate tensile strength and 103 h creep rupture strength.
1000
o
Stress ( MPa )
500 C
o
550 C
10 0
o
6 00 C
o
65 0 C
Fig. 200. Estimated creep rupture
curves of 12Cr-1Mo-1W-0.3V
steel, JIS SUH 616-B.
10
2
10
10
3
10
4
10
5
10
6
T ime to ruptu re ( h )
Tensile
strength
0.2% proof
stress
{
1000 h
{
100000 h
{
Stress [MPa]
{
1000
800
600
500
400
300
200
100
80
60
50
40
30
20
450
500
550
600
Temperature [°C]
650
700
Fig. 201. Estimated 105 h creep
rupture strength for the 9 heats of
12Cr-1Mo-1W-0.3V steel, JIS
SUH 616-B.
2.3.8.4.3 Creep strain data of 12Cr-1Mo-1W-0.3V steel, JIS SUH 616-B
The stress vs. time to reach 0.5, 1, 2 and 5 % total strain, time to tertiary creep and time to rupture for the
heat RAB (Fig. 199) of 12Cr-1Mo-1W-0.3V steel are obtained from [1]. The relationship between stress
and minimum creep rate is shown for the 9 heats in Fig. 202. The relationship between time to rupture
and minimum creep rate is shown for the 9 heats in Fig. 203.
Landolt-Börnstein
New Series VIII/2B
2.3 High Cr steels
600
500
400
10 6
300
10 5
500 °C
550 °C
600 °C
650 °C
n = 124
200
Time to rupture [h]
Stress [MPa]
164
100
80
60
50
40
500 °C
550 °C
600 °C
650 °C
n = 126
30
20
10 -7
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
10 -1
10 4
10 3
10 2
1
Fig. 202. Stress vs. minimum creep rate for 12Cr1Mo-1W-0.3V steel. The corres-ponding data of time
to rupture are included in Fig. 198. n indicates the
total number of data points.
10
10 -7
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
10 -1
1
Fig. 203. Time to rupture versus minimum creep rate for
12Cr-1Mo-1W-0.3V steel. The corresponding data of time
to rupture are included in Fig. 198. n indicates the total
number of data points.
2.3.8.5 Creep behavior and microstructure of 12Cr-1Mo-1W-0.3V steel
2.3.8.5.1 Effect of high aluminum content on the log term creep properties of 12Cr-1Mo-1W-0.3V
steel
The origin of differences in long term creep properties was investigated using the selected 4 heats, RAB,
RAD, RAG and RAJ, of the 9 heats in [1] at temperatures between 500 and 650 °C [4]. Fig. 204 shows
the creep rupture data for the 4 heats. Metallurgical examination after exposure at 600 °C showed that the
microstructure and the dissolved concentration of Mo and W in the matrix were similar among the 4 heats
and had no correlation with their rupture strength. The heats RAD and RAJ, containing low aluminum
content less than 0.007 mass%, showed high strength at long times.
Thermal aging at 600 °C prior to creep test decreases the creep rupture strength at 550 °C, which is
more significant for the heats RAB and RAG with high aluminum content than for the heats RAD and
RAJ with low aluminum content. This is shown in Fig. 205. Aluminum nitride AlN forms in the heats
RAB and RAG (high aluminum content) during thermal aging at 600 °C, as shown in Fig. 206. Therefore,
it appears that a high aluminum content reduces the nitrogen concentration in solution due to the
precipitation of AlN during prolonged rupture test and leads to deterioration in rupture strength.
The rupture ductility variation among the different heats is suggested to arise from the differences in
prior austenite grain size, aluminum content and copper content. The heat RAJ with fine grains, low
aluminum content of 0.005 mass%, and low copper content of 0.03% showed a high ductility over a wide
test range.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 169]
2.3.8 12Cr-1Mo-1W-0.3V steel
RAB
RAD
RAG
RAJ
60
50
40
Stress [kgf/mm 2 ]
165
500 °C
30
20
550 °C
10
600 °C
5
650 °C
10 2
10
10 3
Time to rupture [h]
10 4
Fig. 204. Creep rupture strength
data for the 4 heats, RAB, RAD,
RAG and RAJ, of 12Cr-1Mo-1W0.3V steel.
10 5
50
12Cr-Mo-W-V
Stress [kgf/mm 2 ]
40
RAB
RAD
RAG
RAJ
30
20
Rupture strength
(550 °C, 100 hr)
10
as-received 10 2
10 3
10 4
Heating time at 600 °C [h]
10 5
Fig. 205. Creep rupture strength of 4 heats of 12Cr-1Mo1W-0.3V steel at 550 °C and 102 h, as a function of heating
time at 600 °C prior to creep test at 550 °C.
2.3.8.5.2 Creep deformation behavior and microstructure evolution during creep of 12Cr-1Mo1W-0.3V steel, JIS SUH 616-B
The creep deformation behavior and microstructural evolution during creep were investigated for the 12Cr1Mo-1W-0.3V steel, JIS SUH 616-B, at 600 °C, comparing with those for a simple 12Cr steel, SUS 403-B
without Mo, W and V [5,6]. Main focus was placed on the effect of vanadium addition to 12Cr steel on the
degradation during creep deformation. Fig. 207 shows the creep rate vs. time curve of the RAJ heat of 12Cr1Mo-1W-0.3V steel at 600 °C and 137 MPa, comparing with that of the 12Cr steel at 600 °C and 47 MPa.
The creep specimens had a size of 8 mm in diameter and of 50 mm in gage length. The time to rupture was
almost the same for both steels; 5124 and 6084 h for the 12Cr-1Mo-1W-0.3V and 12Cr steel, respectively.
Both creep rate versus time curves consist of two stages: the transient and tertiary creep stages. No steady
state creep stage was observed. The minimum creep rate of the 12Cr-1Mo-1W-0.3V steel is about half of
that of the 12Cr steel, although the time to rupture of the former is slightly shorter than that of the latter. It
should be also noted that the acceleration of creep rate, defined as d ε& /dt, in the tertiary creep stage, was
larger in the 12Cr-1Mo-1W-0.3V steel by a factor of two than in the 12Cr steel.
Landolt-Börnstein
New Series VIII/2B
166
2.3 High Cr steels
0.03
0.015
12 Cr-Mo-W-V
12 Cr-Mo-W-V
RAB
RAD
RAG
RAJ
0.010
0.02
Al as AIN [%]
N as nitride [%]
RAB
RAD
RAG
RAJ
0.01
0.005
0
as-received 10 2
0
as-received 10 2
10 4
10 3
Heating time at 600 °C [h]
10 3
Heating time at 600 °C [h]
10 4
Fig. 206. Increase in nitrogen and aluminum as aluminum nitride AlN during thermal aging at 600 °C, by alcohol iodine method.
10
-6
0
1000
Time [h]
2000 3000 4000 5000
6000
10 -3
12Cr
873K, 47MPa
Prior creep testing specimens
3
10 - 7
3
10 -4
t r = 21.9 10 6s
3
10 - 8
3
3
-10
10 -6
a
10
10 - 6
12CrMoWV
873K, 137MPa
Prior creep testing specimens
3
10 - 7
10 -3
3
10 -4
3
t r = 18.4 10 6s
10 - 8
3
10 - 9
10
3
Creep rate [h -1]
Creep rate [s -1]
10 - 9
3
10 - 5
3
-10
3
10 - 5
3
10 -6
b
0
10
Time [10 6 s]
20
Fig. 207. Creep rate vs. time curves of the 12Cr and the
12Cr-1Mo-1W-0.3V steels at 600 °C - 47 MPa and 600
°C - 137 MPa, respectively.
After interruption of the creep tests for the 12Cr-1Mo-1W-0.3V steel at 600 °C and 137 MPa and for the
12Cr steel at 600 °C and 47 MPa (Fig. 207) at several different times, the creep specimens were machined
to a smaller size of 6 mm in gage diameter and of 30 mm in gage length to eliminate any mechanical
damages developed near surface during the previous creep tests. Then the creep tests were re-started at
higher stress conditions for the 12Cr-1Mo-1W-0.3V steel at 600 °C and 225 MPa and for the 12Cr steel at
600 °C and 84 MPa. The results are shown in Figure 208. The previous creep test (Fig. 207) is denoted as
Landolt-Börnstein
New Series VIII/2B
Ref. p. 169]
2.3.8 12Cr-1Mo-1W-0.3V steel
167
the first stage creep test, while the later one (Fig. 208) as the enhanced stress creep test. The minimum
creep rate in the enhanced stress creep test becomes larger with increasing creep-interrupted time in the
first stage creep testing both the 12Cr-1Mo-1W-0.3V and 12Cr steels. In the enhanced stress creep test,
the minimum creep rate of the creep-interrupted specimen at 4000 h (14.4 Ms) was only five times larger
in the 12Cr steel than that of the as-tempered specimen without being subjected to the first stage creep
test, while two orders of magnitude larger in the 12Cr-1Mo-1W-0.3V steel.
1
3
10
3
3
10 3
3
10 4
3
14.4 106s
3
10
Time [h]
3 10 2
12Cr
873K
84MPa
10.8 106s
7.2 106s
-5
10
1.08 106s
6
3.6 10 s
3
6
1.8 10 s
3
Creep rate [s -1]
10
10
10 - 3
0.36 106s
-7
3
-8
3
As tempered
10 -4
a
10 - 4
3
10 - 5
3
12CrMoWV
873K
225MPa
6
14.4 10 s
3
10
10
3
10 - 2
3
-6
10 -1
10.8 106s
-6
3
10 - 2
3
3.6 106s
3
-7
1.08 106s
10 - 3
As tempered
3
10 -4
3
10 - 8
b
3×10 8
-3
10
10 -1
Creep rate [h -1]
10
-4
3
3
10 - 2
3
10 -1 3 1
Time [10 6 s ]
3
10
3
10 2
Fig. 208. Creep rate vs. time curves of the creepinterrupted 12Cr and 12Cr-1Mo-1W-0.3V steel
specimens at 84 and 225 MPa, respectively, at
600 °C.
The Vickers hardness and the half-value width of an X-ray diffraction peak at (211) were measured to
evaluate the overall softening during the first stage creep test shown in Fig. 208. The results are shown in
Fig. 209 as a function of t/tr. The overall softening is mainly caused by the coarsening of carbide
precipitates and by the annihilation of pre-existing dislocations. The large decrease in hardness and halfvalue width occurs in the transient creep stage rather than in the tertiary creep stage. The hardness and
half-value width at the time of rupture in the 12Cr-1Mo-1W-0.3V steel is still larger than those of the
12Cr steel on an as-tempered condition.
The 12Cr-1Mo-1W-0.3V steel exhibited tempered martensitic microstructure in as-tempered
condition. The TEM observations showed that the density of precipitated carbides and dislocations
remained high within laths even after 4000 h (14.4 Ms) of the first stage creep test. At the minimum creep
rate stage of the first stage creep test, a recovered zone with a low density of precipitated carbides and
dislocations started to form along prior austenite grain boundaries. The average width of this recovered
zone increased with increasing testing time. The local softening due to the development of the recovered
zone is mainly responsible for the acceleration of creep rate. The loss of creep resistance proceeds due to
local softening along prior austenite grain boundaries, accompanied by the decrease in density of V4C3
carbides and dislocations in the recovered zone.
Landolt-Börnstein
New Series VIII/2B
168
2.3 High Cr steels
Hardness [HV :98N]
350
873 K
300
12CrMoWV, 137MPa
250
12Cr, 47MPa
200
150
Half - value width [° ]
2.2
12CrMoWV, 137MPa
2.0
1.8
1.6
12Cr, 47MPa
0
0.2
0.4
0.6
Normalized time t / t r
Fig. 209. Vickers hardness and the half-value width of
an X-ray diffraction peak at (211) as a function of
normalized time t/tr during the first stage creep test.
0.8
1.0
2.3.8.5.3 Heat-to-heat variation in long term creep strength of 12Cr-1Mo-1W-0.3V steel, JIS SUH
616-B
The heat-to-heat variation in long term creep strength was investigated for the 9 heats of 12Cr-1Mo-1W0.3V steel by analyzing the creep rupture data in [7]. The concentrations of the major alloying elements
C, Si, Mn, Cr, Mo and V are not very different among the 9 heats [5]. Therefore, these parameters are
excluded as main explanation of the observed heat-to-heat variation in time to rupture. Then our interest
is concentrated on the effect of minor alloying elements. It has been well known for ferritic and austenitic
heat resistant steels that nitrogen causes a beneficial effect on the long term creep rupture strength but that
Al causes a deteriorative effect. Fig. 210(a) shows that a monotonous decrease in time to rupture with an
increase in Al content is not observed, suggesting that the time to rupture of the 12Cr-1Mo-1W-0.3V steel
is not directly related with the Al content and other factors should be taken into account. The time to
rupture is again plotted in Figure 210(b) as a function of effective nitrogen concentration, N − Al − Ti in
at %. The effective nitrogen concentration is defined as the nitrogen concentration available for the
formation of fine vanadium nitrides. Because Al is a stronger nitride forming element than Cr and V, the
effective nitrogen concentration is reduced by the formation of AlN. Titanium is also known as a strong
nitride forming element. The time to rupture increases with increasing effective nitrogen concentration,
which becomes more significant with decreasing stress and increasing time.
The creep rate versus time curves in Fig. 211 show that the creep rate is not much different in the
transient region among the different heats but that the onset of acceleration creep occurs at shorter times
in the high-Al heat than in the low-Al heat. This results in higher minimum creep rate and shorter rupture
time in the high-Al heat than in the low-Al heat.
The microstructure evolution was investigated by several researchers. Yokoi and co-workers [4]
investigated the formation of AlN during aging at 600 °C for some heats with different Al content. Their
results show that AlN is already present after heat treatment and before creep test and that additional
precipitation of AlN occurs during aging at 600 °C. On the other hand, Watanabe and co-workers
reported that preferential recovery occurs in the vicinity of prior austenite grain boundaries during creep
at 600 °C [6].
Landolt-Börnstein
New Series VIII/2B
10
5
10
4
10
3
2.3.8 12Cr-1Mo-1W-0.3V steel
o
550 C
176
235
294
333
MPa
MPa
MPa
MPa
Time to rupture ( h )
Tim e to rupture ( h )
Ref. p. 169]
(a)
0
0.01
0.02
0.03
0.04
10
5
10
4
10
3
o
550 C
176 MPa
235
294
333
(b)
0
0.05
169
0.04
0.08
Fig. 210. Time to rupture of 12Cr1Mo-1W-0.3V steel at 550 °C, as a
function of Al content in mass %
and
of
effective
nitrogen
concentration, N(nitrogen)-Al-Ti in
at %. The corresponding creep
rupture data are shown in Fig. 199.
0.12
N - Al - Ti ( at % )
Al content ( m ass % )
o
Creep rate ( 1 / h )
650 C, 69 M P a
10
-3
10
-4
10
-5
10
R AG , 0.0 37% Al
R AA, 0.030% Al
R AJ, 0.005% Al
-2
10
-1
10
0
10
1
10
2
10
3
10
4
Fig. 211. Creep rate vs. time curves of the 3 heats
containing different aluminum content at 650 °C and
69 MPa.
T ime ( h )
Based on above results, it is concluded that Al promotes the preferential recovery in the vicinity of prior
austenite grain boundaries, which may be promoted by the formation of AlN, because this consumes
available fine vanadium nitrides. The preferential recovery in the vicinity of prior austenite grain
boundaries accelerates the onset of acceleration creep. This reduces the time to rupture.
2.3.8.6 References
[1]
[2]
[3]
[4]
National Research Institute for Metals: NRIM Creep Data Sheet, No.10B, (1998).
ASTM Data Series Publication DS50, (1973).
The British Steelmakers Creep Committee (BSCC) High Temperature Data (1972).
Yokoi, S., Shin-ya, N. and Kori, M.: Journal of the Society of Materials Science, Japan, 26 (1977),
241-247.
[5] Matsuo, T., Kikuchi, M., Watanabe, T. and Monma, Y.: Proceedings of the Fifth International
Conference on Creep of Materials, Lake Buena Vista, Florida, USA, 18-21 May (1992), 271-279.
[6] Watanabe, T. and Monma, Y., Matsuo, T. and Kikuchi, M.: Report of the 123rd Committee on HeatResisting Metal and Alloys, Japan Society for the Promotion of Science, Vol.32 (1991) 137-148.
[7] Abe, F.: Proceedings of NRIM-MPA Workshop on Creep and Fatigue Performance of High Cr
Steels for Elevated Temperature Plants, Tsukuba, Japan (2001), 1-10.
Landolt-Börnstein
New Series VIII/2B
170
2.3 High Cr steels
2.3.9 12Cr-1Mo-V steel
2.3.9.1 Introduction
12Cr-1Mo-V steel (X20CrMoV12-1; X20CrMoV(W)12-1, X22CrMoV12-1) has been widely used for
tubing, headers and piping in Europe. The steel was developed in the 1960s together with a modification
for bars and forgings with an increased C-content (0.20 - 0.26 %) - X22CrMoV12-1 as well as a
modification with an additive of W - X22CrMoWV12-1, X20CrMoWV12-1. The maximum long term
service temperature for tubes and pipes is generally limited to 565 °C.
Due to the high Cr- and Mo-content X20CrMoV12-1 steel shows a distinctive martensitic microstructure, which permits its use as thick pipe material. If the standard cooling down in air to room
temperature is applied a residual austenite content of approximately 2 - 5 % could be stated, influenced by
the specific chemical composition of the heat. A change in the standardized heat treatment conditions
leads to modifications in the microstructure and thus to a decrease of the long term creep strength.
X20CrMoV12-1 steel demonstrates high creep rupture ductility also in the long term service range. As
a consequence the sensitivity to the formation of creep cavities is low compared with that of low alloyed
steels. This should be considered if the damage state and life expenditure of the component is evaluated
by means of the replica technique and the consequent appearance of cavities. The steel is sensitive to
intergranular stress corrosion if hardening (austenizing and cooling down) is not followed by tempering.
The steel may be welded if the relevant measures required for the material are observed. Specific care
is required in welding: correct pre and post weld heat treatment should be done in order to avoid cracking.
Pre heating up to 450 °C is necessary in dependence of thickness. After welding an intermediate cooling
down below 130 °C should be performed in order to optimize the martensite formation in the deposit
material (if similar to X20CrMoV12-1) and heat affected zone. Post heat treatment should be done at a
temperature ranging from 750 °C to 770 °C.
2.3.9.2 Materials standards, chemical and tensile requirements
2.3.9.2.1 X20CrMoV12-1 seamless tubes for pressure purposes
Table 114. Heat treatment of X20CrMoV12-1 seamless tubes for pressure purposes; DIN 17175:1979
DIN
17175
Austenizing temperature / cooling medium
1020 - 1070 °C / Air
Tempering temperature / cooling medium
730 - 780 °C/ Air
Table 115. Chemical requirements of X20CrMoV12-1; DIN 17175:1979
Chemical composition [wt%]
Standard Designation/
steel number
C
Si
Mn
P
S
Cr
Mo
DIN
X20CrMoV12-1 0.17
≤0.50 ≤1.00 ≤0.030 ≤0.030 10.00 0.80
17175
1.4922
0.23
12.50 1.20
Ni
0.30
0.80
V
0.25
0.35
Table 116. Mechanical properties of X20CrMoV12-1 at room temperature; DIN 17175:1979
Tensile requirements
Standard Designation/
steel number
Thickness Yield
Tensile
Elongation
Impact energy
stress
strength
after fracture Charpy-V-testpiece
L
T
L
T
[mm]
[MPa]
[MPa]
[%]
[%]
[J]
[J]
DIN
X20CrMoV12-1
490
Not listed ≥17 ≥14
≥16, ≤60
≥34
17175
1.4922
T transverse direction, L longitudinal direction
Landolt-Börnstein
New Series VIII/2B
Ref. p. 179]
2.3.9 12Cr-1Mo-V steel
171
2.3.9.2.2 X20CrMoNiV11-1-1 seamless tubes for pressure purposes
Table 117. Heat treatment of X20CrMoNiV11-1-1-1; ISO 9329-2:1997
ISO
9329
Austenizing temperature / cooling medium
1020 - 1080 °C / Air
Tempering temperature / cooling medium
730 - 780 °C/ Air
Table 118. Chemical requirements of X20CrMoNiV11-1-1; ISO 9329-2:1997
Chemical composition [wt%]
Standard Designation/ steel number
C
Si
Mn
P
S
Cr
Mo Ni
V
ISO 9329 X20CrMoNiV11-1-1
0.17 0.15 ≤1.00 ≤0.030 ≤0.030 10.00 0.80 0.30 0.25
0.23 0.50
12.50 1.20 0.80 0.35
Table 119. Mechanical properties of X20CrMoNiV11-1-1 at room temperature; ISO 9329-2:1997
Tensile requirements
Standard Designation/steel
number
Thickness Yield Tensile
Elongation
Impact energy
stress strength after fracture Charpy-Vtestpiece
L
T
T
L
[mm]
[MPa] [MPa]
[%] [%]
[J]
[J]
ISO 9329 X20CrMoNiV11-1-1
490
Not listed ≥17 ≥14
≥16, ≤60
≥27
≥35
T transverse direction, L longitudinal direction
2.3.9.2.3 X21CrMoV12-1 for large forgings for components in turbine and generator equipment;
SEW 555
Table 120. Heat treatment of X21CrMoV12-1; SEW 555:2001
Austenizing temperature / cooling medium The heat treatment used to achieve the
SEW
specified properties is left up to the
555
manufacturer, if this is not agreed in the order.
Tempering temperature / cooling medium
The customer shall be advised.
Table 121. Chemical requirements of X21CrMoV12-1; SEW 555:2001
Chemical composition [wt%]
Standard Designation/
steel number
C
Si
Mn
P
S
Cr
Mo
0.30
11.00
0.80
SEW
X21CrMoV12-1 0.20
≤0.020 ≤0.007
≤0.20
0.80
12.50 1.20
555
1.4926
0.26
Ni
0.30
0.80
Table 122. Tensile requirements of X21CrMoV12-1; SEW 555:2001
Tensile requirements
Standard Designation
Range of
0.2 %
Tensile Elongation Reduction
significant proof
strength after
of area
dimensions strength
fracture
after
fracture
T
Q
T
Q
[mm]
[MPa]
[MPa]
[%] [%] [%] [%]
SEW
X21CrMoV12-1 1500
600
750 ≥14 ≥11 ≥40 ≥40
555
1.4926
900
T and Q correspond to the relevant direction to the fibre flow
Landolt-Börnstein
New Series VIII/2B
V
0.25
0.35
Impact
energy
Charpy-Vtestpiece
T
Q
[J]
[J]
≥20 ≥12
172
2.3 High Cr steels
2.3.9.2.4 X20CrMoV11-1 for forgings and rolled or forged bars for pressure purposes; ISO 93272:1997; EN 10222-2:2000
Table 123. Heat treatment of X20CrMoV11-1; ISO 9327-2:1997; EN 10222-2:1999
ISO
9327
Austenizing temperature / cooling medium
1020 - 1070 °C / cooling air, oil, water
Tempering temperature
730 - 780 °C
Table 124. Chemical requirements of X20CrMoV11-1; ISO 9327-2:1997; EN 10222-2:1999
Chemical composition [wt%]
Standard Designation/
steel number
C
Si
Mn
P
S
Cr
Mo
Ni
ISO
0.17
0.30
≤0.40 0.30
≤0.035 ≤0.030 10.00 0.80
X20CrMoV11-1 0.23
9327
1.00
12.50 1.20
1.00
1.4922
0.17
0.30
EN
≤0.40 0.30
≤0.025 ≤0.015 10.00 0.80
0.23
1.00
12.50 1.20
0.80
10222
V
0.20
0.35
0.20
0.35
Table 125. Tensile requirements of X20CrMoV11-1; ISO 9327-2:1997; EN 10222-2:2000
Tensile requirements
Standard Designation
Range of
0.2 %
Tensile
Elongation
Impact energy
significant
proof
strength after fracture Charpy-V-testpiece
dimensions
strength
x
y
x-y
y-x
[mm]
[MPa]
[MPa]
[%] [%]
[J]
[J]
ISO
≤100
≥39
≥27
X20CrMoV11-1 >100, ≤200
9327
500
700 - 850 ≥16 ≥16
≥31
≥27
1.4922
≥14
≥27
≥24
>250, ≤300
EN
≤10
≥39
500
700 -850 ≥16 ≥14
10222
≥27
>100, ≤250
≥31
>250, ≤330
≥27
x, y acc. ISO 1546; x, x-y longitudinal; y, y-x transverse
Fig. 212. Microstructure
of X20CrMoV12-1 (as
received state).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 179]
2.3.9 12Cr-1Mo-V steel
173
2.3.9.3 Mechanical properties of X20CrMoV12-1
2.3.9.3.1 X20CrMoV12-1, X22CrMoV12-1, X20CrMoWV12-1, X22CrMoWV12-1 for seamless
tubes and bars
1400
Group 1, without W
Group 1, with W
Group 2, without W
Group 2, with W
Group 3, without W
Group 3, with W
1200
Stress σ (MPa)
1000
800
600
400
200
Fig. 213. High temperature tensile
strength of X20CrMo(W)V12-1.
0
0
100
200
300
400
500
600
700
800
900
100
Test temperature T (°C)
The data of Fig. 212 is published in [1] from which data on proof stress, elongation and reduction of area
can also be obtained. The plot includes data with different heat treatment, consequently the data has to be
divided into the following groups with different tensile strength at room temperature:
Group
1
2
3
Yield strength [MPa]
≥490
≥589
≥785
Tensile strength [MPa]
638 - 883
785 - 932
≥932
It has to be considered that the heat treatment applied to group 2 and 3 heats however does not correspond
to the specified conditions in the codes and standards cited above.
2.3.9.4 Creep properties of X20CrMoV12-1
2.3.9.4.1 Creep strength of X20CrMoV12-1, X20CrMoWV12-1 bars and tubes
The creep rupture strength of X20CrMoV12-1, X20CrMoWV12-1 bars and tubes obtained from
published literature is shown in the following.
The results of 13 heats of X20CrMo(W)V12-1 are shown in Fig. 214. The data of this figure is
published in [1] from which data on stress, elongation and reduction of area of the individual creep tests
can also be obtained. In the figure only test results of heats with a heat treatment which corresponds to the
specified conditions in the codes and standards cited above are plotted. A significant loss in creep strength
at 600 °C especially in the long term range could be observed. This gives the reason for the limitation of
maximum service temperature of this type of steel.
Landolt-Börnstein
New Series VIII/2B
174
2.3 High Cr steels
Stress σ(MPa)
1000
500 °C with W
550 °C with W
550 °C without W
600 °C with W
100
1
10
Fig. 214. Creep rupture
strength of X20CrMo(W)
V12-1; [1].
100
1000
10000
100000
Time to rupture t (h)
1000
Stress σ (MPa)
broken
unbroken
500°C DIN 17175
Fig. 215. Creep rupture
strength data of steel grade
X20CrMoV12-1 at 500 °C (10
heats, bars and tubes) obtained
by the German Creep Committee [2], and average creep
rupture strength values indicated in DIN17175.
100
10
100
1000
10000
100000
1000000
Test duration t (h)
Stress σ (MPa)
1000
100
broken
unbroken
550°C DIN 17175
10
10
100
1000
10000
100000
1000000
Fig. 216. Creep rupture
strength data of steel grade
X20CrMoV12-1 at 550 °C (27
heats, bars and tubes) obtained
by the German Creep Committee [2], and average creep
rupture strength values indicated in DIN17175.
Test duration t (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 179]
2.3.9 12Cr-1Mo-V steel
175
1000
Stress σ (MPa)
broken
unbroken
600°C DIN 17175
100
Fig. 217.
Creep rupture
strength data of steel grade
X20CrMoV12-1 at 600 °C (23
heats, bars and tubes) obtained
by the German Creep Committee [2], and average creep
rupture strength values indicated in DIN17175.
10
10
100
1000
10000
100000
1000000
Test duration t (h)
1000
Stress σ (MPa)
broken
unbroken
100
Fig. 218.
Creep rupture
strength data of steel grade
X20CrMoV12-1 at 600 °C (5
heats, bars and tubes) obtained
by the German Creep Committee; [2].
10
10
100
1000
10000
100000
1000000
Test duration t (h)
2.3.9.4.2 Fact creep data of X20CrMoV12-1 tubes, bars and forgings
The information of fact creep data for X20CrMoV12-1 tubes, bars and forgings can be obtained from [1].
More actual data of heats in accordance to DIN 17175 is available in [2].
2.3.9.4.3 Microstructural change
The typical microstructure X20CrMoV12-1 is a needle shaped martensitic structure. The martensite
needles are decorated with precipitations. Using the lightmicroscope no significant change in the
microstructure due to service exposure is visible. The typical martensitic needle shaped microstructure is
very stable. Laves phase could only be stated after long term service exposure or at non-acceptable high
temperatures. In the long term range the precipitates at the needle boundaries will grow. An incipient
dissolution of the needle shaped microstructure could be observed, Fig. 219. However this is not an
indication of exhaustion or damage of the material.
Landolt-Börnstein
New Series VIII/2B
176
2.3 High Cr steels
Fig. 219. As received state (a) crept state (b)
(550 °C/118 MPa/51447 h), X20CrMoV12-1.
Incipient dissolution of the needle shaped
microstructure due to the growth of carbides.
The formation of creep cavities could be observed at the prior austenite grain boundaries. Also nonmetallic inclusions as MnS are preferred locations for the nucleation of creep cavities. Contents of S and
P considerably below the allowed upper limit given in the standard are advantageous.
Due to the coalescence with non-metallic inclusions a careful metallographic preparation is necessary
in order to identify creep cavities. Especially if the replica technique is used to detect creep damage at
components preferably mechanical polishing should be used. If electropolishing and etching is used there
is a risk that inclusions will be removed from the matrix and the subsequent hole will be identified as
creep cavity. Detailed information about the creep damage development in X20CrMoV12-1 could be
obtained in [3].
2.3.9.4.4 Effect of W, Mo and V on creep rupture strength
Investigations, done by [1] realized that W does not cause an improvement of the creep strength, see
Fig. 220. Optimized creep strength will be obtained if 0.5 - 1 % Mo is balanced by 0.15 - 0.35 % V. Also
the content of C should not fall below the lower bound given in the standard, Fig. 221 [4].
2.3.9.4.5 Effect of heat treatment on long term creep rupture strength
Investigations and failures at components [5] revealed that especially low quenching temperatures will
cause a decrease of the creep strength.
The mechanical properties like tensile strength are no guarantee for optimized creep rupture strength.
Any deviation from the typical needle shaped martensitic microstructure will cause a drop in the creep
strength. A ferritic matrix with carbides is not acceptable. The presence of Laves phase in the as received
state is an indication of wrong heat treatment.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 179]
2.3.9 12Cr-1Mo-V steel
177
1000
Stress σ (MPa)
550 °C with W
550 °C without W
100
10
100
1000
10000
100000
Fig. 220. Effect of W on
creep rupture strength of
X20CrMoV12-1 with heat
treatment according to the
requirements of the cited
standards. The content of W
ranges from 0.013 to 0.59 %.
No increase of creep strength
could be observed; [1].
Creep strength test heat / average creep strength V
Time to rupture t (h)
1.4
1.2
M+20%
1.0
M
0.8
M-20%
500 °C
550 °C
575 °C
580 °C
600 °C
0.6
Fig. 221. Optimal creep strength will be obtained at C
contents higher than 0.17%. M: average value.
Forgings, bars
Pipes, plates
0.11 0.13 0.15 0.17 0.19 0.21 0.23 0.25 0.27
C content [%]
550 °C
Stress [N /mm 2 ]
10 2
10 2
Fig. 222. Austenitizing temperatures below the recommended ones in the relevant codes will lead to a decrease in the creep rupture strength of X20CrMoV12-1;
[6].
600 °C
mean value
lower scatter bound
10
650 °C
10
10 2
Landolt-Börnstein
New Series VIII/2B
10 3
Rupture time [h]
10 4
10 5
L = air
Rp/Rp0.2
[N/mm2]
Rm
[N/mm2]
■ 1/2 h, 920 °C/L; 2 h, 780 °C/L
* 1/2 h, 925 °C/L; 2 h, 780 °C/L
459
558
724
755
○ 1/2 h, 975 °C/L; 2 h, 750 °C/L
605
805
● 1/2 h, 1000 °C/L; 2 h, 750 °C/L 603
810
◆ 1/2 h, 1050 °C/L; 2 h, 750 °C/L 606
814
178
2.3 High Cr steels
2.3.9.4.6 Creep crack growth
Creep crack growth may occur at defects in components operated in the high temperature regime. The
assessment of crack growth at defects is based on fracture mechanics and the relevant data. The crack
growth rate was measured using fracture mechanic specimens (CT, Cs -side grooved CT, D) with
different ratios of crack start length a0 to specimen width W. The stress situation at crack tip was
calculated using the creep fracture mechanics parameter C*.
10 -1
X22 CrMoV 121, AMB, J = 550 °C
10 -2
10 -3 a = 0.14×(C *)
[mm/h]
sn0 [MPa]
CT20 383
CT20 250
CT20 200
CT100 250
CT40 250
CT40 233
D9 304
D9 255
D9 210
10 -4
10
10 -5
10 -6
10 -6 10 -5
10 -4 10 -3 10 -2 10 -1
C * [N/mmh]
10 2
10
1
Fig. 223. Crack growth rate in dependence of the
parameter C* (based on load line displacement rate) for
different specimen types and sizes of X22CrMoV12-1
(forgings). CT20, CT100, CT40 – CT specimens with
thickness 20, 100 and 40 mm, D9 – double edge notched
specimens, thickness 9 mm, σn0 nominal stress at begin
of the test; [7, 8, 9].
10 -1
10 -8
X22 CrMoV 12 1, 220sa/AMB, J = 550 °C
10 -2
sn0 = 312 MPa
(K l0 = 830 N/mm 3/2 )
246
(654)
a [mm/h]
a [mm/h]
10
0.86
10 -9
10 -3
10
10 -10
209
(556)
-4
200
(345)
10 -5
145 160
(381) (276)
240
(422)
a0 / W
Cs20 D9
0.55 0.4
10 -11
10 -12
10 -6
10 -5
10 -4
10 -3
10 -2
C * [N/mmh]
10 -1
Fig. 224. Crack growth rate in
dependence of the parameter C*
(based on load line displacement
rate) for different specimen types of
X20CrMoV12-1 (seamless tube).
Cs20 – side grooved CT specimens
(thickness 20 mm), D9 – double
edge notched specimens (thickness
9 mm); [7, 8, 9].
1
Landolt-Börnstein
New Series VIII/2B
Ref. p. 179]
2.3.9 12Cr-1Mo-V steel
179
2.3.9.4.7 Estimated long term creep rupture strength
Table 126. Average values of creep rupture strength of X20CrMoV12-1, X20CrMoNiV11-1-1 for
seamless tubes, X20CrMoV12-1 for forgings and rolled or forged bars according to the relevant standards
DIN 17175, ISO 9327 and ISO 9329. Long term data is partly based on extended time and stress
extrapolations.
Temperature
°C
500
510
520
530
540
550
560
570
580
590
600
Temperature
°C
500
510
520
530
540
550
560
570
580
590
600
Time to rupture (h)
105 2•105 104
105 2•105 2.5•105
Average creep rupture strength (MPa)
DIN 17175
ISO 9329
294 235 215 290 237 221
215
274 211 191 264 212 196
190
253 186 167 240 189 172
167
232 167 147 217 167 151
145
213 147 128 196 146 130
125
192 128 111 176 127 112
107
173 112
96
157 109
95
90
154 96
81
139
93
80
76
136 82
68
123
80
68
65
119 70
58
107
68
58
56
101 59
48
93
59
50
48
104
Time to rupture (h)
105 2•105 104
105 2•105 2.5•105
Average creep rupture strength (MPa)
EN 10222
ISO 9327
292 236 218 294 248 234
229
269 212 194 274 225 213
208
247 188 170 253 202 190
185
225 167 149 232 180 167
161
205 147 129 213 159 143
137
184 128 112 192 139 122
117
165 111
96
173 121 104
100
147 95
81
154 104
89
84
130 81
68
136
88
76
72
113 69
58
119
75
64
60
97
59
49
101
63
53
50
104
2.3.9.5 References
[1] Ergebnisse deutscher Zeitstandversuche langer Dauer. Verlag Stahleisen mbH (1969).
[2] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade G17CrMoV5-10, compilation of test results; Forschungsvereinigung
Warmfeste Stähle, c.o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D) (2001).
[3] Guideline for the assessment of microstructure and damage development of creep exposed materials
for pipe and boiler components.
VGB-TW 507, VGB Technische Vereinigung der Großkraftwerksbetreiber e. V., Essen.
[4] Jesper, H., and Kautz, H. R.: Eigenschaften, Verarbeitung und Bewährung des Stahles
X20CrMo(W)V 12 1 im Kraftwerk. VGB Konferenz, Werkstoffe und Schweißtechnik im Kraftwerk
(1985), VGB Technische Vereinigung der Großkraftwerksbetreiber e. V., Essen.
Landolt-Börnstein
New Series VIII/2B
180
2.3 High Cr steels
[5] Wachter, O., Müsch, H., and Bendick, W.: Zeitstandschädigung an Komponenten in
Frischdampfleitungen aus dem Werkstoff X 20 CrMoV 12 1. VGB Konferenz „Werkstoffe und
Schweißtechnik im Kraftwerk (1991), VGB Technische Vereinigung der Grosskraftwerksbetreiber
e. V., Essen.
[6] Fabritius, H., and Weber, H.: Zur Betriebssicherheit von Anlagen nach langer
Betriebsbeanspruchung im Zeitstandbereich. Sonderheft VGB-Konferenz „Werkstoffe und
Schweißtechnik im Kraftwerk (1976), S. 179-217. VGB Technische Vereinigung der
Grosskraftwerksbetreiber e. V., Essen.
[7] Granacher, J., Tscheuschner, R., Mao, T. S., Maile, K., and Bareiss, J.: Anwendung numerischer
Verfahren zur Auswertung von Kriechrissversuchen am Stahl X 22CrMoV 12.
Schädigungsmechanismen und Bruch, Berichtsband der 28. Tagung in Bremen. Deutscher Verband
für Materialforschung und –prüfung, Berlin.
[8] Granacher, J., Tscheuschner, R., Maile, K., and Eckert, W.: Rissverhalten warmfester Stähle im
Kriech- und Kriechermüdungsbereich. Abschlussbericht zum AiF-Vorhaben Nr.7251 (1992), MPA
Stuttgart und IfW Darmstadt.
[9] Granacher, J., Kostenko, Y., Maile, K., and Schellenberg, G.: Kriechrissverhalten ausgewählter
Kraftwerksstähle in erweitertem, praxisnahem Parameterbereich. Abschlussbericht zum Vorhaben
Nr.186 (1998), Forschungskuratorium Maschinenbau e.V.; Frankfurt.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 188]
2.3.10 12Cr-1Mo-1W-V-Nb steel
181
2.3.10 12Cr-1Mo-1W-V-Nb steel
2.3.10.1 Introduction
12Cr-1Mo-1W-V-Nb ferritic steel (HCM12) is used as water wall, superheater and reheater tubes in
fossile fired boilers and black liquor recovery boilers. The steel has been developed for improving creep
rupture strength of 410 type ferritic steel at elevated temperatures mainly by substituting a part of Mo by
W. The microstructure of the steel consists of 15 - 30 % delta-ferrite and tempered martensitic matrix
strengthened by M23C6 carbide mainly along grain boundaries and fine dispersed (V,Nb)(C,N)
carbonitride in matrix. (V,Nb)(C,N) is fine and stable even after long term creep exposure at high
temperatures.
2.3.10.2 Material standards, chemical and tensile requirements
Tables 127 and 128 give the chemical requirements and the corresponding tensile requirements of 12Cr1Mo-1W-V-Nb steel tubes and pipes which are designated by the standards: Japanese METI KASUS410J2TB, VD-TÜV 510 HCM12.
Table 127. Chemical requirements of 12Cr-1Mo-1W-V-Nb steel tubes; Japanese METI KASUS410J2TB, VD-TÜV 510 HCM12.
Designation
Grade
Japanese
METI
VD-TÜV
510
KA-SUS
410J2TB
HCM12
C
0.14
0.14
Si
0.50
0.50
Mn
0.30
0.70
0.30
0.70
Chemical composition [wt%]
P
S
Cr
Mo
11.00
0.80
0.030 0.030 13.00
1.20
11.00
0.80
0.030 0.030 13.00
1.20
W
0.80
1.20
0.80
1.20
V
0.20
0.30
0.20
0.30
Nb
0.20
0.20
Table 128. Tensile requirements of 12Cr-1Mo-1W-V-Nb steel tubes; Japanese METI KA-SUS410J2TB,
VD-TÜV 510 HCM12.
Designation
Grade
Min. TS1)
Min. YS2) Min. elongation
Japanese METI KA-SUS 410J2TB 590 MPa
390 MPa
20 %
VD-TÜV 510
HCM12
588 MPa
392 MPa
20 %
1) TS; tensile strength, 2) YS; yield strength as 0.2% proof stress
2.3.10.3 Tensile properties of 12Cr-1Mo-1W-V-Nb steel tubes
Fig. 225 shows tensile strength and yield stress data of 12Cr-1Mo-1W-V-Nb steel tubes [1]. They are
higher than those of T9 steel at all temperatures up to 650 °C. The corresponding tensile elongation and
reduction of area data of 12Cr-1Mo-1W-V-Nb steel tubes are available in [1].
2.3.10.4 Creep rupture properties of 12Cr-1Mo-1W-V-Nb steel tubes
2.3.10.4.1 Creep rupture data of 12Cr-1Mo-1W-V-Nb steel tubes
Fig. 226 shows creep rupture data of 12Cr-1Mo-1W-V-Nb steel tubes with average curves according to
the Larson-Miller parameter method shown in Fig. 227 [1]. The longest creep rupture time of 12Cr-1Mo1W-V-Nb steel tubes is about 60000 h at 650 °C. Their long term creep strengths are very stable in the
temperature range between 500 and 700 °C. Fig. 228 shows a Larson-Miller parameter plot of the creep
rupture data of 12Cr-1Mo-1W-V-Nb steel tubes with a master rupture curve and a 95 % confidence lower
limit. The best fitting was achieved with an optimized constant of 30.21.
Landolt-Börnstein
New Series VIII/2B
182
2.3 High Cr steels
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
0
100
200
300 400 500
Temperature [°C]
600
700
Fig. 225. Tensile strength (circles) and 0.2% proof
stress (triangles) data of 12Cr-1Mo-1W-V-Nb steel
tubes.
500
400
Stress [MPa]
300
500 °C
200
550 °C
600 °C
100
500 °C
550 °C
600 °C
650 °C
700 °C
average curve
80
60
40
1
650 °C
Fig. 226. Creep rupture data of
12Cr-1Mo-1W-V-Nb steel tubes.
700 °C
10 2
10 3
Rupture time [ h]
10
10 4
10 5
500
500 °C
×10 5h
400
Stress [MPa]
300
550 °C
×10 5h
600 °C
×10 5h
200
650 °C
×10 5h
100
80
60
40
24
500 °C
550 °C
600 °C
650 °C
700 °C
25
average curve
minimum curve
26
27
28
29
Fig. 227. Larson-Miller parameter plot of the creep
rupture data of 12Cr-1Mo-1W-V-Nb steel tubes.
30
31
32
33
34
Landolt-Börnstein
New Series VIII/2B
Ref. p. 188]
2.3.10 12Cr-1Mo-1W-V-Nb steel
183
2.3.10.4.2 Creep data of 12Cr-1Mo-1W-V-Nb steel tubes
Fig. 228 shows minimum creep rate of 12Cr-1Mo-1W-V-Nb steel tubes measured at various stress levels
between 500 °C and 700 °C with fitted curves according to the Larson-Miller parameter method. Fig. 229
shows a Larson-Miller parameter plot of the minimum creep rate of 12Cr-1Mo-1W-V-Nb steel tubes with
a master minimum creep rate curve. The best fitting was achieved with an optimized constant of 48.32.
500
400
Stress [MPa]
300
200
500 °C
500 °C
550 °C
600 °C
650 °C
700 °C
average curve
550 °C
600 °C
100
80
60
650 °C
700 °C
40
10 -2
10 -1
500
10 2
10 3
500 °C
0.01%/10 3h
400
550 °C
0.01%/10 3h
300
Stress [MPa]
1
10
Minimum creep rate [% / 10 3 h]
Fig. 228. Minimum creep rate
data of 12Cr-1Mo-1W-V-Nb
steel tubes
600 °C
0.01%/10 3h
200
650 °C
0.01%/10 3h
100
80
60
500 °C
550 °C
600 °C
650 °C
700 °C
average curve
40
34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49
Larson-Miller-parameter T (48.32 + log e ) [×10 -3 ]
Fig. 229. Larson-Miller papameter plot of the minimum
creep rate data of 12Cr-1Mo-1W-V-Nb steel tubes.
2.3.10.5 Allowable stress of 12Cr-1Mo-1W-V-Nb steel tubes and pipe
Fig. 230 shows the allowable tensile stress determined for 12Cr-1Mo-1W-V-Nb steel tubes (Japanese
METI KA-SUS410J2TB) according to the METI standard procedure comparing with that for
conventional 9%Cr steel, ASME SA213-T9 (JIS STBA26).
Landolt-Börnstein
New Series VIII/2B
184
2.3 High Cr steels
180
Allowable tensile stress (MPa)
160
KA-SUS410J2TB
140
120
100
STBA26
80
60
40
Fig. 230. Allowable tensile stress determined for 12Cr1Mo-1W-V-Nb steel tubes (Japanese METI KASUS410J2TB).
20
0
0
100 200 300 400 500 600 700 800
Temperature (℃)
2.3.10.6 Alloying philosophy of 12Cr-1Mo-1W-V-Nb steel tubes
Fig. 231 shows the alloying philosophy of 12Cr-1Mo-1W-V-Nb steel tubes. The steel has been developed
for superheater and reheater tubes used in plants operated at high temperatures between 540 °C and
625 °C. 12%Cr matrix is chosen for corrosion resistance to steam oxidation and hot corrosion at high
temperatures. High temperature creep strength is achieved by using Mo and W in solid solution and fine
precipitation of VC (or (V,Nb)(C,N)) carbide. To optimize the weldability and formability, the C content
is reduced, which causes formation of delta-ferrite to some extent.
Fig. 232 shows change in Charpy impact values with respect to the amount of delta-ferrite in 12Cr
model steels. It can be seen that the toughness of the steels can be maintained reasonably high when
keeping the amount of delta-ferrite below 35 %. Fig. 233 shows the change in the amount of delta-ferrite
with respect to Creq calculated for the model steels. It is concluded that the Creq of the steel should be kept
below 12.5 to suppress too much delta-ferrite and the resultant poor toughness.
Service steam temperature; up to 625℃
High temperature strength
V, Nb
Mo, W
Weldability
Toughness
0.2V-0.05/0.1Nb
Corrosion resistance
12%Cr Steel
Weldability
Formability
Martensitic steel
0.2C
0.1C
Dual phase steel
δ phase ＜ 35%
0.10C-12Cr-1Mo-0.25V-0.05/0.1Nb
Ac1 = 886℃
1050℃Norma. + 800/830℃ Temper
Long-term creep strength
Fig. 231. Alloying philosophy of
12Cr-1Mo-1W-V-Nb steel tubes.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 188]
2.3.10 12Cr-1Mo-1W-V-Nb steel
200
160
140
60
120
δ - ferrite (%)
Impact value at 20℃ (J/cm2 )
Creq = Cr+6Si+4Mo+1.5W+11V+5Nb+12sol.Al+8Ti
- (40C+30N+2Co+Cu+4Ni+2Mn)
80
δ = 35%
180
185
100
80
60
40
20
0
40
δ = 35%
20
0
20
40
60
80
δ - ferrite (%)
100
Fig. 232. Change in Charpy impact values with respect
to the amount of δ -ferrite in 12Cr model steels.
0
5
10
15
Creq (mass%)
20
25
Fig. 233. Change in the amount of delta-ferrite with
respect to Creq calculated for the model steels.
Fig. 234 shows effects of Nb and V contents on the estimated 10000 h creep rupture strength of 12Cr1Mo-1W-V-Nb model steels at 600 °C. Highest creep strength is obtained by the combination of V from
0.2 % to 0.3 % and Nb from 0 % to 0.1 %. Fig. 235 shows effects of Mo and W contents on creep rupture
strength of 12Cr-Mo-W-V-Nb model steels at 600 °C. The creep rupture strength increases with
increasing Moeq (=Mo+1/2W). It is found that the toughness decreases more than 1.5 % with increasing
Moeq which is due to increase in the amount of delta-ferrite. It is therefore concluded that the combination
of 1 % Mo and 1 % W is the best for creep strength with enough toughness.
0.5
600 °C ×104 h
creep rupture
strength [MPa]
Maximum
0.4
V content [mass %]
126
0.3
142
132
152
140
151
0.2
133
147
139
120
101
127
137
123
127
123
0.1
0
0
0.1
0.2
Nb content [mass %]
0.3
0.4
Fig. 234. Effects of Nb and V contents on the estimated
10000 h creep rupture strength of 12Cr-1Mo-1W-V-Nb
model steels at 600 °C.
2.3.10.7 Microstructure of 12Cr-1Mo-1W-V-Nb steel tubes
Fig. 236 shows the microstructure of 12Cr-1Mo-1W-V-Nb steel tubes normalized at 1050 °C and
tempered at 815 °C. The microstructure of the steel consists of about 20 % delta-ferrite and tempered
martensite.
Landolt-Börnstein
New Series VIII/2B
186
2.3 High Cr steels
10 5
Mo + 1/2 W = 1.5
5
Creep rupture time [h]
(A)
10 4
(B)
5
Increase of
δ-ferrite
lower toughness
10 3
A
5
10 2
0
1
2
3
Mo + 1/2 W [ mass %]
B
-
Element
W
Mo
Mo + W
4
5
Fig. 235. Effects of Mo and W contents on creep rupture
strength of 12Cr-Mo-W-V-Nb model steels at 600 °C.
Steel A: 0.1C-12CrMoWVNb, 600 °C × 147 MPa
Steel B: 0.2C-12CrMoWV, 600 °C × 152 MPa
Fig. 236. Optical micrographs of 12Cr-1Mo-1W-V-Nb steel tubes normalized and tempered
showing the duplex micro-structure consisting of delta-ferrite and tempered martensite.
Fig. 237 shows a typical CCT (Continuous Cooling Transformation) diagram determined for 12Cr-1Mo1W-V-Nb steel after heating at 1050 °C. Martensitic microstructure with delta ferrite is easily obtained
even with a fairly slow cooling. It is also noted that the Ac1 temperature is very high, which is
advantageous for establishing a long-term microstructural stability by tempering at higher temperatures.
Fig. 238 shows the effect of tempering temperature on long term creep strength of 12Cr-1Mo-1W-VNb steel tubes. With increasing testing temperature and decreasing stress level the creep rupture lives of
the specimens tempered at 800 °C becomes much longer than those of the ones tempered at 750 °C.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 188]
2.3.10 12Cr-1Mo-1W-V-Nb steel
1000
187
Ac3 931 °C
900
Ac1 886 °C
Temperature T [°C]
800
F+C
700
600
500
400
Ms
300
Vickers hardness
200
349
352
100
10
344
329
279
293 275 199 180
10 3
10 4
10 2
Cooling time from Ac 3 temperature [s]
10 5
Fig. 237. A typical CCT diagram
determined for 12Cr-1Mo-1W-VNb steel after heating at 1050 °C.
400
300
600 °C
Stress [MPa]
200
650 °C
100
80
60
600 °C 650 °C Temperature ×1h AC
750 °C
800 °C
40
10
10 2
10 3
Time to rupture [ h]
10 4
10 5
Fig. 238. Effect of tempering
temperature on long term creep
strength of 12Cr-1Mo-1W-V-Nb
steel tubes.
2.3.10.8 Performance of service exposed tubes
Fig. 239 shows changes in tensile properties and toughness of 12Cr-1Mo-1W-V-Nb steel tubes after
service exposure for up to 16 years in the Wakayama Kyodo Power Station No.3 boiler which has been
operated with outlet steam temperatures of 543 °C for reheater (RH) and 571 °C for superheater (SH)
tubes. The tensile properties after the service exposure have not changed much. Toughness reduced to
some extent after 1 year exposure but saturated to high level even after 16 years exposure. More detailed
performance of the service exposed tubes is available in [3].
Landolt-Börnstein
New Series VIII/2B
2.3 High Cr steels
800
80
600
60
400
40
200
20
0
Elongation [%]
Tensile strength, Yield stress [MPa]
188
0
SH
RH
Impact value at 0 °C [J/cm 2 ]
100
80
60
40
20
0
0
5
10
Service time [a]
15
Fig. 239. Change in tensile properties and toughness of
12Cr-1Mo-1W-V-Nb steel tubes after service exposure at
the Wakayama Kyodo Power Station No.3 boiler. Open
circles: SH; Full circles: RH.
2.3.10.9 References
[1] Sumitomo seamless tubes and pipe Creep Data Sheets, Sumitomo Metal Industries, (1993).
[2] Iseda, A., Sawaragi, Y., Teranishi, H., Kubota, M., and Hayase, Y.: The Sumitomo Search No.40
(1989), 41.
[3] Iseda, A., Sawaragi, Y., Ogawa, K., Kubota, M., and Hayase, Y.: The Sumitomo Search No.48
(1992), 21.
[4] Nishimura, N., and Masuyama, F.: Materials for Advanced Power Engineering, Part 1, Kluwer
Academic Publishers (1994), 351.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 191]
2.3.11 12Cr-1Mo-Ni-V cast
189
2.3.11 12Cr-1Mo-Ni-V cast
2.3.11.1 Introduction
The 12Cr-1Mo-Ni-V cast steel grade GX23CrMoV12-1 (EN 10213-2, material-no 1.4931) was
introduced in the 1960s. Among the steel grades specified in EN 10213-2:1995, GX23CrMoV12-1 shows
the maximum creep rupture strength values and the highest oxidation resistance. The excellent creep
rupture strength is mainly related to a stable microstructure composed of tempered martensite and small
vanadium carbides. If the wall thickness is less than about 100 mm quenching in air is sufficient to
produce a martensite microstructure. Lower cooling rates from austenite temperature promote the
transformation to ferrite and carbides rather than martensite. Large amounts of ferrite, however, are
detrimental to the creep rupture properties, and may reduce the creep rupture strength values by about 30
% at temperatures above 530 °C. For this reason sections thicker than about 100 mm have to be quenched
in oil. The high oxidation resistance at temperatures up to 600 °C is ensured by the high chromium
content of 11.3 to 12.2 %. The main features of GX23CrMoV12-1 are summarized below:
Melting process: Electric arc, argon oxygen decarburization, induction melting.
Heat treatment: Quenched in air or oil, and tempered (cooling in furnace).
Typical microstructure: Tempered martensite.
Weldability: Weldable with similar weld metal; the welding should be performed in the austenite or
austenite/martensite temperature range; after welding cooling below 100 °C is necessary before stress
relief annealing.
High temperature applications: Casings of steam turbines, compressors, gas turbines, valves, nozzles;
service temperatures up to about 600°C.
2.3.11.2 Standard requirements
Table 129. Chemical composition
Chemical composition [wt%]
Standard Designation
C
Si
Mn
P
S
EN
10213-2:
1995
GX23CrMoV12-1 0.20(1.4931)
0.25
0.50≤
0.40 0.80
Cr
11.30- 1.00≤
≤
0.030 0.020 12.20 1.20
Table 130. Heat treatment and tensile properties at room temperature
StanDesignation
Heat treatment
Thickness Min. 0.2 %
dard
proof strength
[mm]
[MPa]
EN
GX23CrMoV12-1 Q:1030°C-1080°C
150
540
10213(1.4931)
T:700°C-780°C
2:1995
Landolt-Börnstein
New Series VIII/2B
Mo
Ni
V
0.25≤
1.00 0.35
Other
W:
≤0.50
Tensile Min. elongastrength tion at rupture
[MPa] [%]
740 880
15
190
2.3 High Cr steels
min_EN
Rm
800
1000
600
800
Rm (MPa)
R p0.2 (MPa)
Rp0.2
400
200
0
600
400
200
0
0
100
200
300
400
500
600
0
100
200
300
400
500
600
Temperature (°C)
Temperature (°C)
Fig. 240. Tensile properties Rp0.2 and Rm of the test materials of cast steel grade GX23CrMoV12-1 creep rupture
tested by the German Creep Committee; [1]. min EN: minimal values by EN 10213-2.
Stress (MPa)
1000
broken
unbroken
500°C_EN
100
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 241. Creep rupture strength data of cast steel grade GX23CrMoV12-1 at 500 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995.
Stress (MPa)
1000
broken
100
unbroken
550°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 242. Creep rupture strength data of cast steel grade GX23CrMoV12-1 at 550°C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:1995
Landolt-Börnstein
New Series VIII/2B
Ref. p. 191]
2.3.11 12Cr-1Mo-Ni-V cast
191
Stress (MPa)
1000
broken
100
unbroken
600°C_EN
10
10
100
1000
10000
100000
1000000
Test duration (h)
Fig. 243. Creep rupture strength data of cast steel grade GX23CrMoV12-1 at 600 °C obtained by the German Creep
Committee [1], and average creep rupture strength values indicated in EN 10213-2:199.
2.3.11.3 Average creep rupture strength
Table 131. Average creep rupture strength values indicated in EN 10213-2:1995.
Time to rupture
Temperature
10000
h
100000 h
200000 h
[°C]
Average creep rupture strength [MPa]
400
504
426
394
450
383
309
279
500
269
207
187
550
167
118
103
600
83
49
39
2.3.11.4 Reference
[1] Results of German long term creep rupture tests; Contribution to the Landolt-Börnstein Creep Data
Book; Cast steel grade GX23CrMoV12-1, compilation of test results; Forschungsvereinigung
Warmfeste Stähle, c. o. Verein Deutscher Eisenhüttenleute, Düsseldorf (D), (2001).
Landolt-Börnstein
New Series VIII/2B
192
2.3 High Cr steels
2.3.12 11Cr-0.4Mo-2W-Cu-V-Nb steel
2.3.12.1 Introduction
11Cr-0.4Mo-2W-Cu-V-Nb ferritic steel (T122, P122; HCM12A) is used as superheater and reheater
tubes, and header and main steam pipe in fossile fired boilers. The steel has been developed for improving
creep rupture strength and corrosion resistance of T91 type 9%Cr steels at elevated temperatures. The
microstructure consists of tempered martensitic matrix strengthened by M23C6 carbide mainly along grain
boundaries and fine dispersed MX such as (V, Nb)(C, N) carbonitride in matrix. MX is fine and stable
even after long term creep exposure at high temperatures.
2.3.12.2 Material standards, and chemical and tensile requirements
Tables 132 and 133 give the chemical requirements and the corresponding tensile requirements of 11Cr0.4Mo-2W-Cu-V-Nb steel tubes and pipes which are designated by the standards: Japanese KASUS410J3TB, KA-SUS410J3TP, ASTM A213-T122, A335-P122, ASME Sec. I CC 2180.
Table 132. Chemical requirements of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes and pipes; Japanese KASUS410J3TB, KA-SUS410J3TP, ASTM A213-T122, A335-P122, ASME Sec. I CC 2180.
Designation
Japanese
METI
ASTMA213
ASTMA335
Grade
(1)
(2)
T122
P122
Std. No.
Cu
0.30
1.70
0.30 ASME
SecI
0.14 0.50 0.70 0.020 0.010 0.50 12.50 0.60 2.50 0.30 0.10 0.100 0.040 0.005 1.70 CC2180
C
0.07
0.14
0.07
Si
0.50
-
Mn
0.70
-
P
0.020
-
S
0.010
-
Chemical composition [wt%]
Ni
Cr
Mo W
V
10.00 0.25 1.50 0.15
0.50 11.50 0.60 2.50 0.30
10.00 0.25 1.50 0.15
Nb
0.04
0.10
0.04
N
0.040
0.100
0.040
Al
0.040
-
B
0.005
-
(1): KA-SUS410J3TB
(2): KA-SUS410J3TP
Table 133. Tensile requirements of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes and pipes; Japanese KASUS410J3TB, KA-SUS410J3TP, ASTM A213-T122, A335-P122, ASME Sec. I CC 2180.
Min.
Standard No.
Designation
Grade
Min. TS1) Min. YS2)
elongation
KA-SUS410J3TB
Japanese METI
620 MPa
400 MPa
20 %
KA-SUS410J3TP
ASTM A213
T122
ASME Sec. I
620 MPa
400 MPa
20 %
CC 2180
ASTM A335
P122
1) TS; tensile strength, 2) YS; yield strength as 0.2% proof stress
2.3.12.3 Tensile properties of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes
2.3.12.3.1 Tensile properties of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes
Fig. 244 shows tensile strength and yield stress data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes [1]. They
are higher than those of T91 type 9%Cr steels in all the temperatures up to 700°C. The corresponding
tensile elongation and reduction of area data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes are available in the
literatures [1,4].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 199]
2.3.12 11Cr-0.4Mo-2W-Cu-V-Nb steel
193
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
0
100
200
300 400 500
Temperature [°C]
600
700
800
Fig. 244. Tensile strength (circles) and yield stress
(triangles) data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes.
2.3.12.3.2 Tensile properties of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes
Fig. 245 shows tensile strength and yield stress data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes [1]. They
are higher than those of P91 type steels at temperatures up to 700 °C. The corresponding tensile
elongation and reduction of area data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes are available in [1].
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
0
100
200
300 400 500
Temperature [°C]
600
700
800
Fig. 245. Tensile strength (circles) and yield stress
(triangles) data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes.
2.3.12.4 Creep rupture properties of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes and pipes
2.3.12.4.1 Creep rupture data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes
Fig. 246 shows creep rupture data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes with average curves assessed
by the Larson-Miller parameter method [1]. The longest creep rupture time of 11Cr-0.4Mo-2W-Cu-V-Nb
steel tubes is about 30000 h at 600 °C. Their long term creep strength is stable at temperatures between
550 °C and 650 °C. Fig. 247 shows a Larson-Miller parameter plot of creep rupture data of 11Cr-0.4Mo2W-Cu-V-Nb steel tubes with a master rupture curve and a 95 % confidence lower limit. The best fitting
was achieved with an optimized constant of 34.64.
Landolt-Börnstein
New Series VIII/2B
194
2.3 High Cr steels
500
400
300
550 °C
Stress [MPa]
200
600 °C
100
650 °C
80
60
40
30
1
550 °C
600 °C
650 °C
700 °C
average curve
700 °C
10 2
10 3
Rupture time [ h]
10
10 4
10 5
Fig. 246. Creep rupture data of
11Cr-0.4Mo-2W-Cu-V-Nb
steel
tubes.
500
550 °C
×10 5h
400
600 °C
×10 5h
300
Stress [MPa]
200
650 °C
×10 5h
100
80
60
40
30
29
550 °C
600 °C
650 °C
700 °C
average curve
minimum curve
Fig. 247. Larson-Miller parameter plot of the creep
rupture data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes.
30 31 32 33 34 35 36 37 38 39
Larson-Miller-parameter T (34.64 + log t ) [×10 -3 ]
2.3.12.4.2 Creep rupture data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes
Fig. 248 shows creep rupture data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes with average curves assessed
by the Larson-Miller parameter method [1]. The longest creep rupture time of 11Cr-0.4Mo-2W-Cu-V-Nb
steel pipes is about 45000 h at 650 °C. Their long term creep strength is very stable at temperatures
between 550 °C and 650 °C. Fig. 249 shows a Larson-Miller parameter plot of creep rupture data of
11Cr-0.4Mo-2W-Cu-V-Nb steel pipes with a master rupture curve and a 95 % confidence lower limit.
The best fitting was achieved with an optimized constant of 35.20.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 199]
2.3.12 11Cr-0.4Mo-2W-Cu-V-Nb steel
195
500
400
300
550 °C
Stress [MPa]
200
600 °C
100
650 °C
80
550 °C
600 °C
650 °C
700 °C
average curve
60
40
30
1
700 °C
10 2
10 3
Rupture time [ h]
10
10 4
Fig. 248. Creep rupture data of
11Cr-0.4Mo-2W-Cu-V-Nb steel
pipes.
10 5
500
550 °C
×10 5h
400
600 °C
×10 5h
300
Stress [MPa]
200
650 °C
×10 5h
100
80
60
550 °C
600 °C
650 °C
700 °C
average curve
minimum curve
40
30
29
Fig. 249. Larson-Miller parameter plot of creep rupture
data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes.
30 31 32 33 34 35 36 37 38 39
Larson-Miller-parameter T (35.20 + log t ) [×10 -3 ]
2.3.12.4.3 Creep data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes
Fig. 250 shows minimum creep rate data of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes measured at various
stress levels at temperatures between 550 °C and 700 °C with curves fitted by the Larson-Miller
parameter method [1]. Fig. 251 shows a Larson-Miller parameter plot of the minimum creep rate data of
11Cr-0.4Mo-2W-Cu-V-Nb steel tubes with a master minimum creep rate curve. The best fitting was
achieved with an optimized constant of 49.86.
2.3.12.4.4 Creep data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes
Fig. 252 shows minimum creep rate data of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes measured at various
stress levels at temperatures between 550 °C and 700 °C with curves fitted by the Larson-Miller
parameter method [1]. Fig. 253 shows a Larson-Miller parameter plot of the minimum creep rate data of
11Cr-0.4Mo-2W-Cu-V-Nb steel pipes with a master minimum creep rate curve. The best fitting was
achieved with an optimized constant of 41.42.
Landolt-Börnstein
New Series VIII/2B
196
2.3 High Cr steels
500
400
500
550 °C
0.01% /10 3h
400
300
300
600 °C
0.01% /10 3h
550 °C
200
Stress [MPa]
Stress [MPa]
200
600 °C
100
80 650 °C
100
80
550 °C
600 °C
650 °C
700 °C
average curve
60
40 700 °C
30
10-2
650 °C
0.01% /10 3h
1
10
10-1
10 2
3
Minimum creep rate [% / 10 h]
550 °C
600 °C
650 °C
700 °C
60
40
10 3
Fig. 250. Minimum creep rate data of 11Cr-0.4Mo2W-Cu-V-Nb steel tubes.
average curve
30
39
40 41 42 43 44 45 46 47 48 49
Larson-Miller-parameter T (49.86 + log e ) [×10 -3 ]
Fig. 251. Larson-Miller parameter plot of the minimum
creep rate of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes.
500
400
500
300
300
550 °C
0.01% /10 3h
400
600 °C
0.01% /10 3h
550 °C
200
Stress [MPa]
Stress [MPa]
200
600 °C
100
80 650 °C
60
40
30
10-2
700 °C
550 °C
600 °C
650 °C
700 °C
average curve
1
10-1
10
10 2
Minimum creep rate [% / 103h]
10 3
Fig. 252. Minimum creep rate data of 11Cr-0.4Mo-2WCu-V-Nb steel pipes.
650 °C
0.01% /10 3h
100
80
60
40
30
39
550 °C
600 °C
650 °C
700 °C
average curve
40 41 42 43 44 45 46 47 48 49
Larson-Miller-parameter T (41.42 + log e ) [×10 -3 ]
Fig. 253. Larson-Miller parameter plot of the minimum
creep rate of 11Cr-0.4Mo-2W-Cu-V-Nb steel pipes.
2.3.12.5 Allowable stress of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes and pipe
Figs. 254 and 255 show the allowable tensile stresses determined for 11Cr-0.4Mo-2W-Cu-V-Nb steel
tubes and pipes (Japanese METI KA-SUS410J3TB and KA-SUS410J3TP) according to the METI
standard procedure comparing with those for the conventional steels ASME SA213-T91 and SA335-P91
(Japanese METI KA-STBA28 and KA-STPA28).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 199]
2.3.12 11Cr-0.4Mo-2W-Cu-V-Nb steel
180
180
160
160
KA-SUS410J3TB
140
120
Allowable tensile stress (MPa)
Allowable tensile stress (MPa)
197
KA-STBA28
100
80
60
40
KA-SUS410J3TP
140
120
KA-STBA28
100
80
60
40
20
20
0
0
0
0
100 200 300 400 500 600 700 800
100 200 300 400 500 600 700 800
Temperature (℃)
Temperature (℃)
Fig. 254. Allowable tensile stress determined for 11Cr0.4Mo-2W-Cu-V-Nb steel tubes (Japanese METI KASUS410J3TB).
Fig. 255. Allowable tensile stress determined for 11Cr0.4Mo-2W-Cu-V-Nb steel pipe (Japanese METI KASUS410J3TP).
2.3.12.6 Alloying philosophy of 11Cr-0.4Mo-2W-Cu-V-Nb steel
11Cr-0.4Mo-2W-Cu-V-Nb steel has been developed for improving creep rupture strength and corrosion
resistance of P91 type 9%Cr steels above 600 °C, which is mainly achieved by a higher Cr content and
substitution of a part of Mo by W. It is also noted that in order to suppress δ - ferrite formation for thick
wall pipes Cu addition is chosen among the γ forming elements shown in Fig. 256. Cu is the γ forming
element which does not reduce Ac1 temperature much and does not enhance coarsening of M23C6 carbide
unlike Ni and Mn. Cu addition allows the combination of a higher Cr content with high W and Mo
contents.
40
Base; 0.1C-11Cr steel
Addition [mass %]
a forming element
2W
DAc 1 [°C ]
1Cr
1Mo
0
2Cu
0.5Ni
-40
-4
0.05N
Cr eq = Cr+6Si+4Mo+1.5W
+11V+5Nb+8Ti+12Al
-40C-30N-4Ni-2Mn
-Cu-2Co [mass %]
Fig. 256. Comparison of alloying elements with respect to
changes in Ac1 temperature and Creq of 0.1C-11Cr model
steels.
g forming element
-2
Landolt-Börnstein
New Series VIII/2B
0
D Cr eq [mass %]
2
4
198
2.3 High Cr steels
2.3.12.7 Microstructural change of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes
Fig. 257 shows a typical CCT diagram determined for 11Cr-0.4Mo-2W-Cu-V-Nb steel after heating at
1050 °C. Full martensitic microstructure is easily obtained even with a very slow cooling. It is noted that
Ac1 temperature is over 800 °C, which is advantageous for a long-term microstructural stability by taking
tempering at higher temperatures.
1000
Ac3 904 °C
Austenitized at 1050 °C × 5 min
900
Ac1 805 °C
Temperature T [°C]
800
700
600
500
400
300
200
100
1
Vickers hardness
10
428 422
437 425 420
416
433 452
481
2
3
10
10
Cooling time from Ac 3 temperature [s]
10 4
408
10 5
Fig. 257. A typical CCT diagram
determined for 11Cr-0.4Mo-2WCu-V-Nb steel after heating at
1050 °C.
Fig. 258 shows TEM micrographs of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes normalized and tempered,
and service exposed for 20508 h in the Wakayama Kyodo Power Station No.3 boiler which has been
operated with an outlet steam temperature of 571 °C [8]. It can be seen that in the specimen normalized
and tempered VN type MX ((V,Nb)(C,N)) is finely dispersed in matrix with coherency strain, while in the
specimen service exposed fine VN still remains in matrix with a smaller coherency strain.
2.3.12.8 Performance of service exposed tubes
More detailed performance data of service exposed tubes is available in [7] and [8].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 199]
2.3.12 11Cr-0.4Mo-2W-Cu-V-Nb steel
ＶＮ
ＶＮ
110 α
100 α
010 α
ＶＮ
100nm
100nm
(a) Bright field image
100 α
199
(b) Dark field image
010 α
VN
Laves
110 α
100nm
(c) Bright field image
100nm
(d) Dark field image
Fig. 258. TEM micrographs of 11Cr-0.4Mo-2W-Cu-V-Nb steel tubes normalized and tempered (a and b), and
service exposed for 20508 h (c and d) at 571 °C.
2.3.12.9 References
[1] Sumitomo seamless tubes and pipe Creep Data Sheets, Sumitomo Metal Industries, (1993).
[2] Sawaragi, Y., Iseda, A., Ogawa, K., and Masuyama, F.: Materials for Advanced Power Engineering,
Part 1, Kluwer Academic Publishers (1994), p.309.
[3] Iseda, A., Sawaragi, Y., Kato, S., and Masuyama, F.: Proc. of the Fifth International Conf. on Creep
Materials, Florida, (1992), 389.
[4] Sawaragi, Y., Igarashi, M., Iseda, A., Yamamoto, S., and Masuyama, F.: Sumitomo Kinzoku Vol.47
No.4 (1995), 29.
[5] Ogawa, K., Iseda, A., Sawaragi, Y., Matsumoto, S., and Masuyama, F.: Sumitomo Kinzoku Vol.47
No.4 (1995), 39.
[6] Sawaragi, Y., Kan, T., Yamadera, Y., Masuyama, F., Yokoyama, T., and Komai, N.: Proc. of the 6th
International Conf. on Materials for Advanced Power Engineering (1998) Liege, Forschungszentrum
Julich GmbH, 62.
[7] Sawaragi, Y., Miyata, K., Yamamoto, S., Masuyama, F., Komai, N., and Yokoyama, T.: Advanced
Heat Resistant Steels For Power Generation, The University Press, Cambridge (1998), p.144.
[8] Miyata, K., Sawaragi, Y., Okada, H., Masuyama, F., Yokoyama, T., and Komai, N.: ISIJ
International, Vol. 40 (2000), 1156.
Landolt-Börnstein
New Series VIII/2B
200
2.3 High Cr steels
2.3.13 12Cr-2.6W-2.5Co-0.5Ni-V-Nb steel
2.3.13.1 Introduction
12Cr-2.6W-2.5Co-0.5Ni-V-Nb steel (NF12) was developed by Nippon Steel Corporation, Japan in the
mid-1990s for boiler tubing and pipe work applications for power plants with high steam parameters. The
targeted creep rupture strength of NF12 steel is at least 1.3 times higher than that of NF616 (P92) steel.
Demonstrative fabrication of full size header and main steam pipe components has been conducted,
although the steel has not yet been used for practical applications.
2.3.13.2 Material standards, chemical and tensile requirements [1]
NF12 steel is not specified in any codes or standards. Table 134 shows the chemical compositions of
tubes and plates produced from NF12 steel on a trial basis and subjected to creep and characterization
tests. The test materials were melted in a vacuum induction furnace or in an electric furnace and then cast
into 20 kgf to 1000 kgf weight ingots. Hot rolling equipment produced 15 mm thick plate specimens and
a hot extrusion machine formed seamless tube specimens with wall thickness from 8.1 to 9.0 mm and
outer diameter from 45 to 50.8 mm. Specimens were austenitized in the range from 1070 °C to 1100 °C,
quenched in air, and then tempered in the range from 770 to 800 °C.
Figs. 259 and 260 show elevated temperature tensile properties and Charpy impact properties of NF12
steel, respectively.
Table 134. Chemical compositions of NF12 steels [1].
Chemical composition [wt%]
Product
C
Si
Mn
Cr
Mo
W
Nb
V
Tube
0.092 0.03
0.51
10.79 0.13
2.47
0.064 0.20
Tube
0.094 0.19
0.50
11.35 0.16
2.57
0.070 0.20
Plate
0.077 0.26
0.57
12.11 0.15
2.75
0.071 0.23
Plate
0.078 0.22
0.47
10.85 0.14
2.60
0.068 0.20
Ni
0.51
0.41
0.58
0.51
Co
1.95
1.49
2.68
1.91
B
0.002
0.004
0.083
0.003
N
0.046
0.040
0.050
0.048
1000
Yield strength, Tensile strength (MPa)
Tensile strength
Yield strength
(0.2% offset)
800
600
400
200
0
Fig. 259. Elevated temperature
tensile properties of NF12 steel;
[1].
0
100
200
300
400
500
600
700
800
Temperature (℃)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 203]
2.3.13 12Cr-2.6W-2.5Co-0.5Ni-V-Nb steel
100
250
200
80
Toughness
Crystallinity
150
60
100
40
50
20
0
-80
-60
-40
-20
0
20
Temperature (℃)
40
Cristallinity (%)
Charpy impact value (J/cm2)
201
Fig. 260.
Charpy impact
properties of NF12 steel; [1].
0
80
60
2.3.13.3 Creep properties
2.3.13.3.1 Creep rupture data [1]
Fig. 261 shows the creep rupture stress vs. time to rupture relationship for NF12 steel tubes and plates at
temperatures from 600 °C to 750 °C. Fig. 262 shows creep rupture elongation and reduction of area
against time to rupture. There is no major difference in creep rupture ductilities among the different
testing temperatures of NF12 steel.
1000
700
500
Stress (MPa)
300
200
600℃
100
70
50
650℃
700℃
30
750℃
20
10
100
101
102
103
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
104
Fig. 261.
Creep rupture
strength data of NF12 steel;
[1].
105
202
2.3 High Cr steels
100
600℃
Elongation, Reduction of area (%)
90
650℃
80
700℃
Reduction of area
70
60
750℃
Elongation
50
40
30
Fig. 262.
Creep rupture
elongation and reduction of
area of NF12 steel; [1].
20
10
0
100
101
102
103
104
105
Time to rupture (h)
2.3.13.3.2 Estimated creep strength [1]
Fig. 263 shows the extrapolated 100,000 h creep rupture strength of NF12 steel compared with those of
NF616 (P92) steel and Mod.9Cr-1Mo (T91) steel by means of the Larson-Miller parametric method. The
extrapolated creep rupture strength for NF616 steel is about 1.3 times higher than that of Mod.9Cr-1Mo
steel. The creep rupture strength of NF12 steel is expected to be about 1.3 to 1.5 times higher than that of
NF616 steel, depending on the extrapolation method. It can thus be inferred that the creep rupture
strength of NF12 steel is around 1.6 times higher than that of Mod.9Cr-1Mo, and sometimes up to 1.9
times higher.
300
Estimated stress for 105h (MPa)
NF12
250
200
NF616
(P92)
150
Mod. 9Cr-1Mo
(P91)
100
Fig. 263. Extrapolated creep
rupture strength of NF12 steel
comparing with NF616 steel and
Mod. 9Cr-1Mo steel; [1].
50
0
500
550
600
Temperature (℃)
650
700
The allowable stresses of NF12 steel were developed based on the ASME criteria and are compared with
NF616 steel and Mod.9Cr-1Mo steel in Fig. 264. The allowable stresses for NF616 steel and NF12 steel
tend to be much higher compared with those of Mod.9Cr-1Mo steel in the relatively high temperature
region. At 600 °C the allowable stress for NF616 steel is 1.3 times higher than that of Mod.9Cr-1Mo
steel, and the stress for NF12 steel is 1.6 times higher than that for Mod.9Cr-1Mo steel.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 203]
2.3.13 12Cr-2.6W-2.5Co-0.5Ni-V-Nb steel
203
200
NF12
180
NF616 (P92)
Allowable stress (MPa)
160
140
Mod. 9Cr-1Mo (P91)
120
100
80
60
Fig. 264. Allowable stresses of
NF12 steel comparing with
NF616 steel and Mod. 9Cr-1Mo
steel; [1].
40
20
0
0
100
200
300
400
Temperature (℃)
500
600
700
2.3.13.4 Reference
[1] Ohgami, M., Hasegawa, Y., Naoi, H., and Fujita, T.: IMechE Conf. Trans. 1997-2, Advanced Steam
Plant-New Materials and Plant Design and Their Practical Implications for Future CCGT and
Conventional Power Stations, (1997), 115.
Landolt-Börnstein
New Series VIII/2B
204
2.3 High Cr steels
2.3.14 12Cr-3W-3Co-V-Nb-Ta-Nd-N steel
2.3.14.1 Introduction
12Cr-3W-3Co-V-Nb-Ta-Nd-N steel (SAVE12) was developed by Sumitomo Metal Industries, Ltd., Japan
in the mid-1990s for thick wall pipes applicable to main steam pipes and headers of power plants with
high steam parameters. This steel has higher creep rupture strength at elevated temperatures up to 630 °C
than Mod.9Cr-1Mo steel (P91), NF616 (P92) steel and HCM12A (P122) steel. To date, laboratory heat
has been used for studies of creep strength and for characterization tests, and the steel has not yet been
used for practical applications. However, several lengths of tubings have been installed in a large capacity
power boiler for field exposure testing.
2.3.14.2 Material standards, chemical and tensile requirements
SAVE12 steel is not specified in any codes or standards. Table 135 [1] shows ranges of chemical
compositions of SAVE12 steel vacuum induction melted in the laboratory as a 150 kgf ingot for creep
strength studies, and subsequently processed by hot forging into 25 mm thick plates. The plates were
normalized for 1 hour at temperatures between 1000 °C and 1200 °C, air cooled, and tempered at 550 °C
to 800 °C, followed again by air-cooling. The plates subjected to creep rupture tests were normalized for
0.5 h at 1050 °C and tempered for 1.5 h at 780 °C.
Fig. 265 [1] shows changes in hardness with Co content of the specimens as a function of the LarsonMiller parameter (LMP) for 0.1C-10Cr-0∼5Co-2.6W-0.2Mo-0.2V-0.05Nb-0.05N-0.005B steel normalized and tempered at 600 °C to 820 °C.
Table 135. Chemical composition of SAVE12 steel; [1].
Chemical composition [wt%]
Product C
Si
Mn P
S
Cr Mo W Co V
Nb
2.0 0 0.15 0
Plate 0.07 0.05 0.01 ≤0.015 ≤0.001 8.0 0
∼
∼ ∼ ∼
∼
∼
∼
∼
∼
13.0 0.50 3.5 7.0 0.35 0.10
0.13 0.75 1.00
Ta
0
∼
0.15
Nd
0
∼
0.15
Hf
0
∼
0.15
N
0.01
∼
0.09
B
0
∼
0.0008
450
0％ Co
1％ Co
400
Hardness（HV）
3％ Co
5％ Co
350
Fig. 265. Changes in hardness
with Co content of specimens
as a function of Larson-Miller
Parameter (LMP) for 0.1C10Cr-0~5Co-2.6W-0.2Mo0.2V-0.05Nb-0.05N-0.005B
steels normalized and tempered; [1].
300
250
200
28
30
32
34
36
T (35＋log t) × 10- 3
38
40
Landolt-Börnstein
New Series VIII/2B
Ref. p. 205]
2.3.14 12Cr-3W-3Co-V-Nb-Ta-Nd-N steel
205
2.3.14.3 Creep properties [1]
Fig. 266 shows creep rupture stress vs. time to rupture relations of SAVE12 plates normalized for 0.5 h at
1050 °C and tempered for 1.5 h at 780 °C. It is found that the creep rupture strength increases with
increasing Co content as in the case of softening resistance of the steels, see Fig. 265. The most marked
increase in creep rupture strength with Co content has been obtained for specimens tested at 600 °C. At
higher testing temperatures with low stress levels, the increase in creep rupture strength with Co content
is smaller than at 600 °C.
500
Stress（MPa）
300
200
600℃
100
70
50
0％ Co
Fig. 266.
Creep rupture
properties
of
0.1C-10Cr0~5Co-2.6W-0.2Mo-0.2V0.05Nb-0.05N-0.005B steels
normalized for 0.5 h at
1050 °C and tempered for 1.5
h at 780 °C; [1].
650℃
1％ Co
3％ Co
700℃
5％ Co
101
102
103
104
105
Time to rupture（h）
2.3.14.4 Reference
[1] Igarashi, M., and Sawaragi, Y.: Proc. International Conf. Power Engineeering-97, Vol.2, Tokyo,
(1997), 107.
Landolt-Börnstein
New Series VIII/2B
206
2.4 Austenitic stainless steels
2.4 Austenitic stainless steels
2.4.1 18Cr-8Ni steel
2.4.1.1 Introduction
Austenitic stainless steels, such as type 304 (18Cr-8Ni), 316 (18Cr-12Ni-Mo), 321 (18Cr-10Ni-Ti) and
347 (18Cr-12Ni-Nb) steels, are widely used as high-temperature components, such as boilers and
superheaters, which require good mechanical properties and corrosion resistance at temperatures up to
650 - 700 °C [1-2]. Among the stainless steels, type 304 steel exhibits the most simple chemical
composition and hence the most simple microstructure where M23C6 carbides and σ phase precipitate at
high temperature. The microstructures of other stainless steels are more complicated.
2.4.1.2 Material standards, chemical and tensile requirements
2.4.1.2.1 18Cr-8Ni stainless steel tubes for boilers and heat exchangers
Table 136. Chemical and tensile requirements for 18Cr-8Ni stainless steel tubes for boilers and heat
exchangers; JIS SUS 304 HTB, ASTM TP 304 H, BS 304 S 51.
Chemical composition [wt%]
Standard Designation
Std.
No.
C
Si
Mn
P
S
Ni
Cr
JIS
SUS 304
0.04-0.10 <0.75 <2.00 <0.040 <0.030 8.00-11.00 18.00-20.00 G3463
HTP
ASTM
TP 304 H
0.04-0.10 <0.75 <2.00 <0.040 <0.030 8.0-11.0
18.0-20.0
A213,
A249
BS
304 S 51
0.04-0.10 <1.00 <2.00 <0.040 <0.030 8.00-11.00 17.0-19.0
3059-2
Standard Designation Yield strength
[MPa]
JIS
SUS 304
>205
HTP
ASTM
TP 304 H
>205
>515
BS
490-690
304 S 51
>230
Tensile strength Std. No.
[MPa]
>520
G3463
A213,
A249
3059-2
2.4.1.2.2 18Cr-8Ni stainless steel plates
Table 137. Chemical and tensile requirements for 18Cr-8Ni stainless steel plates; JIS SUS 304-HP,
ASTM S 30400, BS 304 S 31, DIN X5CrNi 1810, ISO L-No.6 X5CrNi 18-9.
Chemical composition [wt%]
StanDesignation
Std. No.
dard
C
Si
Mn
P
S
Ni
Cr
N
JIS
SUS 304-HP <0.08 <1.00 <2.00 <0.045 <0.030 8.0018.00G4304
10.50
20.00
ASTM S 30400
<0.08 <0.75 <2.00 <0.045 <0.030 8.0018.00<0.10 A240,
10.50
20.00
A666
BS
304 S 31
<0.07 <1.00 <2.00 <0.045 <0.030 8.0-11.0 17.0-19.0
1449-2
DIN
C5CrNi1810 <0.07 <1.00 <2.00 <0.045 <0.030 8.5-10.5 17.0-19.0
17441
ISO
L-No6
<0.07 <1.00 <2.00 <0.045 <0.015 8.0017.00X5CrNi 18-9
10.50
19.50
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
Table 137 cont.
Standard Designation
JIS
ASTM
BS
DIN
ISO
SUS 304-HP
S 30400
304 S 31
C5CrNi1810
L-No6 X5CrNi 18-9
Yield
strength
[MPa]
>205
>205
Tensile
strength
[MPa]
>520
>515
207
Std. No.
G4304
A240, A666
1449-2
17441
2.4.1.3 Data sources for 18Cr-8Ni stainless steel
Information of fact on data for 18Cr-8Ni stainless steel can be obtained from [3], [4], [5] and [6].
2.4.1.4 Creep and creep rupture data for 18Cr-8Ni stainless steel for boiler and heat exchanger
seamless steel tubes, JIS SUS 304H TB
2.4.1.4.1 Creep rupture data for 18Cr-8Ni stainless steel tubes, JIS SUS 304H TB
The complete set of creep and creep rupture data, such as creep rupture time, total elongation, reduction
of area, minimum creep rate and optical micrographs of as-received and crept specimens, has been
obtained for the 9 heats in [3]. The details of steel tube production, processing, thermal history, austenite
grain size number, Rockwell hardness and volume fraction of non-metallic inclusions before creep test,
the chemical compositions, the 0.2% proof stress and ultimate tensile strength data at high temperature
are also available for each heat in [3]
Fig. 267 shows the 0.2% proof stress and tensile strength of the 9 heats of 18Cr-8Ni stainless steel, JIS
SUS 304 HTB, [3], obtained by short-time tensile tests between room temperature and 700 °C. The
tensile and creep specimens, having a geometry of 6 mm in diameter and 30 mm in gauge length, were
taken longitudinally from the middle of wall thickness of the tubes.
Fig. 268 shows stress vs. time to rupture data for 9 heats of 18Cr-8Ni steel, JIS SUS 304 HTB, at
temperatures between 600 and 750 °C.
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
100
200
300
400
500
600
Test temperature (℃)
700 800
0
0
100
200
300
400
600
Test temperature (℃)
Fig. 267. Short-time tensile properties of 18Cr-8Ni stainless steel tubes, JIS SUS 304 HTB.
Landolt-Börnstein
New Series VIII/2B
500
700
800
208
2.4 Austenitic stainless steels
300
o
60 0 C
o
62 5 C
o
65 0 C
o
Stress ( MPa )
67 5 C
o
70 0 C
o
72 5 C
100
80
o
75 0 C
o
77 5 C
o
80 0 C
60
o
82 5 C
o
85 0 C
40
n = 28 7
20
1
10
10
2
10
3
10
4
10
5
10
Fig. 268.
Creep rupture
strength data for 18Cr-8Ni
stainless steel tubes, JIS SUS
304 HTB. n indicates the total
number of data points.
6
Time to ru ptu re ( h )
2.4.1.4.2 Estimated long-term creep rupture strength for 18Cr-8Ni stainless steel tubes, JIS SUS
304H TB
The creep rupture data shown in Fig. 268 were analyzed for each heat using the Manson-Haferd
parameter method. Fig. 269 shows the resulting master rupture curve for all data. The estimated stress vs.
time to rupture curves for the heat ABE which shows an intermediate strength level among the 9 heats are
shown in Fig. 270. The 105 h creep rupture strength was also estimated for the 9 heats. This is shown in
Fig. 271 as a function of temperature, together with 0.2% proof stress, ultimate tensile strength and 103 h
creep rupture strength.
400
600 °C
625 °C
650 °C
675 °C
700 °C
725 °C
750 °C
775 °C
800 °C
825 °C
850 °C
300
Stress [MPa]
200
100
80
60
50
40
30
20
Average
n = 289 (309)
-1.0
-4.0
-3.0
-2.5
-1.5
-3.5
-2.0
Manson-Haferd parameter [( log tR -10.378)/( TK - 650 )] [×10 - 2 ]
Fig. 269. Master rupture curve using the MansonHaferd parameter for the 9 heats of 18Cr-8Ni steel tube,
JIS SUS 304 HTB. n indicates the number of data points
used for analysis. In brackets the total number of data
points is given.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
209
30 0
o
Stress ( MPa )
6 00 C
10 0
80
60
o
6 50 C
40
o
700 C
o
75 0 C
20
10
10
2
10
3
10
4
10
5
10
6
Fig. 270.
Estimated
creep rupture strength
curves for the heat ABE
of 18Cr-8Ni steel tube,
JIS SUS 304 HTB.
Ti me to rupture ( h )
500
400
300
Stress [MPa]
Tensile
strength
{
100
80
{
200
0.2%
proof
stress
1000 h
60
50
40
100000 h
30
20
575
600
625
650
700
675
Temperature [°C]
725
750
775
Fig. 271. Estimated 105 h creep
rupture strength for the 9 heats of
18Cr-8Ni steel tube, JIS SUS 304
HTB.
2.4.1.4.3 Creep strain data of 18Cr-8Ni stainless steel tubes, JIS SUS 304H TB
[3] contains the creep strain data for the heat ABE of 18Cr-8Ni steel tubes, JIS SUS 304 HTB. The stress
vs. time to reach 0.5, 1, 2 and 5 % total strain, time to tertiary creep and time to rupture are obtained from
[3]. The relationship between stress and minimum creep rate is shown Fig. 272. The relationship between
time to rupture and minimum creep rate is shown in Fig. 273.
Landolt-Börnstein
New Series VIII/2B
210
2.4 Austenitic stainless steels
300
10 6
10 5
Time to rupture [h]
Stress [MPa]
200
100
80
600 °C
650 °C
700 °C
ABE
n = 20
60
50
40
10-6
10-5 10-4 10-3 10-2 10-1
Minimum creep rate [%/h]
10 3
600 °C
650 °C
700 °C
ABE
n = 15
10 2
30
20
10-7
10 4
1
Fig. 272. Stress vs. minimum creep rate for the heat
ABE of 18Cr-8Ni steel tubes, JIS SUS 304 HTB at
600, 650 and 700 °C. n indicates the total number of
data points.
10
10-7
10-6
10-5 10-4 10-3 10-2 10-1
Minimum creep rate [%/h]
1
Fig. 273. Time to rupture vs. minimum creep rate for the
heat ABE of 18Cr-8Ni steel tube, JIS SUS 304 HTB at
600, 650 and 700 °C. n indicates the total number of data
points.
2.4.1.5 Creep rupture data for 18Cr-8Ni stainless steel plates for reactor vessels, JIS SUS 304-HP
The complete set of creep rupture data, such as creep rupture time, total elongation, reduction of area and
optical micrographs of as-received and crept specimens, has been obtained for the 2 heats in [4]. The
details of steel plate production, processing, thermal history, austenite grain size number, Rockwell
hardness and volume fraction of non-metallic inclusions before creep test, the chemical compositions, the
0.2% proof stress and ultimate tensile strength data at high temperature are also available for each heat in
[4].
Fig. 274 shows the 0.2% proof stress and tensile strength of the 2 heats of 18Cr-8Ni stainless steel in
[4], obtained by short-time tensile tests between room temperature and 850 °C. The tensile and creep
specimens, having a geometry of 10 mm in diameter and 50 mm in gauge length, were taken
longitudinally from the middle of plates.
Fig. 275 shows stress vs. time to rupture data for the 2 heats of 18Cr-8Ni steel plates, JIS SUS 304HP, at temperatures between 450 and 700 °C.
The 105 h creep rupture strength was estimated for the 2 heats. This is shown in Fig. 276 as a function
of temperature, together with 0.2% proof stress, ultimate tensile strength and 103 h creep rupture strength.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
211
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
100
200
300
400
500
600
700
0
800
0
100
Test temperature (℃)
400
500
600
700
800
450 °C
475 °C
500 °C
525 °C
550 °C
575 °C
600 °C
625 °C
650 °C
675 °C
700 °C
400
300
200
Stress [MPa]
300
Test temperature (℃)
Fig. 274. Short-time tensile properties of 18Cr-8Ni steel, JIS SUS 304-HP.
500
200
100
80
60
50
40
30
n = 71
20
10
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 275. Creep rupture strength
data for the 2 heats of 18Cr-8Ni
steel plates, JIS SUS 304-HP.
n indicates the total number of
data points.
800
600
500
400
100
80
60
50
40
30
450
Tensile
strength
{
200
{
Stress [MPa]
300
0.2%
proof
stress
1000 h
Fig. 276. Estimated 105 h creep
rupture strength for the 2 heats of
18Cr-8Ni steel plates, JIS SUS
304-HP.
100000 h
500
Landolt-Börnstein
New Series VIII/2B
550
650
600
Temperature [°C]
700
750
800
212
2.4 Austenitic stainless steels
2.4.1.6 Creep rupture data for welded joints of 18Cr-8Ni stainless steel plates
2.4.1.6.1 Submerged arc welded joints of 18Cr-8Ni stainless steel plates
The complete set of creep rupture data, such as creep rupture time, total elongation, reduction of area and
optical micrographs of as-received and crept specimens for submerged arc welded joints of 18Cr-8Ni
stainless steel plates are given in [4]. The details of base metal plates and the chemical compositions of
the 2 heats AbW and AbX of 18Cr-8Ni stainless steel are also given in [4]. The details of submerged arc
welding procedure, the welding conditions, and the chemical compositions of filer wires and 308 weld
metals can also be obtained from [4].
The tensile and creep specimens, having a geometry of 10 mm in diameter and 100 mm in gauge
length, were taken longitudinally from the middle of plates. The weld metal was located at the center of
specimen gauge, as shown in Fig. 277. Fig. 278 and 279 show the 0.2% proof stress and tensile strength,
obtained by short-time tensile tests for submerged arc welded joints of the heats AbW and AbX,
respectively.
Fig. 280 and 281 show stress vs. time to rupture data for submerged arc welded joints of 18Cr-8Ni
steel plates at temperatures between 500 and 700 °C. The corresponding total elongation and reduction of
area are available in [4].
Fig. 277. Test specimen of
welded joints.
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
0
100
200
300 400 500
600
Test temperature (℃)
700 800
0
100
200
300 400 500
600
700 800
Test temperature (℃)
Fig. 278. Short-time tensile properties of submerged arc welded joints of the heat AbW of 18Cr-8Ni steel.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
213
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
0
100
200
300 400 500
600
700 800
0
100
200
300
400
500
600
700
800
Test temperature (℃)
Test temperature (℃)
Fig. 279. Short-time tensile properties of submerged arc welded joints of the heat AbX of 18Cr-8Ni steel.
500
400
500 °C
550 °C
600 °C
625 °C
650 °C
675 °C
700 °C
300
Stress [MPa]
200
100
80
60
50
40
n = 79
30
10
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 280. Creep rupture data for
submerged arc welded joints of
the heat AbW of 18Cr-8Ni steel.
n indicates the total number of
data points.
500
400
500 °C
550 °C
600 °C
650 °C
700 °C
300
Stress [MPa]
200
100
80
60
50
40
n = 192
30
10
Landolt-Börnstein
New Series VIII/2B
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 281. Creep rupture data for
submerged arc welded joints of
the heat AbX of 18Cr-8Ni steel. n
indicates the total number of data
points.
214
2.4 Austenitic stainless steels
The 105 h creep rupture strength was estimated for submerged arc welded joints, using the data in Fig.
280 and 281. An example of results estimated is shown in Fig. 282 as a function of temperature, together
with 0.2% proof stress, ultimate tensile strength and 103 h creep rupture strength.
800
600
500
400
100 h
100
80
60
50
40
30
450
Tensile
strength
{
200
{
Stress [MPa]
300
0.2%
proof
stress
10000 h
550
500
650
600
Temperature [°C]
750
700
800
Fig. 282. Estimated 105 h creep
rupture strength for the submerged
arc welded joint JAA of the heat
AbW of 18Cr-8Ni steel plates.
2.4.1.6.2 Electron beam welded joints of 18Cr-8Ni stainless steel plates
The complete set of creep rupture data, such as creep rupture time, total elongation, reduction of area and
optical micrographs of as-received and crept specimens for electron beam welded joints of 18Cr-8Ni
stainless steel plates is given in [4]. The details of electron beam welding procedure and welding
conditions for the base metal, heat AbX, are also given in [4].
The tensile and creep specimens, having a geometry of 10 mm in diameter and 100 mm in gauge
length, were taken longitudinally from the middle of plates. The weld metal was located at the center of
specimen gauge, as shown in Fig. 277. Fig. 283 shows the 0.2% proof stress and tensile strength, obtained
by short-time tensile tests for electron beam welded joints of the heat AbX.
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
0
100
200
300 400 500
600
700 800
Test temperature (℃)
0
100
200
300 400 500
600
700 800
Test temperature (℃)
Fig. 283. Short-time tensile properties of electron beam welded joints.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
215
Fig. 284 shows stress vs. time to rupture data for electron beam welded joints of 18Cr-8Ni steel plates at
temperatures between 475 and 700 °C. The corresponding total elongation and reduction of area are
available in [4]. The 105 h creep rupture strength was estimated using the data in Fig. 284. An example is
shown in Fig. 285 for electron beam welded joints, as a function of temperature, together with 0.2% proof
stress, ultimate tensile strength and 103 h creep rupture strength.
500
475 °C
500 °C
525 °C
550 °C
575 °C
600 °C
650 °C
700 °C
400
300
Stress [MPa]
200
100
80
60
50
40
n = 44
30
10
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 284. Creep rupture strength
data for electron beam welded
joints of the heat AbX of 18Cr8Ni steel. n indicates the total
number of data points.
800
600
500
400
Stress [MPa]
300
Tensile
strength
200
0.2%
proof
stress
100
80
60
50
40
30
450
1000 h
100000 h
500
550
600
Temperature [°C]
650
700
750
Fig. 285. Estimated 105 h creep
rupture strength for the electron
beam welded joint JCA of the
heat AbX of 18Cr-8Ni steel
plates.
2.4.1.6.3 Narrow-gap gas tungsten arc welded joints of 18Cr-8Ni stainless steel plates
The complete set of creep rupture data, such as creep rupture time, total elongation, reduction of area and
optical micrographs of as-received and crept specimens for narrow-gap gas tungsten arc welded joints of
18Cr-8Ni stainless steel plates is given in [4]. The details of the narrow-gap gas tungsten arc welding
procedure and welding conditions for the base metal, heat AbX, are also given in [4].
The tensile and creep specimens, having a geometry of 10 mm in diameter and 100 mm in gauge
length, were taken longitudinally from the middle of plates. The weld metal was located at the center of
specimen gauge, as shown in Fig. 277. Fig. 286 shows the 0.2% proof stress and tensile strength obtained
by short-time tensile tests for narrow-gap gas tungsten arc welded joints of the heat AbX.
Landolt-Börnstein
New Series VIII/2B
216
2.4 Austenitic stainless steels
Fig. 287 shows stress vs. time to rupture data for narrow-gap gas tungsten arc welded joints of 18Cr-8Ni
steel plates at temperatures between 475 and 700 °C. The corresponding total elongation and reduction of
area are available in [4]. The 105 h creep rupture strength was estimated using the data in Fig. 287. This is
shown in Fig. 288 for the narrow-gap gas tungsten arc welded joint, as a function of temperature, together
with 0.2% proof stress, ultimate tensile strength and 103 h creep rupture strength.
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
0
100
200
300
400
500
600
0
700
100
Test temperature (℃)
200
300
400
500
600
700
Test temperature (℃)
Fig. 286. Short-time tensile properties of narrow-gap gas tungsten arc welded joints.
500
475 °C
500 °C
525 °C
550 °C
575 °C
600 °C
650 °C
700 °C
400
300
Stress [MPa]
200
100
80
60
50
40
30
10
n = 21
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 287. Creep rupture strength
data for narrow-gap gas tungsten
arc welded joints of the heat AbX
of
18Cr-8Ni
steel
plates.
n indicates the total number of
data points.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
217
800
600
500
400
Tensile
strength
Stress [MPa]
300
200
0.2% proof
stress
100
80
60
50
40
30
450
1000 h
100000 h
500
550
600
Temperature [°C]
650
700
750
Fig. 288. Estimated 105 h creep
rupture strength for the narrow-gap
gas tungsten arc welded joint of the
heat AbX of 18Cr-8Ni steel plates.
2.4.1.7 Microstructure data of 18Cr-8Ni austenitic steel
2.4.1.7.1 Creep fracture modes of 18Cr-8Ni austenitic steel
The microstructure observations by optical, scanning and transmission electron microscopes were carried
out on the longitudinal cross-section of the specimens after creep-rupture. The creep fracture modes were
characterized by the analysis of microstructure near the fracture portion and are shown for the two heats
ABA and ABE in Fig. 289 [7-9]. The heat ABA exhibits a steep decrease in creep rupture strength at long
times, while the heat ABE exhibits an intermediate level of creep rupture strength among the 9 heats
examined in [3]. The creep fracture modes for the two heats ABA and ABE are divided to one
transgranular fracture (denoted by T) and three types of intergranular fracture: the wedge-type cracking
(denoted by W), the creep cavitation associated with M23C6 carbides at grain boundaries (denoted by C )
and the σ/matrix interface cracking along grain boundaries (denoted by σ), as shown in Fig. 289. The
results suggest that the creep fracture modes at long times above 104 h are closely connected with the
precipitation behavior of M23C6 carbides and σ phase.
2.4.1.7.2 Microstructure evolution in 18Cr-8Ni austenitic steel
The microstructure evolution during thermal aging under no stress has been examined by several
researchers [10, 11] for austenitic stainless steels, 304H, 316H, 321H and 347H steels, for up to 5×104 h
at temperatures between 600 (873) and 800 °C (1073 K). But these studies are limited to specimens tested
in periods not exceeding 6×104 h. Recently, the microstructural evolution during creep and during thermal
aging has comprehensively been investigated for JIS SUS 304HTB steel after long-term creep rupture
tests for up to 1.8×105 h, using specimens tested in the NRIM Creep Data Sheet Project [12, 13]. JIS SUS
304HTB steel is one of the most popular austenitic heat resistant steels. A number of micrographs were
recently published for SUS 304HTB steel as ‘Metallographic Atlas of Long-Term Crept Materials’ [14],
parallel with the NRIM Creep Data Sheets.
The Metallographic Atlas contains not only series of micrographs showing the microstructure
evolution during creep for up to 105 h but also the relating data such as time-temperature-precipitation
(TTP) diagrams, histograms describing the distributions of precipitates and creep-voids, and creep
damage parameters. The TTP diagram was constructed for the heat ABE, using the head or grip portion
under no stress of the crept specimens. This is shown in Fig. 290. Only the M23C6 carbides precipitate in
the specimens at short times less than about 104 h at 650 °C, but both M23C6 carbides and σ phase appear
Landolt-Börnstein
New Series VIII/2B
218
2.4 Austenitic stainless steels
at long times above about 104 h. The TEM micrographs clearly show that the M23C6 carbides are observed
in the form of cube-like particles in Widmanstätten distributions in the matrix and in the form of chains of
enlarged particles along grain boundaries, while the σ phase is observed in the form of large and irregular
shapes on grain boundaries and of needles in the matrix.
The mean size of M23C6 carbides in the specimen head portion is plotted in Fig. 291 as a function of
time. In the initial stage of precipitation in the matrix, the size of M23C6 particles is approximately
described by d ∝ t1/2, suggesting diffusion-controlled growth of the M23C6 carbides from supersaturated
solid solution. The slope of the size vs. time curves decreases at long times and the particle size reaches
about 0.05 - 0.07 µm at 105 h at 600 - 700 °C. For the M23C6 carbides at grain boundaries, the slope of the
size-time curves also decreases at long times similar as in the matrix, although the size is much larger
than that in the matrix.
T
Stress [kgf /mm 2 ]
20
10
8
6
700 °C
725 °C
750 °C
775 °C
800 °C
T
C
s
4
T :Transgranular creep fracture
W :Wedge-type cracking
2 C :Cavity formation
s :Cracking at sigma / austenite interface
1
20
10
8
6
s within grain
750
700
650
M23C6 within grain
600
550
s
creep damage test
2
10 2
10
10 3
10 4
Time to rupture [h]
M23C 6 carbide
Grain boundary
10-1
Matrix
1/2
600 °C
650 °C
700 °C
M23C6 on grain boundary
500
10 -1
1
10
10 2
Time [h]
10 3
10 4
10 5
Fig. 290. Time-Temperature-Precipitation (TTP)
diagram for the specimen head portion of the heat
ABE of JIS SUS 304HTB.
10 5
1
Particle size of M23C6 carbide [ m ]
Temperature [°C]
800
}
4
No precipitation
M23C6 on grain boundary
M23C6
M23C6+s on grain boundary
M23C6+s
s on grain boundary
850
W
C
Fig. 289. Stress vs. time to rupture and creep fracture
modes for the heats ABA (left) and ABE (right) of JIS
SUS304HTB.
900
heat B
40
Stress [kgf /mm 2 ]
600 °C
625 °C
650 °C
675 °C
W
heat A
40
10-2
10-1
1
10
10 3
10 2
Time [ h]
10 4
10 5
10 6
Fig. 291. Mean size of M23C6 carbides in the specimen head
portion of crept specimens under no stress as a function of
time. Specimen: heat ABE of JIS SUS 304HTB.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
219
Area fraction [ %]
The size, area fraction and number density of the σ phase on grain boundaries as a function of time are
shown in Fig. 292. Although the σ phase was observed to have irregular shapes, the particle was assumed
to be spherical having equal volume and the diameter of the sphere was regarded as the size of σ phase.
The area fraction, corresponding to the amount of precipitated σ phase, is much larger in the gauge
portion than in the head portion, indicating an acceleration effect of stress and/or strain on the σ phase
precipitation. In the head portion under no stress, the number of σ phase particles significantly increases
with time, while the size increases only slightly. In the gauge portion under stress, on the other hand, the
number density of σ phase particles decreases or is constant with time, while the size significantly
increases with time. These results indicate that the rate-determining process of the precipitation of the σ
phase on grain boundaries is mainly the nucleation at new sites in the head portion under no stress.
However, in the gauge portion under stress, the nucleation is almost completed in the initial stage of
precipitation and the major process of precipitation is the growth. The available nucleation sites are
restricted to grain boundaries perpendicular to stress direction in the gauge portion, while all grain
boundaries are available for the nucleation in the head portion.
10
s phase on grain boundary
600 °C open:specimen head
650 °C solid:specimen gauge
700 °C
1
Number density [mm-2 ]
10 -1
104
8
6
4
2
103
Particle size [mm 2 ]
102
10
1
10 -1
104
Fig. 292. Area fraction, number density and particle size
of σ phase on grain boundaries in the heat ABE, as a
function of time.
2
4
6
8
105
Time [h]
2.4.1.7.3 Creep voids in 18Cr-8Ni austenitic steel
The creep voids were observed to form during creep, which is more significant at lower stress and longer
time conditions. The area fraction and number density of creep voids were measured on the SEM
micrographs, after interruption of the creep tests at several creep strains. The development of creep voids
Landolt-Börnstein
New Series VIII/2B
220
2.4 Austenitic stainless steels
during creep is shown in Fig. 293, as a function of time normalized by time to rupture tr. The creep voids
form at the later stage of creep substantially above t/tr = 0.5 and significantly develop just before creeprupture above t/tr = 0.9. The creep voids were observed to form at the interface between the σ phase on
grain boundaries and austenite matrix, reflecting the σ/matrix interface cracking along grain boundaries
shown in the creep fracture mode diagram in Fig 289.
0.06
Area fraction of creep void [%]
Area fraction
0.05
750 °C, 37 MPa, t r = 26500.7h
700 °C, 53 MPa, t r = 29289.5h
650 °C, 61 MPa, t r = 100491.4h
0.04
0.03
0.02
0.3
0.01
0.4
0.5
0.6
0.7
0.8
Life consumption rate t/tr
0.9
1.0
0
Number density of creep void [mm-2 ]
100
Number density
80
60
40
Fig. 293. A parameter, area fraction and number density
of creep voids in the heat ABE of JIS SUS304HTB, as a
function of t/tr. The A parameter is defined as the
fraction of grain boundaries on which creep voids have
formed.
20
0
0.3
0.4
0.5
0.6
0.7
0.8
Life consumption rate t /tr
0.9
1.0
2.4.1.7.4 Change in hardness of 18Cr-8Ni austenitic steel
Fig. 294 shows the Vickers hardness of the specimen head portion under no stress and of the specimen
gauge portion under stress of the heat ABE, as a function of time. The Vickers hardness in as-received
condition was 160. The specimen head portion under no stress exhibits two step age hardening: shortterm age hardening for times less than 103 h and long-term age hardening for times above 104 h. The
short-term and long-term age hardening results from the precipitation of M23C6 carbides and σ phase,
respectively.
In the gauge portion of creep-ruptured specimens, the hardness decreases with time for all
temperatures, except for 600 °C where it has an increasing tendency at long times above 104 h. The
change in hardness with time for the gauge portion results from the change in dislocation density
produced by creep deformation as well as from precipitation hardening due to M23C6 carbides and σ
phase, as described above. It should be noted that the solid lines in Fig. 294 are connecting the data points
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
221
for the creep-ruptured specimens which were tested at different stress levels as shown in Fig. 289. In
general, resultant dislocation density and resultant dislocation arrangements in the specimens are strongly
influenced by stress level and test duration in the creep rupture testing. This suggests that the solid lines in
Fig. 293 cannot represent the test duration dependence alone but that they also involve the effect of stress
level.
In order to exclude a possible influence of stress level, the change in hardness was measured as a
function of time during creep at a constant load condition [14]. The creep tests were carried out at 650 °C
and at three different stress levels, 177, 118 and 61 MPa at which the time to rupture tr was 71.9, 2621.3
and 100491.4 h, respectively. The creep tests were interrupted at several strains as indicated by the arrows
in Fig. 295 using different specimens and then the hardness was measured. Fig. 296 shows the change in
hardness during creep as functions of time and normalized time t/tr. In this figure, the hardness in the
specimen head portion under no stress is also shown by the dotted line for comparison. The hardening
behavior during creep at 650 °C depends on stress levels as well as on the precipitation of M23C6 carbides
and σ phase. At a high stress of 177 MPa, the hardness increases up to t/tr = 0.7, then decreases slightly
and again increases just before creep-rupture. The hardening during creep is much larger than the age
hardening, indicating that the hardening during creep is mainly caused by strain hardening. Strain
hardening disappears with decreasing stress and increasing test duration. At a low stress of 61 MPa,
hardening during creep is approximately the same as age hardening for almost the whole range of test
duration except for the final stage of creep where softening occurs similar as at 118 MPa. This suggests
that hardening during creep is determined substantially by precipitation hardening due to M23C6 carbides
at short times and due to σ phase at long times above 105 h. For estimating material degradation or
remaining life for JIS SUS 304HTB austenitic steel components, which are usually operated under stress
presumably less than 61 MPa, the hardness can be approximated by age hardening under no stress, except
for the final stage just before creep-rupture.
Fig. 294 and 295, see next page.
a
240
650 °C, 177 MPa, tr = 71.9 h
650 °C, 118 MPa, tr = 2621.3 h
650 °C, 61 MPa, tr = 100491.4 h
650 °C, thermal aging
as-received
220
200
Vickers hardness [HV5 ]
180
160
10 2
0 10
10 3
Time [h]
10 5
10 4
240
b
220
200
180
160
0
Landolt-Börnstein
New Series VIII/2B
0.2
0.4
0.6
Normalized time t /t r
0.8
1.0
Fig. 296. Vickers hardness of the heat ABE of JIS SUS
304HTB during creep, as functions of (a) time and (b)
normalized time t/tr at 650 oC.
222
2.4 Austenitic stainless steels
260
240
220
200
ruptured
650 °C, 177 MPa, t r = 71.9 h
0.40
specimen head
as-received
550 °C
600 °C
650 °C
700 °C
750 °C
0.30
0.20
Vickers hardness [HV5 ]
180
160
0.10
:interrupted times
140
0
120
260
0
0.30
specimen gauge
10
20
30
40
50
70
80
ruptured
650 °C, 118 MPa, t r = 2621.3 h
240
60
0.25
Creep strain
220
200
180
0.20
0.15
160
0.10
140
0.05
120
0
10
10 2
10 3
Time [h]
10 4
10 5
10 6
Fig. 294. Vickers hardness of the specimen head and
gauge portions of the heat ABE of JIS SUS304HTB, as
a function of time.
0
0
500
1000
1500
2000
2500
3000
0.20
650 °C, 61 MPa, tr = 100491.4 h
ruptured
0.15
0.10
→
Fig. 295. Creep curves of the heat ABE of SUS
304HTB at 177, 118 and 61 MPa at 650 °C.
0.05
0
0
20000
40000
60000 80000 100000 120000
Time [h]
2.4.1.7.5 Heat-to-heat variation in creep rupture strength of 18Cr-8Ni austenitic steel
The heat-to-heat variation in long term creep rupture strength has been investigated for JIS SUS 304HTB
steel in order to clarify the correlation between long term creep rupture strength and minor elements [12,
15]. The materials examined were the 9 heats of JIS SUS 304HTB. As given in [3], the concentrations of
impurities, such as Mo, Ti, Al and N, are widely different among the 9 heats. Fig. 297 shows their creep
rupture data at 600 and 700 °C. At a low temperature of 600 °C, the heat-to-heat variation in time to
rupture is very large over the whole stress range examined. The time to rupture of the heat ABL, which is
the strongest of the steels at 600 °C, is about 10 times longer than that of the heat ABD, which is the
weakest one. It should be noted that the heat-to-heat variation is not caused by data scattering, because
each heat exhibits distinct stress dependence of time to rupture. At a high temperature of 700 °C, the heatto-heat variation in time to rupture is not large at high stress and short time conditions less than 104 h but
it becomes more significant again with increasing test duration at low stress and time conditions longer
than 104 h. Austenite grain size and hardness were not different so much among the 9 heats in as-received
condition. The concentrations of major alloying elements C, Si, Mn, Ni and Cr were also not different
among the 9 heats. Therefore, these parameters are excluded as main explanation of the observed heat-toheat variation in time to rupture. It has been well known for ferritic and austenitic heat resistant steels that
nitrogen causes a beneficial effect on the long-term creep rupture strength but that Al causes a
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
223
deteriorative effect. Proposed mechanisms responsible for reducing creep resistance by Al are the
reduction of dissolved nitrogen or fine vanadium nitrides by the formation of large AlN, the refinement of
grain size, the change in distribution of NbN and M23C6, leaving precipitation-free regions around AlN,
and the promotion of grain boundary cracks associated with AlN at grain boundaries [17-20]. Al is a
stronger nitride forming element than Cr, V ad Nb.
200
ABA
ABB
ABC
ABD
ABE
ABF
ABL
ABM
ABN
100
80
60
10
2
10
3
10
4
Tim e to rupture ( h )
10
SUS 304H
o
700 C
100
Stress ( MPa )
Stress ( MPa )
150
SUS 304H
o
600 C
300
5
80
ABA
ABB
ABC
ABD
ABE
ABF
ABL
ABM
ABN
60
40
25
10
2
10
3
10
4
10
5
Time to rupture ( h )
Fig. 297. Creep rupture data for the 9 heats of SUS 304HTB at 600 and 700 °C.
Fig. 298 shows the time to rupture of the 9 heats at 600 °C-137 MPa and 700 °C-47 MPa, as a function of
Al content, where the 105 h-data are included. The content of Al is high only in the heat ABA (0.047 mass
%) while it is low (0.010 to 0.015 mass %) and not substantially different among the other heats. In Fig.
298 there is no distinct relationship between the time to rupture and the content of Al, while the time to
rupture differs widely from about 104 h to about 105 h at approximately the same content of Al of 0.010 to
0.015 mass %. This suggests other factors as well as Al can affect the time to rupture. Titanium is also
known to form large nitrides during the manufacturing process, because it is a strong nitride forming
element.
Fig. 299 shows the time to rupture of the 9 heats at 700 °C, as a function of available nitrogen
concentration which is defined as the concentration of nitrogen free from AlN and TiN and is given by
Nav = Nt − Al – Ti
(1)
where Nt is the total amount of nitrogen in the steel, Al and Ti are the content of Al and Ti in the steel in
at %. The formation of stoichiometric compounds AlN and TiN is assumed. The time to rupture simply
increases with increasing of Nav at long times above 104 h at 700 °C, while the dependence of Nav is split
into two lines having approximately the same slope at short times. The line with longer time to rupture
among the two split lines represents the data for the heats ABA, ABL, ABM and ABN, which contain
more (Nb+Ta) than the other heats. The presented results indicate that precipitation strengthening due to
Nb-carbides is effective at short times. Fig. 300 shows the extremely fine distribution of Nb-carbides in
the heat ABN (6064.8 h) of SUS 304 HTB steel, after creep rupture testing at 650 °C and 137 MPa. Fig.
301 compares stress vs. time to rupture of the heats with different contents of (Nb+Ta) but the same
amounts of Nav at an intermediate temperature of 650 °C. The difference in time to rupture between the
two heats with high and low contents of (Nb+Ta) disappears with increasing test duration, because of
agglomeration of Nb-carbides during creep. The time to rupture is substantially the same between the
heats with the same amount of Nav at about 105 h. At long times, the heat-to-heat variation in time to
rupture is explained by the variation in Nav.
Landolt-Börnstein
New Series VIII/2B
224
2.4 Austenitic stainless steels
5
10 SUSo 304 HTB
105
700 C
o
600 C, 137 MPa
o
700 C, 47 MPa
104
0
0.01
0.02
0.03
0.04
Time to Rupture, t / h
Time to Rupture, t / h
SUS 304 HTB
3
10
47 MPa
69 MPa
98 MPa
102
0.05
0
0.04
0.08
0.12
Available Nitrogen Concentration (at %)
Total Al Content ( mass % )
Fig. 298. Time to rupture of the 9 heats of SUS
304HTB, as a function of Al content.
4
10
Fig. 299. Time to rupture for SUS 304 HTB steel at
700 °C as a function of available nitrogen concentration
defined by Nav = N−Al−Ti, where the concentrations are
in at %.
Fe
b
Intensity I
Cr
Ni
Nb
Fig. 300. (a) TEM micrograph and (b) EDX analysis of
the heat ABN (6064.8 h) of JIS SUS 304 HTB steel,
after creep rupture testing at 650 °C and 137 MPa,
showing the precipitation of extremely fine Nbcarbonitrides.
Nb
V
0
10.240
Energy E [keV]
20.470
Fig. 302 shows schematic drawings of the difference in time to rupture between the heats with high and
low (Nb+Ta) and between the heats with high and low available nitrogen concentrations Nav. The first
increase in heat-to-heat variation with increasing test duration, which is more pronounced at a lower
temperature of 600 °C, is caused by precipitation strengthening due to very fine Nb-carbides having a size
of 10 nm or less. The precipitation strengthening due to Nb-carbides disappears by about 105 h at 650 °C,
because of their agglomeration during creep. This causes the reduction of heat-to-heat variation. The
second increase in heat-to-heat variation at long times is more pronounced at a higher temperature of 700
°C and at long times above 104 h, but it does not appear at 600 °C for up to 105 h. The available nitrogen
concentration, defined as the concentration of nitrogen free from AlN and TiN, clearly explains the
second heat-to-heat variation. Accelerated void formation in the heat ABA containing Al as high as 0.047
mass% also decreases the creep strength at long times.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 225]
2.4.1 18Cr-8Ni steel
225
200
SUS 304 HTB
o
650 C
(a)
Stress, σ / MPa
100
90
N
AV
80
= 0.043 - 0.044 (at %)
heat ABC, 0.36 Mo, 0.01 Nb+Ta
heat ABN, 0.31 Mo, 0.04 Nb+Ta
70
60
200
(b)
100
90
N
AV
80
heat ABD, 0.06 Mo, 0.01 Nb+Ta
70
heat ABM, 0.32 Mo, 0.03 Nb+Ta
60
10
Fig. 301. Stress vs. time to rupture at 650 °C for (a) the
heats ABC and ABN with available nitrogen
concentration of 0.043-0.044 at % and for (b) the heats
ABD and ABM with available nitrogen concentration of
0.092 at %.
= 0.092 (at %)
10
2
3
10
4
5
10
10
Logarithm of time to rupture
Time to Rupture, t / h
Nav : available nitrogen concentration
Nav = N - Al(sol) - Ti (at %)
Middle Nav
High Nb+Ta
High Nav
Middle Nav
Low Nb+Ta
Middle Nav
Low Nav
Low temp.
& short time
Fig. 302. Schematic drawings showing the difference
in time to rupture between the heats with high and low
Nb+Ta and between the heats with high and low Nav.
High temp.
& long time
2.4.1.8 References
[1] Sourmail, T.: Material Science and Technology, 17 (2001), 1-14.
[2] Pickering, F.B.: in ‘Stainless steels 84’, The Institute of Metals, London (1985), pp.2-28.
[3] National Research Institute for Metals (NRIM) Creep Data Sheet, No.4B, (1986) for 18Cr-8Ni
stainless steel, JIS SUS 304H TB.
[4] National Research Institute for Metals (NRIM) Creep Data Sheet, No.32A, (1995) for 18Cr-8Ni
stainless steel, JIS SUS 304-HP.
[5] ASTM Data Series Publication DS 5-S1, (1965)
[6] Elevated temperature properties for steels for pressure purpose, Part I, British Standards Institution
(BSI), (1990)
Landolt-Börnstein
New Series VIII/2B
226
2.4 Austenitic stainless steels
[7] Shinya, N., Kyono, J., Tanaka, H., Murata, M. and Yokoi, S.: Tetsu to Hagane, 69 (1983) 16681675.
[8] Shinya, N., Tanaka, H., Murata, M., Kaise, M. and Yokoi, S.: Tetsu to Hagane, 71 (1985) 114-120.
[9] Tanaka, H., Murata, M., Kaise, M. and Shinya, N.: Tetsu to Hagane, 74 (1988) 2009-2016.
[10] Biss, V. A. and Sikka, V. K.: Metall. Trans. 12A (1981), 1360-1362.
[11] Minami, Y., Kimura, H. and Ihara, Y.: Mater. Sci. Technol. 2 (1986), 795-806.
[12] Murata, M., Tanaka, H., Abe, F. and Irie, H.: Key Engineering Materials, Trans Tech Publications,
Switzerland, 171-174 (2000), 513-520.
[13] Abe, F., Tanaka, H., Murata, H., Irie, H. and Yagi, K.: Proc. 4th Japan-China Bilateral Symposium on
High Temperature Strength of Materials, Tsukuba, Japan, (2001), 83-88.
[14] National Research Institute for Metals Creep Data Sheet, Metallographic Atlas of Long-Term Crept
Materials, Nat. Res. Inst. for Metals, Tsukuba, No.M-1 (1999).
[15] Tanaka, H., Murata, M., Abe, F. and Irie, H.: Materials Science and Engineering, A319 (2001) 788791.
[16] Miyazaki, H., Tanaka, H., Murata, M. and Abe, F.: J. Japan Institute of Metals, 66 (2002), 12781286.
[17] Yukitoshi. T. and Nishida, K.: Trans. ISIJ, 12 (1972) 429-434.
[18] Shinya, N., Kyono, J., Tanaka, H., Murata, M. and Yokoi, S.: Tetsu-To-Hagane, 69 (1983) 16681675.
[19] Schirra, M. and Anderko, K.: Steel Research, 61 (1990) 242-250.
[20] Naoi, H., Ohgami, M., Liu, X. and Fujita, T.: Metall. Trans., 28A (1997) 1195-1203.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
227
2.4.2 18Cr-12Ni-Mo steel
2.4.2.1 Introduction
The molybdenum-containing AISI type 316 (SUS316HTB, SUS316-HP, SUS316-B, X2CrNiMo17-12-2,
X5CrNiMo17-12-2, X6CrNiMoTi17-12-2, X2CrNiMoN17-13-3, X2CrNiMo17-12-2, SS 2343,
X3CrNiMoBN17-13-3) stainless steel has been used within the power-generating industry. A common
usage is as superheater tubing exposed to high temperatures of 650 °C or higher [1]. It has higher creep
strength than the unstabilized molybdenum-free AISI type 304 steel and better resistance to heat-affected
zone cracking during welding than the niobium- and titanium-stabilized grades, i.e. AISI types 347 and
321. However, in some circumstances it can become embrittled after prolonged exposure at elevated
temperatures as a result of the formation of carbide and intermetallic phases. Moreover, it is now known
that large cast-to-cast variations exist in this grade of steel. Its long-term ductility can vary from below 10
% to over 100 %. The demand of high reliability in modern plants, especially nuclear, requires a
substantial reduction in material variability.
2.4.2.2 Materials standards, and chemical and tensile requirements
2.4.2.2.1 18Cr-12Ni-Mo stainless steel tubes for boilers and heat exchangers
Table 138. Chemical and tensile requirements for 18Cr-12 Ni-Mo stainless
heat exchangers; JIS SUS 316 HTB, ASTM TP 316 H, BS 316 S 33.
Chemical composition [wt%]
StanDesignation
dard
C
Si
Mn
P
S
Ni
JIS
SUS 316
0.04<0.75 <2.00 <0.040 <0.030 11.00HTB
0.10
14.00
ASTM TP 316 H
0.04<0.75 <2.00 <0.040 <0.030 11.00.10
14.0
ASTM TP 316 H
0.04<0.75 <2.00 <0.040 <0.030 11.00.10
14.0
BS
316 S 33
<0.07 1.00
<2.00 <0.040 <0.030 11.014.0
Standard
JIS
Designation Yield strength
[MPa]
SUS 316
>205
HTB
ASTM TP 316 H
>205
ASTM TP 316 H
>205
BS
316 S 33
>245
Landolt-Börnstein
New Series VIII/2B
Tensile
Std.
strength [MPa] No.
>520
G3463
>515
>515
410-710
A213
A249
3606
steel tubes for boilers and
Cr
16.0018.00
16.018.0
16.018.0
16.5018.50
Mo
2.003.00
2.003.00
2.003.00
2.503.00
Std.
No.
G3463
A213
A249
3606
228
2.4 Austenitic stainless steels
2.4.2.2.2 18Cr-12Ni-Mo stainless steel plates
Table 139. Chemical and tensile requirements for 18Cr-8 Ni stainless steel plates; JIS SUS
ASTM S 31600, BS 316 S 31, DIN X5CrNi 17122, ISO L-No.26 X5CrNiMo 17-22-2.
Chemical composition [wt%]
StanDesignation
dard
C
Si
Mn
P
S
Ni
Cr
Mo
N
JIS
SUS 316<0.08 <1.00 <2.00 <0.045 <0.030 11.00- 16.00- 2.00HP
14.00 18.00 3.00
ASTM S 31600
0.04- <0.75 <2.00 <0.040 <0.030 11.00- 16.0- 2.00- <0.1
0.10
14.00 18.0
3.00
BS
316 S 31
0.04- <0.75 <2.00 <0.040 <0.030 10.5- 16.0- 2.00.10
13.5
18.0
2.5
DIN
X5CrNiMo <0.07 <1.00 <2.00 <0.040 <0.030 10.5- 16.50- 2.017122
13.5
18.50 2.5
ISO
L-No26
<0.07 <1.00 <2.00 <0.045 <0.030 10.0- 16.0- 2.0<0.11
X5CrNiMo
13.0
18.0
3.0
17-12-2
Standard
JIS
ASTM
BS
DIN
ISO
Designation
SUS 316-HP
S 31600
316 S 31
X5CrNiMo 17122
L-No26 X5CrNiMo 17-12-2
Yield strength
[MPa]
>206
>205
Tensile strength
[MPa]
>520
>515
316-HP,
Std.
No.
G4304
A240,
A666
1449-2
17441
Std. No.
G4304
A240, A666
1449-2
17441
2.4.2.2.3 18Cr-12Ni-Mo stainless steel bars
Table 140. Chemical and tensile requirements for 18Cr-8 Ni stainless steel bars; JIS SUS 316-B, ASTM
S 31600, BS 316 S 31, DIN X5CrNi 17122, ISO L-No.26 X5CrNiMo 17-22-2.
Chemical composition [wt%]
StanDesignation
Std.
dard
No.
C
Si
Mn
P
S
Ni
Cr
Mo
N
JIS
SUS 316-B <0.08 <1.00 <2.00 <0.045 <0.030 11.00- 16.00- 2.00G4303
14.00 18.00 3.00
ASTM S 31600
0.04- <0.75 <2.00 <0.040 <0.030 11.00- 16.0- 2.00- <0.1 A276,
0.10
14.00 18.0
3.00
A493
BS
316 S 31
0.04- <0.75 <2.00 <0.040 <0.030 10.5- 16.0- 2.00.10
13.5
18.0
2.5
DIN
X5CrNiMo <0.07 <1.00 <2.00 <0.040 <0.030 10.5- 16.50- 2.01654-5
17122
13.5
18.50 2.5
ISO
L-No26
<0.07 <1.00 <2.00 <0.045 <0.030 10.0- 16.0- 2.0<0.11
X5CrNiMo
13.0
18.0
3.0
17-12-2
Standard
JIS
ASTM
BS
DIN
ISO
Designation
SUS 316-HP
S 31600
316 S 31
X5CrNiMo 17122
L-No26 X5CrNiMo 17-12-2
Yield strength
[MPa]
>206
>205
Tensile strength
[MPa]
>520
>515
Std. No.
G4304
A240, A666
1449-2
17441
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
229
2.4.2.3 Data sources for 18Cr-12Ni-Mo stainless steel
Information of fact on data for 18Cr-12Ni-Mo stainless steel can be obtained from [2], [3], [5] and [6].
2.4.2.4 Creep rupture data for 18Cr-12Ni-Mo stainless steel for boiler and heat exchanger seamless
steel tubes, JIS SUS 316H TB
2.4.2.4.1 Creep and creep rupture data for 18Cr-12Ni-Mo stainless steel tubes, JIS SUS 316H TB
The complete set of creep and creep rupture data, such as creep rupture time, total elongation, reduction
of area, minimum creep rate and optical micrographs of as-received and crept specimens, has been
obtained for 9 heats in [2]. The details of steel tube production, processing, thermal history, austenite
grain size number, Rockwell hardness and volume fraction of non-metallic inclusions before creep test,
the chemical compositions, the 0.2% proof stress and ultimate tensile strength data at high temperature
are also available for each heat in [2].
The tensile and creep specimens used in [2], having a geometry of 6 mm in diameter and 30 mm in
gauge length, were taken longitudinally from the middle of wall thickness of the tubes. Fig. 303 shows
the 0.2% proof stress and tensile strength obtained by short-time tensile tests between room temperature
and 750 °C.
Fig. 304 (next page) shows the stress vs. time to rupture data for the 9 heats of 18Cr-12Ni-Mo
stainless steel tubes, JIS SUS 316 HTB, at temperatures between 600 and 750 °C. The fact data are
available in [2].
Tensile strength
0.2% proof stress
700
700
600
600
400
Stress (MPa)
Stress (MPa)
500
300
500
400
200
300
100
0
0
100
200
300
400 500
600
Test temperature (℃)
700
800
200
0
100 200 300
400
500 600 700 800
Test temperature (℃)
Fig. 303. Short-time tensile properties of 18Cr-12Ni-Mo stainless steel tubes, JIS SUS 316 HTB.
2.4.2.4.2 Estimated long-term creep rupture strength for 18Cr-12Ni-Mo stainless steel tubes, JIS
SUS 316H TB
The creep rupture data shown in Fig. 304 were analyzed for each heat using the Manson-Haferd
parameter method. Fig. 305 shows the estimated 105 h creep rupture strength for the 9 heats as a function
of temperature. In this figure, the 0.2% proof stress, ultimate tensile strength and 103 h creep rupture
strength are also included.
Landolt-Börnstein
New Series VIII/2B
230
2.4 Austenitic stainless steels
500
400
300
600 °C
650 °C
700 °C
750 °C
Stress [MPa]
200
100
80
60
50
40
30
20
n = 319
10 3
10 4
Time to rupture [h]
10 5
800
600
500
400
300
Stress [MPa]
200
100
80
60
50
40
30
0.2% proof
stress
10 6
1000 h
20
10
550
Tensile
strength
{
10 2
{
10
10
Fig. 304. Creep rupture strength
data for 18Cr-12Ni-Mo stainless
steel tubes, JIS SUS 316 HTB. n
indicates the total number of data
points.
100000 h
600
700
650
Temperature [°C]
750
800
Fig. 305. Estimated 105 h creep
rupture strength for the 9 heats of
18Cr-12Ni-Mo stainless steel tube,
JIS SUS 316 HTB.
2.4.2.4.3 Creep strain data of 18Cr-12Ni-Mo stainless steel tubes, JIS SUS 316H TB
[2] contains the creep strain data for the heat AAL of 18Cr-12Ni-Mo stainless steel tubes, JIS SUS 316
HTB. The stress vs. time to reach 0.5, 1, 2 and 5 % total strain, time to tertiary creep and time to rupture
are obtained from [2]. The total strain is defined as instantaneous strain and creep strain, which are
produced at loading and during creep, respectively. The relationship between stress and minimum creep
rate is shown in Fig. 306. The relationship between time to rupture and minimum creep rate is shown in
Fig. 307.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
231
Fig. 306. Stress vs. minimum creep rate for
the heat AAL of 18Cr-12Ni-Mo stainless steel
tubes, JIS SUS 316 HTB at 600, 650, 700 and
750 °C. n indicates the total number of data
points.
Fig. 307. Time to rupture vs. minimum creep
rate for the heat AAL of 18Cr-12Ni-Mo
stainless steel tubes, JIS SUS 316 HTB at 600,
650, 700 and 750 °C. n indicates the total
number of data points.
2.4.2.5 Creep rupture data for 18Cr-12Ni-Mo stainless steel plates for reactor vessels, JIS SUS 316HP
2.4.2.5.1 Creep rupture data for 18Cr-12Ni-Mo stainless steel plates, JIS SUS 316-HP
The complete set of creep and creep rupture data, such as creep rupture time, total elongation, reduction
of area, minimum creep rate and optical micrographs of as-received and crept specimens, has been
obtained for 2 heats in [3]. The details of steel plate production, processing, thermal history, austenite
grain size number, Rockwell hardness and volume fraction of non-metallic inclusions before creep test,
the chemical compositions, the 0.2% proof stress and ultimate tensile strength data at high temperature
are also available for each heat in [3].
The details of plate production and the chemical compositions, respectively, of the 2 heats of 18Cr12Ni-Mo stainless steel plates, JIS SUS316-HP, are given in [3]. The tensile and creep specimens, having
a geometry of 10 mm in diameter and 50 mm in gauge length, were taken longitudinally from the middle
of thickness of the plates. Fig. 308 shows the 0.2% proof stress and tensile strength obtained by shorttime tensile tests between room temperature and 850 °C.
Fig. 309 shows the stress vs. time to rupture data for the 2 heats of 18Cr-12Ni-Mo stainless steel
plates, JIS SUS 316-HP, at temperatures between 600 and 850 °C. The fact data are available in [3].
Landolt-Börnstein
New Series VIII/2B
232
2.4 Austenitic stainless steels
Tensile strength
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
600
300
200
200
100
100
0
300
0
200
400
600
800
0
1000
0
200
Test temperature (℃)
400
600
800
1000
Test temperature (℃)
Fig. 308. Short-time tensile properties of 18Cr-12Ni-Mo stainless steel plates, JIS SUS 316-HP.
50 0
o
6 00 C
30 0
o
6 50 C
o
Stress ( MPa )
700 0 C
o
7 50 C
o
8 00 C
10 0
80
o
8 50 C
60
40
20
10 1
n = 57
10 2
1 03
10 4
10 5
10 6
Fig. 309.
Creep rupture
strength data for 18Cr-12NiMo stainless steel plates, JIS
SUS 316-HP. n indicates the
total number of data points.
T ime to rupture ( h )
2.4.2.5.2 Estimated long-term creep rupture strength for 18Cr-12Ni-Mo stainless steel plates, JIS
SUS 316-HP
The creep rupture data shown in Fig. 309 were analyzed for each heat using the Manson-Haferd
parameter method. Fig. 310 shows the resulting master rupture curve for all data. The 105 h creep rupture
strength was also estimated for the 2 heats. This is shown in Fig. 311 as a function of temperature,
together with 0.2% proof stress, ultimate tensile strength and 103 h creep rupture strength.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
400
600 °C
625 °C
650 °C
675 °C
700 °C
750 °C
775 °C
800 °C
850 °C
300
200
Stress [MPa]
233
100
80
60
50
40
30
Fig. 310. Master rupture curve based on the MansonHaferd parameter method for the 2 heats of 18Cr-12NiMo stainless steel plates, JIS SUS 316-HP. n indicates
the total number of data points.
20
Average
n = 60
10
-4.0
-1.0
-5.0
-3.0
-2.0
Manson-Haferd parameter [( log tR -9.223)/( TK - 720 )] [×10 - 2 ]
500
400
300
200
{
Stress [MPa]
{
100
80
60
50
40
1000 h
30
20
500
Tensile
strength
0.2%
proof
stress
100000 h
550
600
750
700
650
Temperature [°C]
800
850
900
Fig. 311. Estimated 105 h creep
rupture strength for the 2 heats of
18Cr-12Ni-Mo stainless steel
plates, JIS SUS 316-HP.
2.4.2.5.3 Creep strain data of 18Cr-12Ni-Mo stainless steel plates, JIS SUS 316-HP
[3] contains creep strain data for 2 heats of 18Cr-12Ni-Mo stainless steel plates, JIS SUS 316-HP. The
stress vs. time to reach 0.5, 1, 2 and 5 % total strain, time to tertiary creep and time to rupture for the AaA
heat can also be obtained from [3]. The total strain is defined as instantaneous strain and creep strain,
which are produced at loading and during creep, respectively. The relationship between stress and
minimum creep rate is shown Fig. 312. The relationship between time to rupture and minimum creep rate
is shown in Fig. 303.
Landolt-Börnstein
New Series VIII/2B
234
2.4 Austenitic stainless steels
300
Stress [MPa]
200
100
80
600 °C
60
50
650 °C
40
30
700 °C
20
750 °C
800 °C 850 °C
n = 28
10
8
10 -7
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
600 °C
625 °C
650 °C
675 °C
700 °C
750°C
775 °C
800 °C
850 °C
10 -1
Fig. 312. Stress vs. minimum creep rate for the heats
AaA and AaB of 18Cr-12Ni-Mo stainless steel plates, JIS
SUS 316-HP at 600, 650, 700, 750, 800 and 850 °C.
n indicates the total number of data points.
1
10 6
600 °C
625 °C
650 °C
675 °C
700 °C
750°C
775 °C
800 °C
850 °C
Time to rupture [h]
10 5
10 4
10 3
Fig. 313. Time to rupture vs. minimum creep rate for the
heats AaA and AaB of 18Cr-12Ni-Mo stainless steel
plates, JIS SUS 316-HP at at 600, 650, 700, 750, 800 and
850 °C. n indicates the total number of data points.
10 2
n = 28
10
1
10 -7
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
10 -1
1
2.4.2.6 Creep rupture data for 18Cr-12Ni-Mo-middle nitrogen-low carbon stainless steel plates, JIS
SUS 316-HP
The creep rupture data, such as creep rupture time, total elongation and reduction of area, and optical
micrographs of as-received specimens, have been obtained for one heat in [4]. The details of steel plate
production, processing, thermal history, austenite grain size number, Rockwell hardness and volume
fraction of non-metallic inclusions before creep test, the chemical composition, the 0.2% proof stress and
ultimate tensile strength data at high temperature are also available in [4].
Fig. 314 shows the 0.2% proof stress and tensile strength of the 1 heat of 18Cr-12Ni-Mo-middle
nitrogen-low carbon stainless steel plates, JIS SUS316-HP , obtained by short-time tensile tests between
room temperature and 800 °C; [4]. The tensile and creep specimens, having a geometry of 10 mm in
diameter and 50 mm in gauge length, were taken longitudinally from the middle of plates.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
235
Tensile strength
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
600
300
200
300
200
100
100
0
0
0
200
400
600
800
1000
0
200
400
600
800
1000
Test temperature (℃)
Test temperature (℃)
Fig. 314. Short-time tensile properties of 18Cr-12Ni-Mo-middle nitrogen-low carbon stainless steel plates, JIS SUS
316-HP.
Fig. 315 shows the stress vs. time to rupture data for 18Cr-12Ni-Mo-middle nitrogen-low carbon stainless
steel plates, JIS SUS 316-HP, at temperatures between 500 and 750 °C. The fact data are available in [4].
Stress ( MPa)
1000
100
o
50 0 C
o
55 0 C
Fig. 315.
Creep rupture
strength data for 18Cr-12NiMo-middle nitrogen-low carbon stainless steel plate, JIS
SUS 316-HP.
o
60 0 C
o
65 0 C
70 0oC
o
10
10
75 0 C
2
3
4
10
10
10
Time to rupture ( h )
5
10
2.4.2.7 Creep rupture data for 18Cr-12Ni-Mo stainless steel bars, JIS SUS 316-B
2.4.2.7.1 Creep rupture data
The complete set of creep and creep rupture data, such as creep rupture time, total elongation, reduction
of area, minimum creep rate and optical micrographs of as-received and crept specimens, has been
obtained for the 6 heats in [5]. The details of steel bar production, processing, thermal history, austenite
Landolt-Börnstein
New Series VIII/2B
236
2.4 Austenitic stainless steels
grain size number, Rockwell hardness and volume fraction of non-metallic inclusions before creep test,
the chemical compositions, the 0.2% proof stress and ultimate tensile strength data at high temperature
are also available for each heat in [5].
Fig. 316 shows the 0.2% proof stress and tensile strength of the 6 heats of 18Cr-12Ni-Mo stainless
steel bars, JIS SUS316-B obtained by short-time tensile tests between room temperature and 850 °C; [5].
The tensile and creep specimens, having a geometry of 10 mm in diameter and 50 mm in gauge length,
were taken longitudinally from the middle of bars.
Fig. 317 shows the stress vs. time to rupture data for the 6 heats of 18Cr-12Ni-Mo stainless steel bars,
JIS SUS 316-B, at temperatures between 600 and 850 °C. The fact data are available in [5].
Tensile strength
700
600
600
500
500
400
400
Stress (MPa)
Stress (MPa)
0.2% proof stress
700
300
300
200
200
100
100
0
0
200
400
600
800
0
1000
0
200
Test temperature (℃)
400
600
800
1000
Test temperature (℃)
Fig. 316. Short-time tensile properties of 18Cr-12Ni-Mo-middle nitrogen-low carbon stainless steel plates, JIS SUS
316-HP.
50 0
o
60 0 C
o
30 0
65 0 C
o
Stress ( MPa )
70 0 C
o
75 0 C
o
80 0 C
o
10 0
80
85 0 C
60
Fig. 317.
Creep rupture
strength data for 18Cr-12NiMo stainless steel bars, JIS
SUS 316-B. n indicates the
total number of data points.
40
n = 1 66
20
10
1
10
2
3
10
10
4
10
5
T ime to rupture ( h )
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
237
2.4.2.7.2 Estimated long-term creep rupture strength for 18Cr-12Ni-Mo stainless steel bars, JIS
SUS 316-B
The creep rupture data shown in Fig. 317 were analyzed for each heat using the Manson-Haferd
parameter method. Fig. 318 shows the resulting master rupture curve for all data. The 105 h creep rupture
strength was also estimated for the 6 heats. This is shown in Fig. 319 as a function of temperature,
together with 0.2% proof stress, ultimate tensile strength and 103 h creep rupture strength.
400
600 °C
650 °C
700 °C
750 °C
800 °C
850 °C
300
Stress [MPa]
200
100
80
60
50
40
30
Fig. 318. Master rupture curve based on the MansonHaferd parameter method for the 6 heats of 18Cr-12NiMo stainless steel bars, JIS SUS 316-B. n indicates the
total number of data points.
20
Average
n = 166
10
-1.0
-4.0
-2.0
-5.0
-3.0
Manson-Haferd parameter [( log tR -8.447)/( TK - 720 )] [×10 - 2 ]
500
400
300
200
{
Stress [MPa]
{
100
80
60
50
40
30
1000 h
20
10
500
Tensile
strength
0.2%
proof
stress
100000 h
550
600
650
750
700
Temperature [°C]
800
850
900
Fig. 319. Estimated 105 h creep
rupture strength for the 9 heats of
18Cr-12Ni-Mo stainless steel bars,
JIS SUS 316-B.
2.4.2.7.3 Creep strain data of 18Cr-12Ni-Mo stainless steel bars, JIS SUS 316-B
[5] contains creep strain data for 6 heats of 18Cr-12Ni-Mo stainless steel bars, JIS SUS 316-B. The stress
vs. time to reach 0.5, 1, 2 and 5 % total strain, time to tertiary creep and time to rupture are obtained from
[5]. The relationship between stress and minimum creep rate is shown in Fig. 320. The relationship
between time to rupture and minimum creep rate is shown in Fig. 321.
Landolt-Börnstein
New Series VIII/2B
238
2.4 Austenitic stainless steels
300
200
Stress [MPa]
600 °C
100
80
60
50
40
700 °C
650 °C
30
20
750 °C
800 °C 850 °C
600 °C
650 °C
700 °C
750 °C
800 °C
850 °C
Fig. 320. Stress vs. minimum creep rate for the 6 heats of
18Cr-12Ni-Mo stainless steel bars, JIS SUS 316-B at 600,
650, 700, 750, 800 and 850 °C. n indicates the total
number of data points.
n = 68
10
10 -7
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
10 6
1
600 °C
650 °C
700 °C
750 °C
800 °C
850 °C
10 5
Time to rupture [h]
10 -1
10 4
10 3
10 2
n = 68
10 1
10 -7
10 -6
10 -4
10 -5
10 -3
Minimum creep rate [%/h]
10 -2
10 -1
Fig. 321. Time to rupture vs. minimum creep rate for the
6 heats of 18Cr-12Ni-Mo stainless steel bars, JIS SUS
316-B at 600, 650, 700, 750, 800 and 850 °C. n indicates
the total number of data points.
2.4.2.8 Microstructure data of 18Cr-12Ni-Mo stainless steel
2.4.2.8.1 Creep fracture modes of JIS SUS316HTB steel
The creep fracture modes were characterized for the two heats AAF and AAL of JIS SUS316HTB steel
[2] by analysis of the microstructure near the fracture portion [7-9]. This is shown in Fig. 322. The heat
AAF contained a high Al concentration (0.095 wt% Al), while the heat AAF contained a low Al
concentration (0.017 wt% Al). The microstructure observations by optical, scanning and transmission
electron microscopes were carried out on the longitudinal cross-section of the specimens after creeprupture. The creep fracture modes are divided to transgranular fracture (denoted by T) and three types of
intergranular fracture: the wedge-type cracking at triple point (denoted by W), the intergranular cavitation
associated with M23C6 carbides at grain boundaries (denoted by C ) and the σ/matrix interface cracking
along grain boundaries (denoted by σ), as shown in Fig. 322. The results suggest that the creep fracture
modes at long times above 104 h are closely connected with the precipitation behavior of M23C6 carbides
and σ phase. The precipitation of M23C6 carbides at grain boundaries promotes the transition from wedgetype cracking to intergranular cavitation, and the precipitation of σ phase at grain boundaries leads to the
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
239
σ/matrix interface cracking. Al contents above 0.03 wt% lead to the precipitation of AlN associated with
the grain boundary σ phase. The precipitation of AlN accelerates the σ/matrix interface cracking,
resulting in the reduction of creep rupture strength and ductility at long times.
30
20
600 °C
700 °
C
10
725
775
800
Stress [kgf / mm2 ]
5
Heat F
625 °C
650 °
C
675 °
C
°C
°C
75
0°
°C
C
3
30
600 °C
Heat L
650 °C
20
700 °
10
625 °C
675 °
C
725 °
C
C
750 °
C
5
3
Transgranular creep fracture
Wedge-type cracking
Cavity formation
Cracking at s/matrix interface
Cracking at c/matrix interface
10 2
775 °
C
10 3
10 4
Time to rupture [h]
10 5
Fig. 322. Stress vs. time to rupture and creep fracture
modes for the heats AAF (Heat F) and AAL (Heat L) of
JIS SUS316HTB steel.
2.4.2.8.2 Microstructure evolution in JIS SUS3164HTB steel
The microstructure evolution during thermal aging under no stress and during creep has been examined
by several researchers [10-16] for type 316 stainless steel at temperatures between 600 (873) and 800 °C
(1073 K). But these studies are limited to specimens tested in periods not exceeding 6×104 h. Recently,
the microstructure evolution during creep and during thermal aging has been comprehensively
investigated for SUS 316HTB steel after long-term creep rupture tests over 105 h, using specimens tested
in the NIMS Creep Data Sheet Project [8-9]. A number of micrographs were recently published for SUS
316HTB steel as ‘Metallographic Atlas of Long-Term Crept Materials’ [8], parallel with the NIMS Creep
Data Sheets.
The Metallographic Atlas of Long-Term Crept Materials contains not only series of micrographs
showing the microstructure evolution during creep for up to 105 h but also the relating data such as timetemperature-precipitation (TTP) diagrams, histograms describing the distributions of precipitates and
creep-voids, and creep damage parameters. The TTP diagram was constructed for the heat AAL, using
the head or grip portion of the crept specimens under no stress. This is shown in Fig. 323. Only the M23C6
carbides precipitate in the specimens at short times less than about 103 h at 650 °C but Laves and σ
phases also appear at long times above about 103 and 104 h, respectively. The χ phase appears at long
times in the temperature range between 750 and 900 °C. The TEM micrographs clearly show that the
M23C6 carbides are observed in form of cube-like particles in Widmanstätten distributions in the matrix
and in form of chains of enlarged particles along grain boundaries, as shown in Fig. 324. The Laves phase
is observed in rod shape in the matrix, while the σ and χ phases are in form of large and irregular shapes
on grain boundaries.
Landolt-Börnstein
New Series VIII/2B
240
2.4 Austenitic stainless steels
Fig. 323.
Time-Temperature-Precipitation
(TTP) diagram for the specimen head portion
of the heat AAL of JIS SUS 316HTB steel.
Fig. 324. TEM micrographs of the heat AAL of JIS SUS 316HTB steel after aging under no stress.
The change in particle size of M23C6 carbides, Laves phase and σ phase is shown as a function of time for
the heat AAL of JIS SUS316HTB steel in Fig. 325 and Fig. 326. A series of optical, scanning and
transmission electron micrographs are included in the Metallographic Atlas of Long-Term Crept
Materials [8]. The area fraction, corresponding to the amount of precipitated σ phase, is much larger in
the gauge portion under stress than in the head portion under no stress, indicating an acceleration effect of
stress and/or strain on the σ phase precipitation, as shown in Fig. 326.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
241
The concentrations of the major components Cr, Fe, Mo and Ni in the M of M23C6 carbides were
measured on extracted replicas by EDX in SEM. The summation of metallic elements is assumed to be
100 %. The results are shown in Fig. 327 as a function of time. At 650 °C, the concentration of Cr
increases from about 50 - 55 % at 102 h to about 60 % at 105 h in the specimen head, while that of Fe
decreases from 25 % at 102 h to 15 % at 105 h. Because the equilibrium concentration of Cr in the M23C6
carbides is much higher than the concentration of Cr in the matrix, it is considered that at first the M23C6
carbides precipitate with a Cr concentration much lower than the equilibrium one and that the
concentration of Cr in the M23C6 carbides gradually increases during exposure at high temperatures
toward the equilibrium one. The change in the concentrations of Cr and Fe with time shifts to shorter
times with increasing temperature, suggesting a diffusion-controlled process. The change in the
concentrations of Cr and Fe with time was accelerated under stress.
In the Laves phase and σ phase, the concentrations of Fe, Cr, Mo and Ni do not change so much with
time, as shown in Fig. 328 and Fig. 329. There was also no difference in concentrations between the
specimen head and gauge portions. The present results indicate that the Laves phase and σ phase, having
the equilibrium concentrations of Fe, Cr, Mo and Ni, precipitate from the initial stage. This is quite
different from the remarkable change in concentrations for the M23C6 carbides during high temperature
exposure.
Fig. 325. Change in particle size of (a) M23C6
carbides and (b) Laves phase within grains in
specimen head portion as a function of time
for the heat AAL of JIS SUS316HTB steel.
Landolt-Börnstein
New Series VIII/2B
242
2.4 Austenitic stainless steels
Fig. 326. Change in area fraction, number density and
particle size of σ phase on grain boundaries in specimen
head and gauge portions as a function of time for the
heat AAL of JIS SUS316HTB steel.
2.4.2.8.3 Change in hardness in JIS SUS3164HTB steel
Fig. 330 shows the Vickers hardness of the specimen head portion under no stress and the specimen
gauge portion under stress of the heat AAL of JIS SUS316HTB steel, as a function of time. The Vickers
hardness in as-received condition was 160. The specimen head portion under no stress exhibits hardening
with time for up to 105 h or more, indicating precipitation hardening, although detailed relationship
between the hardening and precipitation behavior is not clear. Presumably, the short-term and long-term
age hardening results from the precipitation of M23C6 carbides and intermetallic compounds, respectively.
The gauge portion of specimens, creep-ruptured at short times less than 103 h, exhibits hardening
compared with the as-received condition and the hardness decreases with time for up to 105 h or more.
The change in hardness with time for the gauge portion results from the change in dislocation density
produced by creep deformation as well as from the precipitation hardening described above. It should be
noted that the solid lines in Fig. 330 are connecting the data points for the creep-ruptured specimens
which were tested at different stress levels as shown in Fig. 322. In general, resultant dislocation density
and resultant dislocation arrangements in the specimens are strongly influenced by stress level and test
duration in the creep rupture testing. This suggests that the solid lines in Fig. 330 cannot represent the test
duration dependence alone but that they also involve the effect of stress level.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
80
2.4.2 18Cr-12Ni-Mo steel
80
550 °C
60
40
20
Chemical composition of M in M23 C 6 carbide [ %]
600 °C
60
Fe
Cr
Mo
Ni
40
243
20
10
10
10 2
80
650 °C
10 3
10 4
10 5
10 6
10
10
10 2
80
700 °C
60
60
40
40
20
20
10
10
10 2
80
750 °C
10 3
10 4
10 5
10 6
10
10
10 2
10 4
10 5
10 6
10 3 10 4
Time [h]
10 5
10 6
10 3
60
40
Fig. 327. Change in chemical composition of
M in M23C6 carbides within grains in
specimen head portion as a function of time
for the heat AAL of JIS SUS316HTB steel.
20
10
10
50
10 3 10 4
Time [h]
10 2
10 5
10 6
50
550 °C
600 °C
40
30
20
40
Fe
Cr
Mo
Ni
30
20
Chemical composition of Laves phase [ %]
10
10
0
10
10 2
50
650 °C
40
10 3
10 4
10 5
10 6
0
10
50
10 4
10 5
10 6
10 3 10 4
Time [h]
10 5
10 6
10 3
700 °C
40
30
30
20
20
10
10
0
10
10 2
50
750 °C
40
10 2
10 3
10 4
10 5
10 6
0
10
10 2
30
Fig. 328. Change in chemical composition of
Laves phase within grains in specimen head
portion as a function of time for the heat AAL
of JIS SUS316HTB steel.
20
10
0
10
10 2
10 3 10 4
Time [h]
Landolt-Börnstein
New Series VIII/2B
10 5
10 6
244
60
50
2.4 Austenitic stainless steels
40
30
20
Chemical composition of s phase [ %]
60
550 °C
600 °C
50
Fe
Cr
Mo
Ni
10
0
10
10 2
60
650 °C
50
40
No precipitation
30
20
10 3
10 4
10 5
10
0
10 6 10
60
10 4
10 5
10 6
10 3 10 4
Time [h]
10 5
10 6
10 3
700 °C
50
40
10 2
40
30
30
20
20
10
0
10
10 2
60
750 °C
50
10 3
10 4
10 5
10
0
10 6 10
10 2
40
30
Fig. 329. Change in chemical composition of
σ phase within grains in specimen head
portion as a function of time for the heat AAL
of JIS SUS316HTB steel.
20
10
0
10
10 2
260
10 5
10 6
Specimen head
as - received
550 °C
600 °C
650 °C
700 °C
750 °C
240
220
200
180
Vickers hardness [HV5]
10 3 10 4
Time [h]
160
140
120
260
Specimen gauge
240
220
200
180
160
140
120
1 10
10 2
10 4
10 3
Time [h]
10 5
10 6
Fig. 330.
Comparison of Vickers hardness for
specimens tested at various temperatures, (a) specimen
head and (b) gauge portions of the heat AAL of JIS
SUS316HTB steel.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 246]
2.4.2 18Cr-12Ni-Mo steel
245
30
650 °C, 88 MPa, t r = 55367.0 h
700 °C, 61 MPa, t r = 34728.8 h
750 °C, 37 MPa, t r = 32781.8 h
Creep strain [%]
25
20
15
10
5
Fig. 331. Creep curves and creep-interruption conditions
for creep void observations for the heat AAL of JIS
SUS316HTB steel.
open :interrupted points
solid :ruptured points
0
0
10000
20000
30000 40000
Time [h ]
50000 60000
A parameter
0.3
A parameter
750 °C, 37 MPa, t r = 32781.8 h
700 °C, 61 MPa, t r = 34728.8 h
650 °C, 88 MPa, t r = 55367.0 h
0.2
0.1
0
0.04
Area fraction of
creep voids [%]
Area fraction
0.03
0.02
0.01
0
250
Number density of
creep voids [mm2 ]
Number density
200
150
100
Fig. 332. A parameter, area fraction and number density
of creep voids as a function of life consumption rate t/tr
for the heat AAL of JIS SUS316HTB steel.
50
0
0
Landolt-Börnstein
New Series VIII/2B
0.2
0.8
0.4
0.6
Life consumption rate, t / t r
1.0
246
2.4 Austenitic stainless steels
2.4.2.8.4 Creep voids in JIS SUS316HTB steel
The creep voids were observed to form during creep, which is more significant at lower stress and longer
time conditions. The area fraction and number density of creep voids were measured on the SEM
micrographs, after interruption of the creep tests at several creep strains, as shown in Fig. 331. The
development of creep voids during creep is shown in Fig. 332, as a function of time normalized by time to
rupture tr. The creep voids form at the later stage of creep substantially above t/tr = 0.4. The creep voids
were observed to form mainly at the interface between the σ phase on grain boundaries and austenite
matrix, reflecting the σ/matrix interface cracking along grain boundaries shown in the creep fracture
mode diagram in Fig. 322.
2.4.2.9 References
[1] Lai, J. K. L.: Materials Science and Engineering, 61 (1983), 101-109.
[2] National Research Institute for Metals (NRIM) Creep Data Sheet, No.6B (2000) for 18Cr-12Ni-Mo
stainless steel tubes, JIS SUS 316H TB.
[3] National Research Institute for Metals (NRIM) Creep Data Sheet, No.14B (1988) for 18Cr-12Ni-Mo
stainless steel plates, JIS SUS 316-HP.
[4] National Research Institute for Metals (NRIM) Creep Data Sheet, No.45 (1997) for 18Cr-12Ni-Momiddle N-low C hot rolled stainless steel plate, JIS SUS 316-HP.
[5] National Research Institute for Metals (NRIM) Creep Data Sheet, No.15B (1988) for 18Cr-12Ni-Mo
stainless steel bars, JIS SUS 316-B.
[6] Elevated temperature properties for steels for pressure purpose, Part I, British Standards Institution
(BSI), (1990).
[7] Shinya, N., Tanaka, H., Murata, M., Kaise, M. and Yokoi, S.: Tetsu-To-Hagane, 71 (1985), 114120.
[8] National Institute for Material Science Creep Data Sheet, Metallographic Atlas of Long-Term Crept
Materials, National Institute for Materials Science, Tsukuba, No.M-2 (2003).
[9] Tanaka, H., Murata, M., Abe, F. and Yagi, K.: Proceedings of the 28th MPA-Seminar, Stuttgart
University, Stuttgart, Germany (2002), 38.1-38.10.
[10] Lai, J. K. and Wickents, A.: Acta Metall., 27 (1979), 217-230.
[11] Mimino, T., Kinoshita, K., Shinoda, T. and Minegishi, I.: Tetsu-To-Hagane, 54 (1968), 464-472.
[12] Morris, D. G. and Harries, D. R.: Metal Science, 12 (1978), 525-549.
[13] Weiss, B. and Stickler, R.: Metallurgical Transactions, 3 (1972), 851-866.
[14] Lai, J. K. L.: Materials Science and Engineering, 58 (1983), 195-209.
[15] Lai, J. K. L.: Materials Science and Engineering, 61 (1983), 101-109.
[16] Minami, Y., Kimura, H. and Ihara, Y.: Mater. Sci. Technol. 2 (1986), 795-806.
Landolt-Börnstein
New Series VIII/2B
2.4.3 18Cr-10Ni-Ti steel
247
2.4.3 18Cr-10Ni-Ti steel
2.4.3.1 Introduction
18Cr-10Ni-Ti steel is an austenitic stainless steel with an addition of about 0.4 wt% Ti on 18Cr-10Ni base
compositions. This steel was developed in order to improve the corrosion resistance of 18Cr-10Ni steels.
However, it is widely used as super-heater and/or re-heater tubes in boilers because its creep rupture
strength is higher than that of 18Cr-10Ni steel. In the 1950s there was a burst problem of boiler tubes
consisting of 18Cr-10Ni-Ti steel in the USA. The reason was the low creep rupture strength due to lower
solution heat treatment temperature. After that higher solution heat treatment temperature was specified
and coarse grain was required.
2.4.3.2 Chemical composition
Chemical requirements of 18Cr-10Ni-Ti steel tubes are shown in Table 141.
Table 141. Chemical requirements of 18Cr-10Ni-Ti steel.
Chemical composition [wt%]
Standards Designation
C
Si
Mn
P
S
Ni
0.049.00JIS
SUS321HTB
≤0.75
≤2.00
≤0.030 ≤0.030
0.10
13.00
0.049.00ASTM
TP321H
≤0.75
≤2.00
≤0.040 ≤0.030
0.10
13.0
0.049.00BS
321S51
≤1.00
≤2.00
≤0.040 ≤0.030
0.10
12.0
9.0DIN
X6CrNiTi1810 ≤0.08
12.0
Cr
17.0020.00
17.020.0
17.019.0
17.019.0
Ti
4×C%
- 0.60
4×C%
- 0.60
5×C%
- 0.80
5×C%
- 0.80
2.4.3.3 Mechanical properties
The mechanical requirements of this steel is as follows:
Tensile strength is more than 520 MPa. Yield strength is more than 205 MPa. Elongation is more than
35 % in the case of outer diameter of tubes over 20 mm.
2.4.3.4 Creep rupture properties
The creep rupture data of 18Cr-10Ni-Ti steel tubes is shown in Fig. 333 [1]. From [1], data on elongation,
reduction of area, minimum creep rate and microstructure of as-received materials and crept specimens
can also be obtained.
Landolt-Börnstein
New Series VIII/2B
248
2.4 Austenitic stainless steels
500
６００℃
６５０℃
７００℃
７５０℃
400
300
Stress (MPa)
200
100
90
80
70
60
50
40
Fig. 333.
Creep rupture
strength data of 18Cr-10Ni-Ti
steel; [1]. n indicates the total
number of data points.
30
20
n=278
10 1
10 2
10 3
10 4
10 5
10 6
Time to rupture (h)
Average creep rupture strength of 18Cr-10Ni-Ti steel predicted from curvilinear regression using the
Manson-Haferd parameter method is summarized in Table 142.
Table 142. Average creep rupture strength of 18Cr-10Ni-Ti steel [MPa]
Temperature
100 h
1,000 h
10,000 h
100,000 h
600 °C
307
244
166
100
650 °C
238
167
106
64
700 °C
168
110
70
750 °C
114
76
40
2.4.3.5 Reference
[1] National Research Institute for Metals: NRIM Creep Data Sheet, No. 5B (1987).
Landolt-Börnstein
New Series VIII/2B
2.4.4 18Cr-12Ni-Nb steel
249
2.4.4 18Cr-12Ni-Nb steel
2.4.4.1 Introduction
18Cr-12Ni-Nb steel is an austenitic stainless steel with an addition of about 0.8 wt% Nb on 18Cr-12Ni
base compositions. It was developed in order to improve the corrosion resistance of 18Cr-10Ni steels.
However, it is widely used as super-heater and/or re-heater tubes in boilers because its creep rupture
strength is higher than that of 18Cr-10Ni steel.
2.4.4.2 Chemical composition
Chemical requirements of 18Cr-12Ni-Nb steel tubes are shown in Table 143.
Table 143. Chemical requirements of 18Cr-12Ni-Nb steel.
Chemical composition [wt%]
Standards Designation
C
Si
Mn
P
S
Ni
0.04≤1.00
≤2.00
≤0.030 ≤0.030 9.00JIS
SUS347HTB
13.00
0.10
0.04≤0.75
≤2.00
≤0.040 ≤0.030 9.00ASTM
TP347H
0.10
13.0
0.04≤1.00
≤2.00
≤0.040 ≤0.030 9.0BS
347S51
0.10
13.0
9.0≤0.08
DIN
X6CrNiNb1810
12.0
Cr
17.0020.00
17.0020.0
17.019.0
17.019.0
Nb
8×C%
-1.00
8×C%
-1.0
10×C%
-1.2
10×C%
-1.00
2.4.4.3 Mechanical properties
The mechanical requirements of this steel is as follows:
Tensile strength is more than 520 MPa. Yield strength is more than 205 MPa. Elongation is more than 35
% in the case of outer diameter of tubes over 20 mm.
2.4.4.4 Creep rupture properties
The creep rupture data of 9 heats obtained from [1] is shown in Fig. 334. From [1] data on elongation,
reduction of area, minimum creep rate and microstructure of as-received materials and crept specimens
can also be obtained.
Landolt-Börnstein
New Series VIII/2B
250
2.4 Austenitic stainless steels
500
400
６００℃
300
７００℃
６５０℃
７５０℃
Stress (MPa)
200
100
90
80
70
60
50
40
Fig. 334.
Creep rupture
strength data of 18Cr-12Ni-Nb
steel; [1]. n indicates the total
number of data points.
30
n=206
20
10
1
10
2
10
3
10
4
10
5
Time to rupture (h)
The creep rupture strength of 9 heats predicted from curvilinear regression using the Manson-Haferd
parameter method is summarized in Table 144. There is a large difference in creep rupture strength
between the heats.
Table 144. Creep rupture strength of 18Cr-12Ni-Nb steel tubes [MPa]
Temperature 100 h
1,000 h
10,000 h 100,000 h
600 °C
283-308 222-303
153-242
85-181
650 °C
223-290 157-223
95-165
53-109
700 °C
161-220 103-155
59-102
38-60
750 °C
111-167 64-114
43-70
2.4.4.5 Reference
[1] National Research Institute for Metals: NRIM Creep Data Sheet, No. 28B (2001).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 257]
2.4.5 Fine grained 18Cr-12Ni-Nb steel
251
2.4.5 Fine-grained 18Cr-12Ni-Nb steel
2.4.5.1 Introduction
Fine-grained 18Cr-12Ni-Nb austenitic steel (TP347HFG) is widely used as superheater and reheater tubes
in fossile fired boilers. The steel has been developed for improving steam oxidation resistance of
conventional TP347H stainless steel by grain refinement through a specially established thermomechanical process. The microstructure consists of a austenitic fine-grained matrix strengthened by
M23C6 carbides mainly along grain boundaries and finely dispersed NbC carbides in the matrix. NbC is
fine and stable even after long-term creep exposure at high temperatures.
2.4.5.2 Material standards, and chemical and tensile requirements
Tables 145 and 146 give the chemical requirements and the corresponding tensile requirements for finegrained 18Cr-12Ni-Nb steel tubes which are designated by the standards; ASTM A213 TP347HFG,
ASME Sec. I CC 2159, EN 10216-5
Table 145. Chemical requirements of fine-grained 18Cr-12Ni-Nb steel tubes; ASTM A213 TP347HFG,
ASME Sec. I CC 2159, EN 10216-5.
Chemical composition [wt%]
DesigGrade
Std. No.
nation
C
Si
Mn P
S
Ni
Cr
Nb+Ta
0.06 ≤
9.00 17.0 ≤
(Nb+Ta)/C
≤
≤
≤
A213
ASTM TP347HFG
0.10 0.75 2.00 0.040 0.030 13.0 20.0 1.0 >8
0.06 ≤
9.00 17.0 ≤
(Nb+Ta)/C
≤
≤
≤
EN
TP347HFG
10216-5
0.10 0.75 2.00 0.040 0.030 13.0 20.0 1.0 >8
Table 146. Tensile requirements of fine-grained 18Cr-12Ni-Nb steel tubes; ASTM A213 TP347HFG,
ASME Sec. I CC 2159, EN 10216-5
Designation
Grade
Min. TS1)
Min. YS2)
Min.
elongation
Standard No.
ASTM
TP347HFG
550 MPa
205 MPa
35 %
A213
EN
TP347HFG
550 MPa
205 MPa
35 %
10216-5
1) TS; Tensile strength, 2) YS; Yield strength as 0.2% proof stress
2.4.5.3 Tensile properties of fine-grained 18Cr-12Ni-Nb steel tubes
Fig. 335 shows tensile strength and yield stress data of fine-grained 18Cr-12Ni-Nb steel tubes [1]. They
are higher than those of conventional TP347H steel for temperatures up to 750 °C. The corresponding
tensile elongation and reduction of area data of fine-grained 18Cr-12Ni-Nb steel tubes are in the same
level of those of TP347H steel, which are available in [1], [3] and [4].
Landolt-Börnstein
New Series VIII/2B
252
2.4 Austenitic stainless steels
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
100
200
300 400 500
Temperature [°C]
600
700
800
Fig. 335. Tensile strength (circles) and yield stress
(triangles) data of fine-grained 18Cr-12Ni-Nb steel tubes.
2.4.5.4 Creep rupture properties of fine-grained 18Cr-12Ni-Nb steel tubes
2.4.5.4.1 Creep rupture data of fine-grained 18Cr-12Ni-Nb steel tubes
Fig. 336 shows creep rupture data of fine-grained 18Cr-12Ni-Nb steel tubes with average curves
according to the Larson-Miller parameter method [1]. The longest creep rupture time of fine-grained
18Cr-12Ni-Nb steel tubes is about 60000 h at 600 °C. Their long term creep strength is very stable and no
degradation in creep strength is expected at temperatures up to 750 °C. Fig. 337 shows a Larson-Miller
parameter plot of the creep rupture data of fine-grained 18Cr-12Ni-Nb steel tubes with a master rupture
curve and a 95% confidence lower limit. The best fitting was achieved with the optimized constant of
19.14.
500
400
300
Stress [MPa]
200
650 °C
700 °C
750 °C
800 °C
600 °C
100
80
60
600 °C
650 °C
700 °C
750 °C
800 °C
Average curve
40
20
10
1
10
10 2
10 3
Rupture time [h]
10 4
10 5
Fig. 336. Creep rupture strength
data of fine-grained 18Cr-12NiNb steel tubes.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 257]
2.4.5 Fine grained 18Cr-12Ni-Nb steel
500
400
300
600 °C
×10 5h
650 °C
×10 5h
200
Stress [MPa]
253
700 °C
×10 5h
100
80
60
600 °C
650 °C
700 °C
750 °C
800 °C
average curve
minimum curve
40
20
10
16
17 18 19 20 21 22 23 24 25 26
Larson-Miller-parameter T (19.14 + log t ) [×10 -3 ]
Fig. 337. Larson-Miller parameter plot of the creep
rupture data of fine-grained 18Cr-12Ni-Nb steel tubes.
2.4.5.4.2 Creep data of fine-grained 18Cr-12Ni-Nb steel tubes
Fig. 338 shows minimum creep rate data of fine-grained 18Cr-12Ni-Nb steel tubes measured at various
stress levels in the temperature range between 600 °C and 900 °C with average curves according to the
Larson-Miller parameter method [1]. Fig. 339 shows a Larson-Miller parameter plot of the minimum
creep rate data of fine-grained 18Cr-12Ni-Nb steel tubes with a master minimum creep rate curve. The
best fitting was achieved with the optimized constant of 32.88.
500
400
300
600 °C
200
Stress [MPa]
650 °C
100 700 °C
80
60
750 °C
40
20
800 °C
850 °C
10
10 -2
Landolt-Börnstein
New Series VIII/2B
10 -1
900 °C
Average curve
1
10
Minimum creep rate [% /10 3 h]
10 2
600 °C
650 °C
700 °C
750 °C
800 °C
850 °C
900 °C
10 3
Fig. 338. Minimum creep rate
data of fine-grained 18Cr-12NiNb steel tubes.
254
2.4 Austenitic stainless steels
500
400
300
600 °C
0.01%/10 3h
650 °C
0.01%/10 3h
700 °C
0.01%/10 3h
Stress [MPa]
200
100
80
60
750 °C
0.01%/10 3h
600 °C
650 °C
700 °C
750 °C
800 °C
850 °C
900 °C
40
20
average curve
10
26
Fig. 339. Larson-Miller parameter plot of the minimum
creep rate data of fine-grained 18Cr-12Ni-Nb steel
tubes.
28
40
36
30
32
34
38
Larson-Miller-parameter T (32.88 + log e) [×10 -3 ]
2.4.5.5 Allowable tensile stress of fine-grained 18Cr-12Ni-Nb steel tubes
Fig. 340 shows the allowable stress designated for fine-grained 18Cr-12Ni-Nb steel according to the
ASME standard procedure comparing with that for the conventional steel TP347H.
180
Allowable tensile stress (MPa)
160
TP347HFG
140
120
100
80
TP347H
60
40
20
Fig. 340. Allowable tensile stress for fine-grained 18Cr12Ni-Nb steel tubes according to the ASME standard.
0
0
100 200 300 400 500 600 700 800
Temperature (℃)
2.4.5.6 Manufacturing process of fine-grained 18Cr-12Ni-Nb steel tubes
Fig. 341 shows the manufacturing process of fine-grained 18Cr-12Ni-Nb steel tubes, which is
characterized by the double stage heat treatment to achieve fine grain microstructure with very fine
dispersion of NbC in the matrix. In the new process, NbC resolves into the matrix during the pre-solution
treatment at higher temperatures and re-precipitates finely in the matrix during the subsequent solution
treatment at lower temperatures. This gives rise to the fine grain microstructure with very fine dispersion
of NbC in the matrix.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 257]
2.4.5 Fine grained 18Cr-12Ni-Nb steel
255
Fine NbC
precipitation
Cold
working
CW≧30%
Fine NbC
dissolution
Solution
Pre-solution
Fig. 342 shows the initial microstructures of fine-grained 18Cr-12Ni-Nb steel tubes and conventional
TP347H steel. Homogeneous and very fine grain structure has been achieved by applying the double
stage heat treatment in the new process.
Fig. 343 shows the change in the amount of Nb precipitated as NbC after pre-solution treatment at
various temperatures. More NbC precipitated when applying the lower temperature solution treatment
after a higher temperature pre-solution treatment.
Fig. 344 shows the effect of the final-solution treatment temperature on the grain size of fine-grained
18Cr-12Ni-Nb steel tubes and the conventional TP347H steel. The grain size of fine-grained 18Cr-12NiNb steel tubes is much smaller than that of the conventional steel even with increasing the solution
treatment temperature. This has been achieved by more precipitation of fine NbC carbides during the presolution treatment at lower temperatures.
TP347HFG
CW 20～30%
Conventional
TP347H
ASTM
ASTM
G.S No. 8
G.S No. 4～5
Fig. 341. A comparison of manufacturing processes for fine-grained 18Cr-12Ni-Nb steel and the conventional
TP347H steel tubes. Left: New process; right: conventional process.
(a) Fine-grained 18Cr-12Ni-Nb steel
(b) conventional TP347H
Fig. 342. Optical micrographs of the fine grain microstructure in fine-grained 18Cr-12Ni-Nb steel by taking the
double stage heat treatment (a) and of the microstructure in conventional TP347H steel (b).
Landolt-Börnstein
New Series VIII/2B
256
2.4 Austenitic stainless steels
0.8
0.6
0.4
0.2
0
12
Pre-solution treatment
○ 1250℃× 10min
● 1300℃× 10min
10
Grain size [ASTM No.]
Extracted Nb content (%)
1.0
Precipitated Nb as NbC
during final-solution treatment
As presolution
1100
1150
1200
1250
Solution treatment temperature ( ℃ )
Fig. 343. Effects of pre- and final-solution treatment
temperatures on the amount of NbC precipitation.
8
6
4
new process
conventional process
2
0
1100
1300
1200
Solution treatment temperature [°C]
Fig. 344. Effect of solution treatment temperatures on
grain size and the resultant creep strength of finegrained 18Cr-12Ni-Nb steel.
2.4.5.7 Corrosion resistance of fine-grained 18Cr-12Ni-Nb steel tubes
Inner-scale thickness (µm)
Fig. 345 shows the oxidation behavior of fine-grained 18Cr-12Ni-Nb steels with different grain size
tested at 650 °C and 700 °C. It can be seen that the steam oxidation resistance is improved by reduction of
the grain size. Steam oxidation resistance is found to be achieved by the thin and tight protective Cr2O3
corundum type inner oxide layer formed in the fine-grained steel (see Fig. 346).
650℃
□
○
△
100
700℃
■
TP347H（
GS6）
TP347HFG(GS8）
▲
●
TP347HFG(GS9）
50
20
In steam
10
500
Outer scale
1000
Fig. 345. Steam oxidation resistance of fine-grained
18Cr-12Ni-Nb steel tested at 650 °C and 700 °C.
3000
Time (h)
Inner scale
Metal
Outer scale
Inner scale
Metal
Fe3O4
Fe3O4
Cr2O3
(Fe,Cr)3O4
Cr2O3
a) Fine-grained
(Fe,Cr)3O4
Fig. 346.
Schematic
illustrations of oxide layers
to be formed in fine- and
coarse-grained steels.
b) Coarse-grained
Landolt-Börnstein
New Series VIII/2B
Ref. p. 257]
2.4.5 Fine grained 18Cr-12Ni-Nb steel
257
2.4.5.8 Microstructural change of fine-grained 18Cr-12Ni-Nb steel during long term creep
exposure
Fig. 347 shows TEM micrographs of crept specimens from fine-grained 18Cr-12Ni-Nb and coarsegrained TP347H steel. The initial microstructures are completely different from each other. In finegrained 18Cr-12Ni-Nb steel fine NbC carbide precipitates homogeneously and acts as obstacle against
dislocation motion even after creep exposure, resulting in higher creep strength. In conventional TP347H
steel the size and distribution of NbC carbide is inhomogeneous, which is the main reason for lower creep
strength.
As-solution treated
1µm
crept at 750℃, 98 MPa
crept at 650℃, 196 MPa
ruptured after 643h
ruptured after 2958h
(a) Fine-grained 18Cr-12Ni-Nb steel (TP347HFG)
ruptured after 399h
ruptured after 9746h (177MPa)
(b) Conventional TP347H with coarse-grain
Fig. 347. TEM micrographs of the crept specimens from fine-grained 18Cr-12Ni-Nb steel and the conventional
TP347H steel with coarse-grain.
2.4.5.9 Performance of service exposed tubes
Performance of service exposed tubes is available in [7] and [9].
2.4.5.10 References
[1] Sumitomo seamless tubes and pipe Creep Data Sheets, Sumitomo Metal Industries, (1993).
[2] Teranishi, H., Yoshikawa, K., Fujikawa, H., Kubota, M., Tokimasa, K., and Miura, M.: Proc. of
Int.Conf. on Coatings and Bi-Metallics for Energy Systems Chemical Process Environment, (1984).
[3] Teranishi, H., Fujikawa, H., Yoshikawa, K., Kubota, M., Yamamoto, S.: Sumitomo Metals, 36
(1984), 134.
[4] Yukitoshi, T., Yoshikawa, K., Teranishi, Y.: Tetsu-to-Hagane, 70 (1984), 1962.
[5] Yuzawa, H., Teranishi, H., Fujikawa, H., Yoshikawa, K., Kubota, M.: The Thermal and Nuclear
Power, 36 (1985), 1325.
[6] Teranishi, H., Yoshikawa, K., and Sawaragi, Y.: Proc. of Int. Conf. on Creep, (1986), 233.
[7] Teranishi, H., Sawaragi, Y., Kubota, M., Hayase, Y.: The Sumitomo Search, No.38 (1989), 63.
[8] Sawaragi, Y., Teranishi, H., Iseda, A., Yoshikawa, K.: Sumitomo Metals, 42 (1990), 260.
[9] Sawaragi, Y., Otsuka, N., Senba, H., and Yamamoto, S.: Sumitomo Search, No.56 (1994), 34.
Landolt-Börnstein
New Series VIII/2B
258
2.4 Austenitic stainless steels
2.4.6 16Cr-13Ni-Nb steel
2.4.6.1 Introduction
16Cr-13Ni-Nb steel (X8CrNiNb16-13, 1.4961, 16-13Nb) is a creep resisting austenitic steel used
primarily in Germany. It has a similar composition to AISI Type 347 and is used for components in gas
and steam turbines, turbo-rotor wheels, turbine rotor power plants, superheater and heat exchanger
tubings and steam pipes.
2.4.6.2 Material standards, chemical composition and tensile requirements
Table 147. Chemical requirements of X8CrNiNb16-13 concerning EN 10028-7 (plate) and EN 10216-5
(tube).
Chemical composition [wt%]
StanStd. No Designation
dard
C
Si
Mn
P
S
Cr
Ni
Others
0.04 0.30
15.00 12.00 Nb ≥
EN
10028-7 X8CrNiNb16-13Nb
≤1.50 ≤0.035 ≤0.015
0.10 0.50
17.00 14.00 10×C-1.2
0.04 0.30
15.00 12.00 Nb ≥
EN
10216-5 X8CrNiNb16-13Nb
≤1.50 ≤0.035 ≤0.015
0.10 0.50
17.00 14.00 10×C-1.2
X8CrNiNb16-13Nb is usually solution heat treated at 1050 to 1110°C.
Table 148. Room temperature minimum mechanical property requirements for X8CrNiNb16-13
Rp1.0
Rm
Rp0.2
Standard Std. No. Designation
[Nmm-2] [Nmm-2] [Nmm-2]
EN
10028-7 X8CrNiNb16-13 205
245
510
Table 149. Minimum 0.2% and 1.0% proof strength values at elevated temperatures for X8CrNiNb16-13
Minimum 0.2% (Rp0.2) and 1.0% (Rp1.0) proof strength, at
StanStd. No. Designation
various temperatures.
dard
T [°C]
100 150 250 350 450 500 550
Rp1.0 [Nmm-2] 175 157 137 128 118 118 113
EN
10028-7 X8CrNiNb16-13
Rp0.2 [Nmm-2] 205 186 167 157 147 147 142
2.4.6.3 Creep rupture strength
The creep rupture strength of X8CrNiNb16-13 is shown in Fig. 348. The analysis from which the data in
the figure are derived was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found from their published data sheets [1].
The test data were available from 28 casts with test temperatures of 500 - 800 °C. The test distribution
of test durations are shown in Table 150.
Table 150. Distribution of test durations used to derive the stress rupture properties of X8CrNiNb 16-13.
Number of test points at the various test durations
20,001- 30,00170,001<10,000 h 10,000-20,000 h
50,001-70,000 h
>100,000 h
30,000 h 50,000 h
100,000 h
209
28
17
13
12 (5)
(7)
1 (4)
( ) denotes unbroken tests
The assessment was made by the German Creep Committee in 1969 by using the graphical method [2, 3].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 259]
2.4.6 16Cr-13Ni-Nb steel
259
1000
Stress [MPa]
10,000h
100,000h
200,000h
100
Fig. 348. Creep rupture strength data of
X8CrNiNb16-13.
10
550
600
650
700
750
800
Temperature [°C]
2.4.6.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 348 the 100,000 h and 200,000 h rupture strength values for a range of
temperatures are as follows:
100,000 h rupture strengths at specified temperatures
Temperature [°C]
580
590
600
610
620
630
640
650
660
Stress [Nmm-2]
129
119
108
98
89
80
72
64
57*
Temperature [°C]
Stress [Nmm-2]
100,000 h rupture strengths at specified temperatures
670
680
690
700
710
720
730
50*
44*
39*
34*
30*
27*
25*
740
22*
750
20*
Temperature [°C]
Stress [Nmm-2]
200,000 h rupture strengths at specified temperatures
580
590
600
610
620
630
640
115*
105*
94*
85*
77*
69*
61*
650
53*
660
47*
740
17*
750
15*
200,000 h rupture strengths at specified temperatures
Temperature [°C]
670
680
690
700
710
720
730
Stress [Nmm-2]
41*
36*
31*
27*
25*
22*
19*
* Values which have involved extended time extrapolation.
2.4.6.5 References
[1] ECCC Data sheet for X8CrNiNb16-13, European Creep Collaborative Committee, BRITE EURAM
Thematic Network BET2-0509 “Weld Creep”, (1999).
[2] G. Bandel and H. Gravenhorst: Verhalten warmfester Stähle im Zeitstandversuch bei 500 bis 700°C,
Teil II. Auswertungsverfahren, Archiv Eisenhüttenwes. 28 (1957), pp.253-258.
[3] ECCC- WG1 Recommendations “Data Assessment” volume 5, part 1, app. D5; European Creep
Collaborative Committee (1996).
Landolt-Börnstein
New Series VIII/2B
260
2.4 Austenitic stainless steels
2.4.7 18Cr-9Ni-3Cu-Nb-N steel
2.4.7.1 Introduction
18Cr-9Ni-3Cu-Nb-N austenitic steel (SUPER304H) is used as superheater and reheater tubes in fossile
fired boilers. The steel has been developed for substituting conventional 304H and 321H steels by the
addition of copper and nitrogen to increase creep strength at elevated temperatures and toughness after
long term exposure at high temperatures. The microstructure consists of an austenitic matrix strengthened
by M23C6 carbides mainly along grain boundaries and finely dispersed Cu-phase and NbCrN nitrides in
the matrix. The Cu-phase is fine and coherent to the matrix which gives rise to a significant increase in
creep strength at elevated temperatures. NbCrN is also fine and stable even after long term exposure at
high temperatures. No σ phase is expected to form even after 100000 h in the temperature range between
600 and 800 °C, which is mainly achieved by stabilization of the austenitic matrix with copper and
nitrogen.
2.4.7.2 Material standards, chemical and tensile requirements
Tables 151 and 152 give the chemical requirements and the corresponding tensile requirements of 18Cr9Ni-3Cu-Nb-N steel tubes which are designated by the standards; Japanese METI KA-SUS304J1HTB,
ASTM A213 UNS No.S30432 (ASME Sec. I CC 2328).
Table 151. Chemical requirements of 18Cr-9Ni-3Cu-Nb-N steel tubes; Japanese
SUS304J1HTB, ASTM A213 UNS No.S30432 (ASME Sec. I CC 2328), and EN 10216-5.
Chemical composition [wt%]
DesigGrade
nation
C
Si
Mn P
S
Ni
Cr
Nb N
Cu Al
7.50 17.00 0.30 0.05 2.50
Japanese (1)
0.07 ≤
≤
≤
≤
METI
0.13 0.30 1.00 0.040 0.010 10.50 19.00 0.60 0.12 3.50
7.50 17.00 0.20 0.05 2.50 0.003
ASTM (2)
0.07 ≤
≤
≤
≤
0.13 0.30 1.00 0.040 0.010 10.50 19.00 0.60 0.12 3.50 0.030
7.50 17.00 0.20 0.05 2.50 0.003
EN
(3)
0.07 ≤
≤
≤
≤
0.13 0.30 1.00 0.040 0.010 10.50 19.00 0.60 0.12 3.50 0.030
(1) KA-SUS304J1HTB; (2) UNS No. S30432; (3) SUPER304H
METI KA-
B
Std.
No.
0.001 A213
0.010
0.001 10216
0.010 -5
Table 152.
Tensile requirements of 18Cr-9Ni-3Cu-Nb-N steel tubes; Japanese METI KASUS304J1HTB, ASTM A213-01 UNS No.S30432 (ASME C.C.2328), and EN 10216-5.
Designation
Grade
Min. TS1) Min. YS2) Min elongation Standard No.
Japanese METI KA-SUS304J1HTB 590 MPa 235 MPa 35 %
ASTM
UNS No.S30432
590 MPa 235 MPa 35 %
A213
EN
SUPER304H
590 MPa
235 MPa
35 %
10216-5
1) TS; tensile strength, 2) YS; yield strength as 0.2% proof stress
2.4.7.3 Tensile properties of 18Cr-9Ni-3Cu-Nb-N steel tubes
Fig. 349 shows tensile strength and yield stress data of 18Cr-9Ni-3Cu-Nb-N steel tubes [1], [3] and [4].
They are higher than those of TP304H steel for temperatures up to 750 °C. The corresponding tensile
elongation and reduction of area data of 18Cr-9Ni-3Cu-Nb-N tubes are in the same level as those of
TP304H steel, which are available in [1].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 264]
2.4.7 18Cr-9Ni-3Cu-Nb-N steel
261
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
100
200
300 400 500
Temperature [°C]
600
700
800
Fig. 349. Tensile strength (circles) and yield stress
(triangles) data of 18Cr-9Ni-3Cu-Nb-N steel tubes.
2.4.7.4 Creep rupture properties of 18Cr-9Ni-3Cu-Nb-N steel tubes
2.4.7.4.1 Creep rupture data of 18Cr-9Ni-3Cu-Nb-N steel tubes
Fig. 350 shows creep rupture data of 18Cr-9Ni-3Cu-Nb-N steel tubes with average curves assessed by the
Larson-Miller parameter method [1]. The longest creep rupture datum of 18Cr-9Ni-3Cu-Nb-N steel tubes
is over 85000 h at 600 °C. Their long-term creep strength is very stable and no degradation in creep
strength is expected for temperatures up to 750 °C. Fig. 351 shows a Larson-Miller parameter plot of the
creep rupture data of 18Cr-9Ni-3Cu-Nb-N steel tubes with a master rupture curve and a 95 % confidence
lower limit. The best fitting was achieved with an optimized constant of 19.68.
500
400
Stress [MPa]
300
600 °C
200
650 °C
100
80
60
600 °C
650 °C
700 °C
750 °C
average curve
40
10
Landolt-Börnstein
New Series VIII/2B
10 2
700 °C
750 °C
10 3
Rupture time [h]
10 4
10 5
Fig. 350. Creep rupture strength
data of 18Cr-9Ni-3Cu-Nb-N steel
tubes.
262
2.4 Austenitic stainless steels
500
600 °C
×10 5h
400
Stress [MPa]
300
650 °C
×10 5h
200
700 °C
×10 5h
750 °C
×10 5h
100
600 °C
650 °C
700 °C
750 °C
80
60
40
17
average curve
minimum curve
Fig. 351. Larson-Miller parameter plot of the creep
rupture data of 18Cr-9Ni-3Cu-Nb-N steel tubes.
18 19 20 21 22 23 24 25 26 27
Larson-Miller-parameter T (19.68 + log t ) [×10 -3 ]
2.4.7.4.2 Creep data of 18Cr-9Ni-3Cu-Nb-N steel tubes
Fig. 352 shows minimum creep rate data of 18Cr-9Ni-3Cu-Nb-N steel tubes measured at various stress
levels in the temperature range between 600 °C and 750 °C with average curves according to the LarsonMiller parameter method [1]. Fig. 353 shows a Larson-Miller parameter plot of the minimum creep rate
data of 18Cr-9Ni-3Cu-Nb-N steel tubes with a master curve. The best fitting was achieved with an
optimized constant of 26.96.
500
400
300
Stress [MPa]
600 °C
200
650 °C
100 700 °C
600 °C
650 °C
700 °C
750 °C
80
60 750 °C
40
average curve
10 -2
10 -1
1
10
Minimum creep rate [% /10 3 h]
10 2
10 3
Fig. 352. Minimum creep rate
data of 18Cr-9Ni-3Cu-Nb-N steel
tubes.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 264]
2.4.7 18Cr-9Ni-3Cu-Nb-N steel
263
500
400
600 °C
0.01%/10 3h
Stress [MPa]
300
650 °C
0.01%/10 3h
200
700 °C
0.01%/10 3h
100
600 °C
650 °C
700 °C
750 °C
80
60
40
20
Fig. 353. Larson-Miller parameter plot of the minimum
creep rate data of 18Cr-9Ni-3Cu-Nb-N steel tubes.
average curve
21 22 23 24 25 26 27 28 29 30
Larson-Miller-parameter T (26.96 + log e) [×10 -3 ]
2.4.7.5 Allowable stress of 18Cr-9Ni-3Cu-Nb-N steel tubes
Fig. 354 shows the allowable stress determined for 18Cr-9Ni-3Cu-Nb-N steel (Japanese METI KA-SUS
304J1HTB) according to the METI standard procedure comparing with that for the conventional steel,
ASME SA213-TP347H (Japanese METI KA-SUSTP347HTB).
180
T ensile region
Allowable tensile stress (MPa)
160
Solid solution
strengthening (N)
140
120
KA-SUS304J1HTB
C reep region
Precipitation
strengthening
Nb (C,N)
NbCrN
M 23C6
Cu phase
100
KA-SUSTP347HTB
80
60
40
20
0
0
100 200 300 400 500 600 700 800
Temperature (℃)
Fig. 354. Allowable tensile stress determined for 18Cr-9Ni-3Cu-Nb-N steel tubes.
2.4.7.6 Microstructural change of 18Cr-9Ni-3Cu-Nb-N steel tubes
The microstructural change in 18Cr-9Ni-3Cu-Nb-N steel tubes after aging for up to 10000 h in the
temperature range between 600 °C and 750 °C is available in [3] and [4]. There is no significant
microstructural change observed even after aging for 10000 h at 750 °C. A detailed TEM observation of
Landolt-Börnstein
New Series VIII/2B
264
2.4 Austenitic stainless steels
the specimens aged for 3000 h in the temperature range between 600 °C and 750 °C has shown that fine
coherent Cu phase is dispersed in the matrix as well as fine NbCrN nitrides and no harmful blocky
precipitation such as σ phase is found. This fine dispersion of the precipitates is the major strengthening
mechanism of this steel in creep regions at higher temperatures, which is schematically depicted in Fig.
354.
2.4.7.7 Performance of service exposed tubes
Performance of service exposed tubes is available in [5] and [8].
2.4.7.8 References
[1] Sumitomo seamless tubes and pipe Creep Data Sheets, Sumitomo Metal Industries, (1993).
[2] Sawaragi, Y., and Hirano, S.: Proc. of Int. Conf. on New Alloys for Pressure Vessels and Piping,
(1990).
[3] Sawaragi, Y., Otsuka, N., Ogawa, K., Kato, S., and Hirano, S.: Sumitomo Metals, 43 (1991), 24.
[4] Sawaragi, Y., Otsuka, N., Ogawa, K., Kato, S., and Hirano, S.: The Sumitomo Search, No.48
(1992), 50.
[5] Sawaragi, Y., Otsuka, N., Senba, H., and Yamamoto, S.: Sumitomo Search, No.56 (1994), 34.
[6] Sawaragi, Y., Hirano, S., and Masuyama, F.: Proc. of Int. Conf. on Microstructure and Mechanical
Properties of Aging Materials, (1992).
[7] Ogawa, K., Sawaragi, Y., Otsuka, N., Hirata, H., Natori, A., and Matsumoto, S.: ISIJ International,
35 (1995), 1258.
[8] Kan, T., Sawaragi, Y., Yamadera, Y., and Okada, H.: Proc. of 6th Liege Conf. on Materials for
Advanced Power Engineering, (1998), 60.
[9] Senba, H., Sawaragi, Y., Ogawa, K., Natori, A., and Kan, T.: Materia, 41 (2002), 120.
Landolt-Börnstein
New Series VIII/2B
2.4.8 18Cr-10Ni-Ti-Nb steel
265
2.4.8 18Cr-10Ni-Ti-Nb steel
2.4.8.1 Introduction
18Cr-10Ni-Ti-Nb steel (TEMPALOY A-1) is a austenitic stainless steel made by adding small amounts
of Ti and Nb to SUS304H for precipitation strengthening in order to improve its long-term creep rupture
strength. Its allowable stress at 650 °C for 105 h is 1.5 times that of conventional SUS321H and
SUS347H. It is suitable as material for super-heater tubes and re-heater tubes in boilers.
2.4.8.2 Chemical composition
Chemical requirement of 18Cr-10Ni-Ti-Nb steel tube is shown in Table 153.
Table 153. Chemical requirement of 18Cr-10Ni-Ti-Nb steel (mass%)
Chemical requirements [wt%]
Designation
C
Si
Mn
P
S
Ni
Cr
Ti
KA-SUS
0.07- ≤1.00 ≤2.00 ≤0.040 ≤0.030 9.00- 17.50- ≤0.20
321J1HTB 0.14
12.00 19.50
Nb
≤0.40
(Ti+Nb/2)/C
0.6-2.5
2.4.8.3 Mechanical properties
The room temperature mechanical requirements are as follows: tensile strength is more than 520 MPa,
yield strength is more than 205 MPa, and elongation is more than 35 % in the case of outer diameter of
tubes over 20 mm.
2.4.8.4 Creep and rupture properties
The creep rupture data obtained from NKK co. is shown in Fig. 355 [1]. The average creep rupture
strength predicted using the Larson-Miller parameter method is summarized in Table 154.
Table 154. Average creep rupture strength of 18Cr-10Ni-Ti-Nb steel [MPa]
Temperature 100 h 1,000 h 10,000 h 100,000 h
600 °C
304
240
191
139
650 °C
225
162
123
93
700 °C
157
108
74
58
750 °C
111
78
54
34
Landolt-Börnstein
New Series VIII/2B
266
2.4 Austenitic stainless steels
400
600
650
700
750
800
300
Stress (MPa)
200
100
90
80
70
60
50
40
30
20
10
1
10
2
3
4
10
10
Time to rupture (h)
10
5
10
6
Fig. 355.
Creep rupture
strength data of 18Cr-10Ni-TiNb steel obtained from NKK
co.
2.4.8.5 Reference
[1] NKK Technical Bullitin: TEC. No.243-312 Boiler Tubing.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 269]
2.4.9 20Cr-25Ni-1.5Mo-NbTiBN steel
267
2.4.9 20Cr-25Ni-1.5Mo-NbTiBN steel
2.4.9.1 Introduction
20Cr-25Ni-1.5Mo-NbTiBN steel [1] is designed for superheater tubes for Ultra Super Critical power
plants. The high chromium content improves corrosion resistance, and the well-balanced Nb, Ti and N
contents increase creep rupture strength compared with conventional austenitic heat resistant steels. 20Cr25Ni-1.5Mo-NbTiBN steel is designated as KA-SUS310J2TB in METI standard with a creep rupture
strength of 86 MPa at 700 °C for 100,000 h.
2.4.9.2 Material standards, chemical and tensile requirements
Table 155. Chemical requirements of 20Cr-25Ni-1.5Mo-NbTiBN steel tubes for superheater tubes;
KA-SUS310J2TB
Chemical composition [wt%]
StanDesignation
dards
C
Si
Mn
P
S
Cr
Mo
METI KA-SUS310J2TB ≤0.10 ≤1.00 ≤1.50 ≤0.030 ≤0.010 19.00-23.00 1.00-2.00
Standards
Designation
METI
KA-SUS310J2TB
Ni
22.0028.00
Nb
0.100.40
N
0.100.25
Ti
0.020.20
B
0.0020.010
2.4.9.3 Creep properties of 20Cr-25Ni-1.5Mo-NbTiBN steel tubes
[2] contains creep data of 20Cr-25Ni-1.5Mo-NbTiBN steel tubes, namely rupture data, minimum creep
rate, rupture elongation and reduction of area.
2.4.9.3.1 Creep rupture data of 20Cr-25Ni-1.5Mo-NbTiBN steel tubes
Fig. 356 shows the creep rupture data of 20Cr-25Ni-1.5Mo-NbTiBN steel tubes of 6 heats. Creep tests
continue over 50,000 h.
1000
Stress (MPa)
600°C
650°C
100
700°C
750°C
800°C
850°C
950°C
Fig. 356.
Creep rupture
strength data of 20Cr-25Ni1.5Mo-NbTiBN steel tubes;
[2].
900°C
10
10
100
1000
Time to Rupture (h)
Landolt-Börnstein
New Series VIII/2B
10000
100000
268
2.4 Austenitic stainless steels
2.4.9.3.2 Time-Temperature-Parametric prognostication of the creep rupture strength
Fig. 357 shows the Larson-Miller Parametric plot of the rupture data [2]. Fig. 358 shows a creep rupture
curve regression by a fifth-degree expression predicting the creep rupture strength for times longer than
that of the experiment for temperatures between 600 °C and 850 °C.
18000
20000
22000
800°C,105h
750°C,105h
700°C,105h
10
16000
650°C,105h
100
600°C,105h
Stress (MPa)
1000
24000
26000
Fig. 357. Master rupture curve by LarsonMiller parameter method for 20Cr-25Ni1.5Mo-NbTiBN steel tubes.
LMP=[T+273.15][17.825+log(tr)]
1000
Stress (MPa)
600°C
650°C
100
700°C
850°C
900°C
950°C
750°C
800°C
10
10
100
1000
10000
Fig. 358. Estimated creep
rupture curves of 20Cr-25Ni1.5Mo-NbTiBN steel tubes.
100000
Time to Rupture (h)
2.4.9.3.3 Microstructural changes during creep [3]
Fig. 359 shows an optical and an electron micrograph of steel crept at 700 °C for 5,000 h. Mainly, three
types of precipitates are observed in the crept specimens: Massive particles (type A) of 0.4 - 0.5 µm in
width, string Cr-Nb nitrides (type B) of less than 0.03 µm in width and fine granular and needle-like
M23C6 carbides (type C) of less than 0.2 µm in width. The long-term creep properties are presumed to be
due to the many fine precipitates, which are very fine particles and string type and needle type
precipitates, the depression of coarsening of the precipitates by Nb and Ti, and the absence of
precipitation of σ-phase.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 269]
Cr
2.4.9 20Cr-25Ni-1.5Mo-NbTiBN steel
269
b
a
c
Ni
Cr
Si Mo
Fe
Mo
Nb
Nb
Cr
Nb
Mo
Fe Ni
Mo
Fe
Mo
Fig. 359. Optical and electron micrographs and EDX analyses of steel crept at 700 °C for 5,000 h; [3].
2.4.9.4 References
[1] Kikuchi, M., Sakakibara, M., Otoguro, Y., Mimura, H. Takahashi, T., and Fujita, T.: The
International Conference on High Temperature Alloys, Petton (1985).
[2] Nippon Steel Corporation: Nippon Steel Creep Database (2000).
[3] Takahashi, T., Sakakibara, M., Kikuchi, M., Ogawa, T., Araki, S., and Fujita, T.: Tetsu-to-Hagane,
76 (1990), No. 7, 1131.
Landolt-Börnstein
New Series VIII/2B
270
2.4 Austenitic stainless steels
2.4.10 21Cr-32Ni-Ti-Al steel
2.4.10.1 Introduction
21Cr-32Ni-Ti-Al steel (Alloy 800H) is a version of Alloy 800 having higher creep and rupture strength.
The two alloys have the same chemical composition with the exception that the carbon content of Alloy
800H is restricted to the upper portion of the standard range for Alloy 800. In addition to a controlled
carbon content, Alloy 800H receives an annealing treatment that produces a coarse grain size.This alloy is
popularly useful for applications involving long term exposure to elevated temperatures or corrosive
atmospheres. It is used in chemical and petrochemical processing and often used in domestic appliances
for sheathing on electric heating elements.
2.4.10.2 Material standards, chemical and tensile requirements
Chemical requirements of 21Cr-32Ni-Ti-Al steel plates are shown in Table 156.
Table 156. Chemical requirements of 21Cr-32Ni-Ti-Al steel plates ; JIS NCF800H, ASTM B409, UNS
8810.
Chemical composition [wt%]
Specification
C
Cr
Mn
Cu
Ni
P
S
Si
JIS
NCF800H
0.050.10
19.00 23.00
≤1.50
≤0.75
30.00 35.00
≤0.030
≤0.015
≤1.00
ASTM B409
0.050.10
19.00 23.00
≤1.50
≤0.75
30.00 35.00
-
≤0.015
≤1.00
UNS
No.8810
0.050.10
19.00 23.00
1.50
0.75
30.00 35.00
0.015
0.75
Others
Al:0.15-0.60
Ti:0.15-0.60
Bal: Fe
Al:0.15-0.60
Ti:0.15-0.60
Bal: Fe
Al:0.15-1.00
Ti:0.15-0.60
Bal: Fe
2.4.10.3 Creep properties of 21Cr-32Ni-Ti-Al steel plates
Information of fact on creep data for 21Cr-32Ni-Ti-Al steel plates can be obtained from [1].
2.4.10.3.1 Creep rupture data of 21Cr-32Ni-Ti-Al steel plates
The creep rupture strength of 21Cr-32Ni-Ti-Al steel plates obtained from [1]is illustrated in Fig. 360.
Data on elongation, reduction of area and minimum creep rate is shown in Fig. 361-Fig. 364.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 272]
2.4.10 21Cr-32Ni-Ti-Al steel
600
500
400
300
600 °C
700 °C
800 °C
900 °C
1000 °C
200
Stress [MPa]
271
100
80
60
50
40
30
20
10
8
6
5
4
3
10
Fig. 360. Creep rupture strength
data of 21Cr-32Ni-Ti-Al steel
plates. n indicates the total
number of data points.
n = 155
100
10 2
10 3
10 4
Time to rupture [h]
600 °C
100
80
80
60
60
40
40
n = 22
600 °C
0 n = 22
700 °C
80
60
40
20
n = 28
100
100
Reduction of area [%]
Elongation [%]
100
0
10 6
20
20
0
10 5
60
40
20
0
800 °C
80
800 °C
80
60
60
40
20
20
Landolt-Börnstein
New Series VIII/2B
n = 28
100
40
n = 38
0
10 5
10
10 2 10 3 10 4
Time to rupture [h]
700 °C
80
10 6
0
10
n = 38
10 5
10 2 10 3 10 4
Time to rupture [h]
10 6
Fig. 361. Elongation and reduction of area at
600, 700 and 800 °C for 21Cr-32Ni-Ti-Al
steel plates. n indicates the total number of
data points in each diagram.
272
2.4 Austenitic stainless steels
100
900 °C
80
80
60
60
40
40
20
0
n = 40
100
1000 °C
80
Reduction of area [%]
Elongation [%]
100
20
0
1000 °C
80
60
40
40
20
20
10 6
n = 40
100
60
n = 27
0
10 5
10
10 2 10 3 10 4
Time to rupture [h]
900 °C
n = 27
0
10 5
10
10 2 10 3 10 4
Time to rupture [h]
400
300
600 °C
700 °C
800 °C
900 °C
1000 °C
10 5
100
80
60
Time to rupture [h]
Stress [MPa]
10 6
10 6
200
40
20
10
8
6
4
10 -7
Fig. 362. Elongation and reduction of area at
900 and 1000 °C for 21Cr-32Ni-Ti-Al steel
plates. n indicates the total number of data
points in each diagram.
600 °C
700 °C
800 °C
900 °C
1000 °C
n = 52
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
10 -1
10 4
10 3
10 2
n = 52
10
1
Fig. 363. Stress vs. minimum creep rate for 21Cr32Ni-Ti-Al steel plates. n indicates the total number
of data points.
10 -7
10 -6
10 -5 10 -4 10 -3 10 -2
Minimum creep rate [%/h]
10 -1
1
Fig. 364. Time to rupture vs. minimum creep rate for
21Cr-32Ni-Ti-Al steel plates. n indicates the total number
of data points.
2.3.4.10.4 Reference
[1] National Research Institute for Metals: NRIM Creep Data Sheet, 27B (2000).
Landolt-Börnstein
New Series VIII/2B
2.4.11 22Cr-15Ni-NbBN steel
273
2.4.11 22Cr-15Ni-NbBN steel
2.4.11.1 Introduction
22Cr-15Ni-NbBN (TEMPALOY A-3) is a steel with corrosion resistance equal to, and high temperature
strength greater than that of Incoloy 800H, and has superior economic efficiency. It is suited for use in
environments of 700 °C or higher. Other high Cr steels containing more than 20 % Cr include SUS309
and SUS310, but their high temperature strength is low. Its high temperature strength is improved by
Nb(C,N) and M23C6 formed by adding Nb and N to 22 %Cr-15 %Ni.
SUS310 and other high Cr steels become brittle as large quantities of the σ-phase are precipitated
during long use. The σ-phase precipitation of 22Cr-15Ni-NbBN is equal to that of 18 %Cr-8 %Ni type
steel, providing it with suitable toughness after long-term use. Its Ni content lower than that of SUS310
and Incoloy 800H provides superior economic benefits.
2.4.11.2 Chemical composition
Chemical requirement of 22Cr-15Ni-NbBN steel is shown in Table 157.
Table 157. Chemical requirement of 22Cr-15Ni-NbBN steel.
Designation
Chemical composition [wt%]
C
Si
Mn
P
S
Ni
Cr
KA-SUS
0.03- ≤1.00 ≤2.00 ≤0.040 ≤0.030 14.50- 21.00309J4HTB 0.10
16.50 23.00
Nb
0.500.80
B
N
≤0.005 0.100.20
2.4.11.3 Mechanical requirements
The room temperature mechanical requirements of this steel are as follows: tensile strength is more than
590 MPa, yield strength is more than 235 MPa, elongation is more than 35 % in the case of outer
diameter of tubes over 20 mm.
2.4.11.4 Creep rupture properties
The creep rupture data obtained from NKK co. [1] is shown in Fig. 365.
Landolt-Börnstein
New Series VIII/2B
274
2.4 Austenitic stainless steels
500
400
300
Stress (MPa)
200
100
90
80
70
60
50
600℃
650℃
700℃
750℃
800℃
40
30
20
101
102
103
104
105
Fig. 365.
Creep rupture
strength data of 22Cr-15NiNbBN steel; [1].
Time to rupture (h)
2.4.11.5 Reference
[1] NKK Technical Bullitin: TEC. No. 243-312 Boiler Tubing.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 278]
2.4.12 23Cr-18Ni-3Cu-1.5W-Nb-N steel
275
2.4.12 23Cr-18Ni-3Cu-1.5W-Nb-N steel
2.4.12.1 Introduction
23Cr-18Ni-3Cu-1.5W-0.4Nb-0.2N steel (SAVE25) has been used for superheater and reheater tubings in
power boilers in field exposure test experiences since June 1999, and has been approved as a ministerial
ordinance material by METI Japan. SAVE25 was developed by Sumitomo Metal Industries, Ltd., Japan
in the 1990s to improve creep rupture strength and to obtain a high corrosion resistance by optimizing
chemical composition as well as to achieve relatively low cost. The chromium content is 23 %, based on
requirements for hot corrosion and steam oxidation resistance. Nitrogen, as a substitute element for
nickel, is sufficient at 0.2 % for stabilizing an austenite structure. Copper and tungsten are characteristic
alloying elements to improve creep strength.
2.4.12.2 Material standards, chemical and tensile requirements
Table 158 and Table 159 respectively show the chemical and tensile requirements of 23Cr-18Ni-3Cu1.5W-Nb-N steel tubes approved by METI in the ministerial ordinance interpretation designated as KASUS310J3TB. Fig. 366 shows charpy impact values at 0 °C of SAVE25 steel aged for 3000 h.
Table 158. Chemical requirements of 23Cr-18Ni-3Cu-1.5W-Nb-N steel (SAVE25)
Stan- Designation
dard
METI KA-SUS310J3TB
Chemical composition [wt%]
C
Si
Mn
P
S
Ni
Cr
W
Cu
Nb
N
0.05∼ ≤1.50 ≤2.00 ≤0.030 ≤0.010 15.00~ 21.00~ 0.80~ 2.00~ 0.30~ 0.15~
22.00 24.00 2.80 4.00 0.60 0.30
0.12
Table 159. Tensile requirements of 23Cr-18Ni-3Cu-1.5W-Nb-N steel (SAVE25)
Standard Designation
Yield strength [MPa] Tensile strength [MPa] Elongation [%]
METI
KA-SUS310J3TB ≥650
2295
≤30
120
Charpy impact value (J/cm2)
600℃
650℃
100
700℃
750℃
80
60
40
20
0
500
1000
1500
2000
2500
3000
Fig. 366. Charpy impact values at 0 °C of the
SAVE25 steel aged for up to 3000 h; [1].
Aging time (h)
2.4.12.3 Creep properties
2.4.12.3.1 Creep rupture data and creep rupture strength
The chemical compositions of SAVE25 steel tubes used for creep rupture testing are listed in Table 160.
These tubes were melted in a VIF forged from 180 kg ingots, hot-extruded, cold-drawn and solution heattreated at 1150 - 1175 °C .
Landolt-Börnstein
New Series VIII/2B
276
2.4 Austenitic stainless steels
The creep rupture properties of SAVE25 steel are shown in Fig. 367. The creep rupture master-curve of
the Larson-Miller parameter method is shown in Fig. 368.
Table 160. Chemical compositions of SAVE25 steel tested [1].
Chemical composition [wt%]
Product
C
Si
Mn Ni Cr W
Cu Nb N
Tube A 0.10 0.13 0.48 19.8 22.6 1.47 3.51 0.44 0.22
Tube B 0.10 0.14 0.48 16.7 22.6 2.51 3.50 0.45 0.18
Tube C 0.10 0.14 0.49 15.3 22.7 1.48 3.52 0.45 0.20
500
Stress (MPa)
300
200
650℃
700℃
100
750℃
70
50
Fig. 367.
Creep rupture
strength data of SAVE25 steel;
[1].
800℃
100
101
102
103
104
105
Time to rupture (h)
500
650℃
700℃
300
750℃
Stress (MPa)
800℃
200
100
Fig. 368. Creep rupture master-curve of
SAVE25 steel by Larson-Miller parametric
method; [1].
70
50
16
18
20
22
24
26
T (17.9817＋log t) × 10- 3
2.4.12.3.2 Creep deformation behavior
Fig. 369 compares the creep curves of 0.75Mo and 1.5W steels having base composition of 23Cr-20Ni3Cu-0.12N with the same Mo equivalents. It is found that the tertiary creep of the 1.5W steel is retarded,
due to the fact that σ phase precipitation and carbide/nitride coarsening are suppressed, thus resulting in
higher creep strength.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 278]
2.4.12 23Cr-18Ni-3Cu-1.5W-Nb-N steel
277
1.5
Elongation (%)
800℃，78.4MPa
1.0
0.75Mo
0.5
1.5W
Fig. 369. Effect of 0.75% Mo and 1.5% W on
creep curves of 23Cr-20Ni-3Cu-0.12N steels;
[1].
0
0
200
400
Time (h)
600
800
2.4.12.3.3 Effect of chemical compositions on creep strength
Fig. 370 shows that the creep rupture strength at 700 °C and 156.8 MPa increases remarkably with the
addition of 2.5 to 3.5 % Cu.
2400
Cu solubility
Time to rupture (h)
2200
2000
1800
1600
1400
700℃ 156.8MPa
1200
0
1
2
Cu content (mass %)
3
4
Fig. 370. Effect of Cu content on creep
rupture strength of 23Cr-15.5Ni-2.5W-0.2N
steels; [2].
Fig. 371 shows that the creep rupture strength of 1 % Mo steel is much lower than that of 1.5 % W steel,
although the Mo-equivalent (=Mo + 1/2W) of 1 % Mo steel is higher than that of 1.5 % W steel.
Chemical analysis of extracted residues in the 1 % Mo steel revealed that about 90 % of the alloyed Mo is
dissolved even after long-term aging at 650 - 800 °C . Most of the alloyed Mo seems to be still in solution
during the long-term creep stage and contributes to creep rupture strength by solid solution strengthening.
Time to rupture (h)
3000
650℃ 235.2MPa
700℃ 156.8MPa
1Mo
1Mo
750℃ 98MPa
2000
1000
700
500
Landolt-Börnstein
New Series VIII/2B
1.5W
1.5W
1Mo
1.5W
Fig. 371. Effect of Mo and W on creep
rupture strength of 23Cr-20Ni-3Cu-Mo-W0.12N steels; [2].
278
2.4 Austenitic stainless steels
2.4.12.4 Estimated creep strength
Allowable stresses of SAVE25 developed based on the ASME criteria are compared with those of HR3C
steel (ASME SA-213 TP310HCbN) and TP347H steel in Fig. 372. The allowable stress of SAVE25 at
700 °C is 60 MPa, which is 36 % higher than that of HR3C steel and 88 % higher than that of TP347H.
180
Allowable tensile stress (MPa)
160
SAVE25
140
HR3C
120
100
TP347H
80
60
Fig. 372. Allowable stresses of SAVE25 steel
compared with HR3C steel and TP347H steel;
[1].
40
20
0
200
400
Temperature (℃)
600
800
2.4.12.5 References
[1] Semba, H., Igarashi, M., Sawaragi, Y., and Iseda, A.: Proc. International Conf. Advanced Materials
and Processes, Munich, Germany, (2001), 1.
[2] Semba, H., Igarashi, M., and Sawaragi, Y.: Proc. International Conf. Power Engineering-97, Vol.2,
Tokyo, (1997), 125.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 282]
2.4.13 23Cr-45Ni-6W-Nb-Ti-B steel
279
2.4.13 23Cr-45Ni-6W-Nb-Ti-B steel
2.4.13.1 Introduction
23Cr-45Ni-6W-Nb-Ti-B steel (HR6W) was developed by Sumitomo Metal Industries, Ltd., Japan in the
1980s for heat exchanger tubings such as power boiler superheaters and reheaters, steam generators and
chemical plants operated at very high temperatures of around 700 °C. HR6W steel has a high creep
rupture strength compared with conventional high alloy steels such as Hastelloy X, Alloy 800H and
Incoloy 807 due to the solution strengthening effect of W and N and the precipitation strengthening effect
provided by a fine Laves phase, M23C6 and nitrides. The corrosion resistance properties of HR6W steel,
such as steam oxidation resistance and hot corrosion resistance are superior to those of 18Cr-8Ni stainless
steels due to the higher Cr content.
2.4.13.2 Material standards, chemical and tensile requirements [1]
HR6W steel is not specified in any codes or standards except the internal specifications of Sumitomo
Metal Industries, Ltd.
Table 161 shows the chemical requirements for HR6W steel by Sumitomo Metal Industries, Ltd.
Table 2 lists the chemical composition of HR6W steel used for creep rupture and characterization tests, as
well as Hastelloy X, Alloy 800H and Incoloy 807 for comparison. Fig. 373 shows Charpy impact values
at 0 °C for HR6W steel aged for 3,000 h, as compared to the competitive alloys.
Table 161. Chemical requirements of HR6W steel
Chemical composition [wt%]
Product
C
Si
Mo
P
S
Ni
Cr
W
Ti
Nb
Tube
≤0.10 ≤1.00 ≤2.00 ≤0.030 ≤0.030 35.0-45.0 21.0-25.0 4.0-8.0 ≤0.20 ≤0.40
Table 162. Chemical compositions of alloys tested [1]
Chemical composition [wt%]
Alloy
C
Si
Mn Ni
Cr
Mo V
Ti
Nb
M3
0.08 0.42 1.20 42.21 22.91 2.98 0.07 0.19
M5
0.07 0.40 1.21 44.20 22.77 4.91 0.08 0.18
W5
0.08 0.42 1.16 41.69 22.64 5.07 0.07 0.18
W7
0.08 0.39 1.20 44.15 22.76 7.38 0.09 0.16
Hastelloy X 0.07 0.38 0.70 44.80 22.47 8.71 1.10 Alloy 800H 0.08 0.43 0.97 33.86 21.08 0.53 Incoloy 807 0.07 0.43 0.97 38.82 20.41 4.55 0.30 0.95
Landolt-Börnstein
New Series VIII/2B
Others
0.0034 B
0.0032 B
0.0033 B
0.0030 B
1.1 Co
0.43 Al
0.38 Al, 7.7 Co
280
2.4 Austenitic stainless steels
160
120
100
80
Impact value after aging for 3000h
Impact value before aging
140
Charpy impact value (J/cm2)
0.6
M3
M5
W5
W7
Hastelloy X
Alloy 800H
Incoloy 807
60
40
0.4
0.2
20
0
700
750
0
800
700
Aging temperature (℃）
750
800
Aging temperature (℃）
Fig. 373. Charpy impact properties of test alloys at 0 °C after 3000 h aging; [1].
2.4.13.3 Creep properties
2.4.13.3.1 Creep rupture data and creep rupture strength [1]
Fig. 374 shows the creep rupture properties for HR6W steel. The 100,000 h extrapolated creep rupture
strength at 700 °C is estimated to be over 90 MPa.
300
Stress (MPa)
200
700℃
100
750℃
70
50
800℃
102
103
104
105
Fig. 374.
Creep rupture
strength data of HR6W steel;
[1].
Time to rupture (h)
2.4.13.3.2 Effect of chemical compositions on creep properties [1]
The creep rupture strength of four alloys (M3, M5, W5, W7) is superior to that of Alloy 800H, and the
rupture strength levels of the alloys containing W (W5 and W7) are higher than those of the alloys
containing Mo (M3 and M5), as shown in Fig. 375.
The creep rupture ductilities for alloy M5 with 5 % Mo and alloy W7 with 7 % W are shown in Fig.
376. The rupture elongation and the reduction of area deteriorate with increased testing temperature and
Landolt-Börnstein
New Series VIII/2B
Ref. p. 282]
2.4.13 23Cr-45Ni-6W-Nb-Ti-B steel
281
time in M5, with inferior creep rupture strength. In W7, having excellent rupture strength, the
deterioration in creep rupture ductility is smaller than that in M5; in particular the change of the reduction
of area in W7 is small even after higher temperature and longer term testing. Fig. 377 shows that lowering
of the W content to 3 % results in decreased creep rupture strength.
19
200
T (18 + log t ) [ 10 3 ]
20
21
22
200
23
5
700 °C ×10 h
700 °C
100
W [%]
W5 5.07
W7 7.38
750 °C
Stress [MPa]
Stress [MPa]
100
50
200
M3
M5
Hasteloy X
50
200
5
700 °C ×10 h
700 °C
100
100
Mo [%]
M3 2.98
M5 4.91
750 °C
50
W5
W7
Alloy 800 H
Incoloy 807
50
10 2
10 3
10 4
Time to rupture [h]
10 5
21
22
23
24
T (20 + log t ) [ 10 3 ]
25
Fig. 375. Creep rupture properties for test alloys M3, M5, W5, and W7; [1].
Alloy
M5
700℃ 750℃
E1.
（4.91% Mo）R.A
W7
Elongation, Reduction of area (%)
E1.
（7.38% W） R.A
100
80
60
40
Fig. 376.
Comparison of creep rupture
ductility between alloys M5 (4.91% Mo) and
W7 (7.38% W); [1].
20
0
102
103
104
Time to rupture (h）
Table 163. Chemical compositions of Alloys A, B and C given in Fig. 377.
Chemical composition [wt%]
Alloy
C
Ni
Cr
W Ti
Nb B
A
0.083 42.75 23.12 6.58 0.07 0.16 0.0042
B
0.080 40.60 23.08 5.03 0.07 0.15 0.0038
C
0.085 35.56 22.98 3.10 0.07 0.15 0.0035
Landolt-Börnstein
New Series VIII/2B
282
2.4 Austenitic stainless steels
Alloy 700℃ 750℃
A
B
C
Stress (MPa)
300
200
100
Fig. 377. Effect of W content on creep
rupture strength of 23Cr-45Ni-Nb-Ti-B
steels; [1].
70
50
10
102
103
104
105
Time to rupture (h)
2.4.13.4 Reference
[1] Sawaragi, Y., Hayase, Y., and Yoshikawa, K.: Proc. International Conf. Stainless Steels, Chiba,
Japan (1991), 633.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 285]
2.4.14 25Cr-12Ni steel
283
2.4.14 25Cr-12Ni steel
2.4.14.1 Introduction
25Cr-12Ni steel is a cast heat resistant stainless steel. It is basically austenitic, but in some composition
balance of Cr, Ni and C, partially it is duplex of ferritic and austenitic. The partially ferritic alloy is
adapted to operating conditions that are subjected to changes in temperature level and applied stress. The
wholly austenitic alloy is used extensively in high temperature applications because of its combination of
relatively high strength and oxidation resistance at temperatures up to 1100 ºC.
2.4.14.2 Chemical composition
Chemical requirements of 25Cr-12Ni steel are shown in Table 164.
Table 164. Chemical requirements of 25Cr-12Ni steel
Chemical composition [wt%]
StanDesigdards
nation
C
Si
Mn
P
0.20JIS
SCH13
≤2.00 ≤2.00 ≤0.040
0.50
0.20ASTM HH
≤2.00 ≤2.00 ≤0.04
0.50
0.20ASTM Type I
≤1.75 ≤2.50 ≤0.05
0.45
0.20ASTM Type II
≤1.75 ≤2.50 ≤0.05
0.45
Type I: partially feritic
Type II: wholly austenitic
S
Ni
11.00≤0.040
14.00
11.00≤0.04
14.00
10.00≤0.05
14.00
10.00≤0.05
14.00
Cr
24.0028.00
24.0028.00
23.0028.00
23.0028.00
Mo
N
≤0.50
≤0.20
≤0.20
2.4.14.3 Mechanical properties
The mechanical requirements of this steel is shown in Table 165.
Table 165. Mechanical property requirements of 25Cr-12Ni steel.
Standards Designation Yield strength Tensile strength
[N/mm2]
[N/mm2]
JIS
SCH13
≥235
≥490
ASTM
HH
≥240
≥535
ASTM
Type I, II
≥550
2.4.14.4 Creep rupture properties
The creep rupture data of 25Cr-12Ni steel obtained from [1] is shown in Fig. 378. From [1] data on
elongation and reduction of area can also be obtained (Fig. 379 and Fig. 380).
Landolt-Börnstein
New Series VIII/2B
284
2.4 Austenitic stainless steels
Stress (MPa)
200
700℃
800℃
900℃
950℃
100
90
80
70
60
50
40
30
20
10
101
102
103
104
105
Fig. 378.
Creep rupture
strength data of 25Cr-12Ni
steel. n indicates the total
number of data points.
Time to rupture (h)
n = 88
Fig. 379, see next page
n = 20
100
900 °C
80
80
60
60
40
40
20
0
100
n = 15
950 °C
80
Reduction of area [%]
Elongation [%]
100
100
950 °C
80
40
40
20
20
10 6
n = 15
0
60
10 5
10 2 10 3 10 4
Time to rupture [h]
900 °C
20
60
0
10
n = 20
0
10
Fig. 380.
Elongation and
reduction of area at 900 and
950 ºC for 25Cr-12Ni steel.
n indicates the total number of
data points in each diagram.
10 5
10 2 10 3 10 4
Time to rupture [h]
10 6
Landolt-Börnstein
New Series VIII/2B
Ref. p. 285]
100
n = 29
100
700 °C
80
80
60
60
40
40
20
20
0
0
100
n=5
100
750 °C
80
80
60
60
40
40
20
0
100
n = 27
Reduction of area [%]
Elongation [%]
2.4.14 25Cr-12Ni steel
80
40
40
20
20
0
0
80
100
850 °C
60
40
40
20
20
10 5
10 2 10 3 10 4
Time to rupture [h]
10 6
750 °C
n = 27
800 °C
n=5
850 °C
80
60
0
10
n=5
80
60
n=5
700 °C
0
60
100
n = 29
20
100
800 °C
285
0
10
10 5
10 2 10 3 10 4
Time to rupture [h]
10 6
Fig. 379.
Elongation and
reduction of area at 700, 750,
800 and 850 ºC for 25Cr-12Ni
steel. n indicates the total number
of data points in each diagram.
2.4.14.5 Reference
[1] National Research Institute for Metals: NRIM Creep Data Sheet, 37A (1992).
Landolt-Börnstein
New Series VIII/2B
286
2.4 Austenitic stainless steels
2.4.15 25Cr-20Ni steel
2.4.15.1 Introduction
25Cr-25Ni centrifugally cast steel is a heat resistant and high corrosion resistant alloy having similar
chemical composition as type 310 stainless steel but a rather high carbon content. It is popularly used in
chemical plant pipings or reactors, especially in steam or ethylene reformer tubes.
2.4.15.2 Materials standards, and chemical requirements
Table 166. Chemical requirements of 25Cr-20Ni steel castings; JIS SCH21, ASTM A351 HK40, ACI
HK40, BS 310 310C45 and DIN 17465 1.4848
Chemical composition [wt%]
Specification
C
Cr
Mn
Mo Ni
P
S
Si
Others
N: ≤0.2
JIS SCH21
0.35-0.45 23-27 ≤1.5
0.04
≤0.5 19-22 0.04
≤1.75
Bal: Fe
ASTM A351 HK40 0.35-0.45 23-27 1.5
19-22 0.04
0.04
1.75
Bal: Fe
Bal: Fe
ACI HK40
0.2-0.6
24-28 ≤2
≤0.5 18-22 ≤0.04 ≤0.04 1.75
22-27 ≤1.5
0.04
BS 310 310C45
≤0.5
≤0.5 19-22 0.04
≤1.75 Bal: Fe
DIN 17465 1.4848 0.3-0.5
24-26 0.5-1.5
19-21 0.045 0.03
1-2.5 Bal: Fe
2.4.15.3 Creep properties of 25Cr-20Ni steel castings
Information of fact on creep data for 25Cr-20Ni steel castings can be obtained from [1].
2.4.15.3.1 Creep rupture data of 25Cr-20Ni steel castings
The creep rupture strength of 25Cr-20Ni steel castings obtained from the available creep data source is
illustrated in Fig. 381. The results of creep tests from 14 JIS SCH21 heat cast tubes are given in [1]. Data
on elongation, reduction of area, minimum creep rate and microstructures of as-received and crept
specimens can also be obtained from [1].
Log-log plots of relations between stress and time to rupture in the temperature range 800 - 1000 °C are
almost of a linear shape but tend to be slightly inverse sigmoidal.
2.4.15.3.2 Creep rupture strength of 25Cr-20Ni steel castings
The creep rupture time at each stress level is widely scattered, and almost a 0.8 - 1.0 order of magnitude
difference can be found (Fig. 381). There are no significant differences of the data scattering tendency
between the test temperatures. The creep rupture strength depends on manufacturing conditions, chemical
composition, and initial microstructure. In particular, the rupture strength of castings is influenced by
casting condition and initial (as-cast) microstructure.
The master rupture curve obtained by the Larson-Miller parameter method for centrifugally cast 25Cr20Ni steel tubes is illustrated in Fig. 382. A set of of Larson-Miller parameters fitted to all creep-rupture
data in [1] is given in Table 167.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 291]
2.4.15 25Cr-20Ni steel
287
100
８００℃
９００℃
１０００℃
10
n=211
10
1
10 2
10 3
10 4
10 5
10 6
Fig. 381.
Creep rupture
strength data of centrifugally
cast 25Cr-20Ni steel tubes.
n indicates the total number of
data points.
Time to rupture (h)
200
800 °C
870 °C
900 °C
982 °C
1000 °C
1050 °C
1100 °C
Stress [MPa]
100
80
60
50
40
30
20
10
8
6
4
12
average curve
n = 263
18
20
16
14
Larson-Miller-parameter TK ( log t R-10.315) [10 3 ]
Fig. 382. Master rupture curve obtained by the LarsonMiller parameter method for centrifugally cast 25Cr-20Ni
steel tubes; [1]. n indicates the total number of data points.
Table 167. A set of Larson-Miller parameters fitted to all creep-rupture data in [1] according to
log tR=(T+273.15)−1[b0+b1log S+b2 (log S)2]−C
n C
b0
b1
b2
263 1.031462 × 10 2.204935 × 104 −3.439566 × 103 −4.937139 × 102
2.4.15.3.3 Elongation and reduction of area of ruptured specimens
Elongation and reduction of area of ruptured specimens tested at 800, 900 and 1000 °C as a function of
time to rupture are shown in Fig. 383. The ductility of the specimens tested at 800 °C and 900 °C is
almost constant and not high, approximately some percent. The ductility of the specimens tested at
1000 °C up to 104 h has a similar tendency as those tested at lower temperatures, however, it increases
with increasing time for rupture times longer than 103 h.
Landolt-Börnstein
New Series VIII/2B
288
2.4 Austenitic stainless steels
100
100
800 °C
80
80
60
60
40
40
20
20
0
0
n = 70
100
900 °C
Reduction of area [%]
100
Elongation [%]
n = 72
80
60
40
20
800 °C
n = 70
900 °C
n = 71
1000 °C
80
60
40
20
0
0
100
n = 72
n = 71
1000 °C
100
80
80
60
60
40
40
20
20
0
10
10 5
10 2 10 3 10 4
Time to rupture [h]
10 6
0
10
10 5
10 2 10 3 10 4
Time to rupture [h]
10 6
Fig. 383.
Elongation and
reduction of area of ruptured
specimens tested at 800, 900 and
1000 °C as a function of time to
rupture. n indicates the total
number of data points in each
diagram.
2.4.15.3.4 Microstructural change
The microstructure of centrifugally cast 25Cr-20Ni steel tubes varies with cast condition. Fig. 384 is an
example set of variation in as-cast microstructure from [1]. Large directed grain structures (TAA and
TAG) or rather fine equiaxed grain structure (TAD and TAF) can be found. As reported in [2], there is no
clear relationship between grain size and creep rupture strength in this alloy.
The microstructure of 25Cr-20Ni steels changes during creep deformation (Fig. 385: All the data from
cast TAA of Fig. 384). The microstructure of as-cast materials is an austenitic matrix having dendritic
structure with primary carbides along dendrite interfaces. With increasing creep temperature and/or creep
time, carbide coarsening and coalescencing takes place. Under the creep conditions tested below 900 °C,
fine carbides precipitate around the final solidified zone during creep tests.
There are some reports that σ phase precipitates are found during creep deformation. The σ phase
morphology and quantity very depend on the alloy chemistry. In general, massive σ phase, which
precipitates after long creep (more than 30000 h) beside coarsened carbide, can be a creep void initiation
site. Details are described later.
2.4.15.3.5 Creep deformation process in 25Cr-20Ni steel
As seen in Fig. 385, voids at/along dendrite interfaces can be found at low magnified structures. At a
higher magnification observations reveals that voids initiate within the primary carbides. An estimated
process of creep damage progress is as follows:
A. void initiation at dendrite interface, especially at primary carbides
B. connection of adjoining voids
C. microcrack initiation
D. crack propagation
Landolt-Börnstein
New Series VIII/2B
Ref. p. 291]
2.4.15 25Cr-20Ni steel
289
There are reports on the relationship between creep strength and σ phase precipitation [3]. In general, a
σ/matrix interface tends to be a creep crack initiation and propagation site. Time to rupture of alloys
involving more σ phase are short. Therefore, improved alloy design avoids σ phase precipitation. For
example, PHACOMP alloy design method was applied and improved alloys have been developed [4].
800 °C
900 °C
1000 °C
Fig. 385. A comparison of microstructure after creep ruptured at similar times as a function of test temperature.
Landolt-Börnstein
New Series VIII/2B
290
2.4 Austenitic stainless steels
Fig. 384. Variation in microstructure of centrifugally cast 25Cr-20Ni steel.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 291]
2.4.15 25Cr-20Ni steel
291
2.4.15.4 References
[1] National Research Institute for Metals Japan: NRIM Creep Data Sheet, 16B (1990).
[2] Ohta, S., and Kori, M.: Report of the 123rd Committee on Heat-Resisting Metals and Alloys, Japan
Society for the Promotion of Science, 20, No.1 (1984) 97 in Japanese.
[3] Ohta, S., and Kori, M.: Annmonia to Kougyou, 31 (1978) 1 in Japanese.
[4] Kihara, S., Ohtomo, A., Shinozaki, S., Saiga, Y., and Kawasumi, Y: J. Jpn. Petrol. Inst. 24 (1981)
27.
Landolt-Börnstein
New Series VIII/2B
292
2.4 Austenitic stainless steels
2.4.16 25Cr-20Ni-Nb-N steel
2.4.16.1 Introduction
25Cr-20Ni-Nb-N austenitic steel (TP310HCbN; HR3C) is used as superheater and reheater tubes in
fossile fired, black liquor recovery and refuse fired boilers. The steel has been developed for improving
310 type stainless steel by the addition of niobium and nitrogen to increase creep strength at elevated
temperatures. The microstructure consists of an austenitic matrix strengthened by M23C6 carbide mainly
along grain boundaries and finely dispersed NbCrN nitride in the matrix. NbCrN is fine and stable even
after long term creep exposure at high temperatures. No σ phase has been found after more than 30000 h
in the temperature range between 600 and 800 °C, which is achieved by optimizing the Ni-balance.
2.4.16.2 Material standards, and chemical and tensile requirements
Tables 168 and 169 give the chemical requirements and the corresponding tensile requirements of 25Cr20Ni-Nb-N steel tubes which are designated by the standards; Japanese METI KA-SUS310J1TB, ASTM
A213 TP310HCbN (ASME Sec. I CC 2115), and EN 10216-5.
Table 168. Chemical requirements of 25Cr-20Ni-Nb-N steel tubes; Japanese METI KA-SUS310J1TB,
ASTM A213 TP310HCbN (ASME Sec. I CC 2115), and EN 10216-5.
Chemical composition [wt%]
DesigGrade
Std.No.
nation
C
Si
Mn P
S
Ni
Cr
Nb
N
Japanese
17.0
23.0
0.20 0.15
KA≤
≤
≤
≤
≤
SUS310J1TB 0.10 1.50 2.00 0.030 0.030 23.0
METI
27.0
0.60 0.35
0.04 ≤
17.00 24.00 0.20 0.15
≤
≤
≤
ASTM
TP310HCbN
A213
0.10 0.75 2.00 0.030 0.030 23.00 26.00 0.60 0.35
17.0
23.0
0.20 0.15 10216≤
≤
≤
≤
≤
EN
TP310HCbN
0.10 1.50 2.00 0.030 0.030 23.0
27.0
0.60 0.35 5
Table 169. Tensile requirements of 25Cr-20Ni-Nb-N steel tubes; Japanese METI KA-SUS310J1TB,
ASTM A213 TP310HCbN (ASME Sec. I CC 2115), and EN 10216-5.
Min.
Standard No.
Designation Grade
Min. TS1)
Min. YS2)
elongation
Japanese
KA660 MPa
295 MPa
30 %
METI
SUS310J1TB
ASTM
TP310HCbN
665 MPa
295 MPa
30 %
SA213-01
EN
HR3C
665 MPa
295 MPa
30 %
10216-5
1) TS; tensile strength, 2) YS; yield strength as 0.2% proof stress
2.4.16.3 Tensile properties of 25Cr-20Ni-Nb-N steel tubes
2.4.16.3.1 Tensile properties of 25Cr-20Ni-Nb-N steel tubes
Fig. 386 shows tensile strength and yield stress data of 25Cr-20Ni-Nb-N steel tubes [1], [2] and [5]. They
are higher than those of TP310S steel at all temperatures up to 750 °C. The corresponding tensile
elongation and reduction of area data of 25Cr-20Ni-Nb-N steel tubes, which are also available in [1], [2]
and [5], are at the same level as those of TP310S steel at temperatures below 300 °C but much better
above 300 °C.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 296]
2.4.16 25Cr-20Ni-Nb-N steel
293
1000
Tensile strength,Yield stress [MPa]
900
800
700
600
500
400
300
200
100
0
100
200
300 400 500
Temperature [°C]
600
700
800
Fig. 386. Tensile strength (circles) and yield stress
(triangles) data of 25Cr-20Ni-Nb-N steel tubes.
2.4.16.4 Creep rupture properties of 25Cr-20Ni-Nb-N steel tubes
2.4.16.4.1 Creep rupture data of 25Cr-20Ni-Nb-N steel tubes
Fig. 387 shows creep rupture data of 25Cr-20Ni-Nb-N steel tubes with average curves assessed by the
Larson-Miller parameter method [1]. The longest creep rupture datum of 25Cr-20Ni-Nb-N steel tubes is
about 90000 h at 700 °C. Their long-term creep strength is very stable and no degradation in creep
strength is expected at temperatures up to 750 °C. Fig. 388 shows a Larson-Miller parameter plot of the
creep rupture data of 25Cr-20Ni-Nb-N steel tubes with a master rupture curve and a 95 % confidence
lower limit. The best fitting was achieved with an optimized constant of 16.96.
500
400
300
Stress [MPa]
600 °C
200
650 °C
100
80
60
40
700 °C
600 °C
650 °C
700 °C
750 °C
average curve
1
Landolt-Börnstein
New Series VIII/2B
10
750 °C
10 2
10 3
Rupture time [h]
10 4
10 5
Fig. 387. Creep rupture data of
25Cr-20Ni-Nb-N steel tubes.
294
2.4 Austenitic stainless steels
500
600 °C
×10 5h
400
Stress [MPa]
300
650 °C
×10 5h
200
700 °C
×10 5h
100
600 °C
650 °C
700 °C
750 °C
80
60
40
14
average curve
minimum curve
Fig. 388. Larson-Miller parameter plot of the creep
rupture data of 25Cr-20Ni-Nb-N steel tubes.
15 16 17 18 19 20 21 22 23 24
Larson-Miller-parameter T (16.96 + log t )[×10 -3 ]
2.4.16.4.2 Creep data of 25Cr-20Ni-Nb-N steel tubes
Fig. 389 shows minimum creep rate data of 25Cr-20Ni-Nb-N steel tubes measured at various stress levels
at temperatures between 600 °C and 750 °C with fitted curves obtained by the Larson-Miller parameter
method shown in Fig. 390 [1]. Fig. 390 shows a Larson-Miller parameter plot of the minimum creep rate
data of 25Cr-20Ni-Nb-N steel tubes with a master curve. The best fitting was achieved with an optimized
constant of 19.06.
500
400
Stress [MPa]
300
200 600 °C
650 °C
100
600 °C
650 °C
700 °C
750 °C
average curve
80 700 °C
60
750 °C
40
10 -2
10 -1
1
10
Minimum creep rate [% / 10 3 h]
10 2
10 3
Fig. 389. Minimum creep rate
data of 25Cr-20Ni-Nb-N steel
tubes.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 296]
2.4.16 25Cr-20Ni-Nb-N steel
295
500
400
600 °C
0.01%/10 3h
Stress [MPa]
300
650 °C
0.01%/10 3h
200
700 °C
0.01%/10 3h
100
600 °C
650 °C
700 °C
750 °C
80
60
average curve
40
13
14 15 16 17 18 19 20 21 22 23
Larson-Miller-parameter T (19.06 + log e ) [×10 -3 ]
Fig. 390. Larson-Miller parameter plot of the minimum
creep rate data of 25Cr-20Ni-Nb-N steel tubes.
2.4.16.5 Allowable stress of 25Cr-20Ni-Nb-N steel tubes
Fig. 391 shows the allowable stress determined for 25Cr-20Ni-Nb-N steel (Japanese METI KASUS310HCbN) according to the METI standard procedure comparing with that for the conventional steel
ASME SA213-TP347HTB (Japanese METI KA-SUSTP347HTB).
180
KA-SUS310J1TB
Allowable tensile stress (MPa)
160
140
120
100
KA-SUSTP347HTB
80
60
40
Fig. 391. Allowable tensile stress determined for 25Cr20Ni-Nb-N steel tubes according to the Japanese METI
standard.
20
0
0
100
200
300
400
500
600
700
800
Temperature ( ℃)
2.4.16.6 Microstructural change of 25Cr-20Ni-Nb-N steel tubes
The microstructural change in 25Cr-20Ni-Nb-N steel tubes after aging for up to 10000 h in the
temperature range between 600 °C and 750 °C is available in [2] and [4]. There is no significant
microstructural change observed even after aging for 10000 h at 750 °C. A detailed TEM observation of
the specimens aged for 3000 h in the temperature range between 600 °C and 750 °C has shown that fine
NbCrN nitride dispersion is identified in the matrix and no harmful blocky precipitation such as σ phase
is found.
Landolt-Börnstein
New Series VIII/2B
296
2.4 Austenitic stainless steels
2.4.16.7 Performance of service exposed tubes
Performance of service exposed tubes is available in [5].
2.4.16.8 References
[1] Sumitomo seamless tubes and pipe Creep Data Sheets, Sumitomo Metal Industries, (1993).
[2] Sawaragi, Y., Teranishi, H., Makiura, H., Miura, M., and Kubota, M.: Sumitomo Metals, 37 (1985),
66.
[3] Sawaragi, Y.: The Journal of Japan Welding Society, 58 (1989), 193.
[4] Sawaragi, Y., Teranishi, Y., Iseda, A., Yoshikawa, K.: Sumitomo Metals, 42 (1990), 260.
[5] Natori, A., Sawaragi, Y.: Sumitomo Metals, 45 (1993), 96.
[6] Sawaragi, Y., Teranishi, H., and Yoshikawa, K.: Proc. of Int. Conf. on Creep, Tokyo, (1986), 239.
[7] Yoshikawa, K., Sawaragi, Y., and Yuzawa, H.: Proc. of Int. Conf. on Improved Coal-Fired Power
Plants, EPRI, Palo Alto, (1986), 5.
[8] Sawaragi, Y., Teranishi, H., Iseda, A., and Yoshikawa, K.: The Sumitomo Search, No.44 (1990),
146.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 300]
2.4.17 25Cr-35Ni steel
297
2.4.17 25Cr-35Ni steel
2.4.17.1 Introduction
This material is used as centrifugally cast tubes and cast blocks of 25Cr-35Ni-0.4C steel for reformer
furnaces. It is a heat resisting steel casting and is manufactured by the centrifugal casting process or the
mould casting process. Heat treatment is not performed, but this steel is used as-casted.
2.4.17.2 Material standards, chemical compositions and tensile properties
The following information is obtained from [1].
Table 170 shows the specification (SCH 24(JIS G 5122)) of the chemical composition of 25Cr-35Ni0.4C steel and analysis results for the chemical compositions of three typical heats used for data treatment
in this article. Three heats were made by the centrifugal casting process and one heat was made by the
mould casting process. The chemical compositions are almost equivalent and there is little difference
between the heats and the casting processes.
Table 170. Specification of chemical composition and analysis results for 25Cr-35Ni-0.4C steel
SCH24(JIS G 5122) and HP(ASTM).
Chemical composition [wt%]
Product
C
Si
Mn
P
S
Ni
Cr
Fe
Requirement
0.35-0.75 ≤2.00 ≤2.00 ≤0.040 ≤0.040 33.0-37.0 24.0-28.0 Rem.
Cast tube (typical
Heat 1 0.48
0.83 1.01 0.009 0.005 34.47
25.64
Rem.
example)
Heat 2 0.48
0.97 0.81 0.014 0.009 35.63
27.49
Rem.
Heat 3 0.42
0.82 1.02 0.01
0.018 34.4
24.75
Rem.
Cast block
Heat 1 0.47
0.9
0.97 0.008 0.009 35.02
26.11
Rem.
Table 171 shows the specification of the tensile properties at room temperature and the test results of the
heats tested here. As for the heat resisting steel castings which were made by the centrifugal casting
process, the difference between the heats is small. However, the tensile strength of the heat resisting steel
casting made by the mould casting process is quite low compared with those of the heats made by the
centrifugal casting process. Although the 0.2% proof stress shows the same tendency, the difference is
small compared with the tensile strength. Moreover, the tensile elongation and the reduction of area of the
mould casted heat are quite low compared with the centrifugal casted heat.
Table 171. Specification of tensile properties
HP(ASTM).
Product
Tensile properties
Tensile strength
[MPa]
Requirement
≥440
Cast tubes
Heat 1
490
(typical example) Heat 2
530
Heat 3
510
Cast block
Heat 4
400
for 25Cr-35Ni-0.4C steel SCH24(JIS G 5122) and
0.2% proof stress
[MPa]
≥235
245
260
250
230
Elongation
[%]
≥5
10
15
16
9
Reduction of area
[%]
13
12
16
11
Next, the high temperature tensile properties will be reviewed. First, the tensile properties of the heat
resisting steel casting made by the centrifugal casting process are shown below. Fig. 392 a - d show the
tensile strength, 0.2% proof stress, tensile elongation and reduction of area from room temperature to
900 °C. The solid lines express the average level for each tensile property. The decrease of tensile
strength is relatively large in the low temperature region up to near 200 °C , but after that, the tensile
Landolt-Börnstein
New Series VIII/2B
298
2.4 Austenitic stainless steels
600
600
500
500
400
400
0.2 % proof stress [MPa]
Tensile strength (MPa)
strength is almost constant up to near 400 °C . Then, a large decrease takes place in the higher
temperature region, especially the fall of tensile strength is remarkable at 700 °C or more. The tendency
of 0.2% proof stress is almost the same as that of tensile strength up to near 600 °C. However, proof
stress rises near 700 °C and at higher temperatures it decreases again.. The tensile elongation and the
reduction of area are in a tendency completely contrary to 0.2% proof stress, and there is a phenomenon
of falling near 700 °C.
300
200
300
200
100
100
0
0
0
200
400
600
800
0
1000
200
70
70
60
60
Reduction of area (%)
Elongation (%)
600
800
1000
800
1000
b
a
50
40
30
50
40
30
20
20
10
10
0
0
0
200
400
600
800
1000
0
Temperature (℃)
c
400
Temperature (℃)
Temperature (℃)
200
400
600
Temperature (℃)
d
Fig. 392. Tensile strength (a), 0.2% proof stress (b), elongation (c) and reduction of area (d) for 25Cr-35Ni-0.4C
steel made by centrifugal casting process.
Next, the tensile properties of the heat resisting steel casting manufactured by the mould casting process
are shown in Fig. 393 a - d. The tensile strength is lower than that of the steel manufactured by the
centrifugal casting process at room temperature. Although the tensile strength tends to decrease with
increasing temperature, the degree of the decrease is relatively small. The tensile strength of both steels
becomes almost equivalent near 800 °C. 0.2% proof strength of both steel casts is very similar, but there
is no increase of 0.2% proof strength near 700 °C in the mould cast heat resisting steel. Although tensile
elongation and reduction of area are lower than those of the steel manufactured by the centrifugal casting
process , they become almost equivalent at 400 °C or more. The mould casted steels tend to higher values
up to 700 °C and then decrease again up to 800 °C, whereas the centrifugally casted steels have a
minimum at about 700 °C and then strongly increase up to 800 °C.
Landolt-Börnstein
New Series VIII/2B
2.4.17 25Cr-35Ni steel
700
700
600
600
500
500
0.2 % proof stress [MPa]
Tensile strength (MPa)
Ref. p. 300]
400
300
200
299
400
300
200
100
100
0
0
0
200
400
600
800
0
1000
200
600
800
1000
800
1000
b
a
70
70
60
60
Reduction of area (%)
Elongation (%)
400
Temperature (℃)
Temperature (℃)
50
40
30
50
40
30
20
20
10
10
0
0
0
200
400
600
800
1000
0
Temperature (℃)
c
200
400
600
Temperature (℃)
d
Fig. 393. Tensile strength (a), 0.2% proof stress (b), elongation (c) and reduction of area (d) for 25Cr-35Ni-0.4C
steel made by the mould casting process.
2.4.17.3 Creep rupture properties of 25Cr-35Ni steels
Fig. 394 shows the creep rupture strength of 25Cr-35Ni-0.4C steel manufactured by the centrifugal
casting process. Creep tests were carried out at temperatures in 50 °C pitch from 850 °C to 1000 °C and at
stresses between 10 MPa and 69 MPa. The longest time to rupture was over 50,000 h. The test data shows
a slightly larger scattering for high temperatures.
Fig. 395 shows the creep rupture strength of 25Cr-35Ni-0.4C steel manufactured by the mould casting
process. Although there are less data points as in Fig. 394, it can be said that the creep rupture strength is
at an equivalent level for both steels. However, there is a tendency that the degree of the decreasing of
creep rupture strength in the mould steel casting is lower than that in the centrifugal steel casting for
1000 °C and 1100 °C.
Landolt-Börnstein
New Series VIII/2B
300
2.4 Austenitic stainless steels
Stress (MPa)
100
10
T=850℃
T=900℃
T=950℃
T=1000℃
T=1100℃
101
102
Fig. 394.
Creep rupture
strength data for 25Cr-35Ni0.4C steel made by the
centrifugal casting process.
103
104
105
Time to rupture (h)
Stress (MPa)
100
10
T=850℃
T=900℃
T=950℃
T=1000℃
T=1100℃
101
102
Fig. 395.
Creep rupture
strength data for 25Cr-35Ni0.4C steel made by the mould
casting process.
103
104
105
Time to rupture (h)
2.4.17.4 Reference
[1] National Research Institute for Metals : NRIM Creep Data Sheet, No. 38A, (1991).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 308]
2.4.18 27Cr-32Ni-Nb-Ce steel
301
2.4.18 27Cr-32Ni-Nb-Ce steel
2.4.18.1 Introduction
27Cr-32Ni-Nb-Ce steel (Grade X6NiCrNbCe 32-27, also called AC 66) is a development based on alloy
800.
Originally it was invented to be used in coal gasification plants with process gases including H2S,
SO2, HCl, CO etc. operating at 950 °C. It is characterized by an austenitic microstructure with 28 %
chromium, 32 % nickel and 0.05 % - 0.10 % cerium. Owing to its generally good corrosion resistance,
further applications were found in waste incineration plants in which AC 66 tubes have been successfully
operated as super-heater bundles at 500 to 600 °C.
The material has a fully austenitic grain structure (Fig. 396) with a high thermal stability. The
mechanical properties remain unchanged after exposure up to 30,000 h at temperatures up to 1000 °C.
Embrittlement due to intermetallic phases like the σ phase does not occur.
Fig. 396. Tenitic structure of AC 66.
For tubes and pipes with a wall thickness of less than 20 mm, grain sizes of ASTM 5-7 are usually
obtained. A grain size of ASTM 3-4 must be assumed for pipes with larger wall thickness.
2.4.18.2 Material standards, chemical composition and tensile requirements
Table 172. Chemical requirements of X6NiCrNbCe 32-27 steel tubes concerning: ASTM A 213, A 321,
VdTÜV data sheet 497 (Grade-no. 1.4877) and EN 10095.
Std.
Chemical composition [wt%]
Designation
Standard
No.
C
Si
Mn P
S
Cr Ni Al
Nb Ce
A213 UNS S
0.04
26.0 31.0
0.6 0.05
ASTM
≤0.30 ≤1.0 ≤0.020 ≤0.015
≤0.025
A312 33228
0.08
28.0 33.0
1.0 0.10
0.04
26.0 31.0
0.6 0.05
VdTÜV 497
1.4877
≤0.30 ≤1.0 ≤0.015 ≤0.010
≤0.025
0.08
28.0 33.0
1.0 0.10
X6NiCrNbCe 0.04
26.0 31.0
0.6 0.05
EN
10095
≤0.30 ≤1.0 ≤0.020 ≤0.010
≤0.025
32-27
0.08
28.0 33.0
1.0 0.10
The formation of protective oxide layers is guaranteed by the high chromium content of about 27 %,
which is high enough to ensure the healing of the scale cracks caused by mechanical or temperatureinduced stressing. Scale spalling due to temperature changes has been significantly reduced through the
addition of cerium. The limitation of the aluminium and silicon contents is necessary in order to ensure
the required resistance to internal oxidation. The niobium content improves the tensile properties.
Landolt-Börnstein
New Series VIII/2B
302
2.4 Austenitic stainless steels
Normally all products are delivered in a solution annealed condition with an annealing temperature
between 1120 °C and 1180 °C, according to the VdTÜV data sheet 497. According to the EN 10095
standard, the recommended annealing temperature range is 1050 - 1150 °C.
Table 173. Room temperature mechanical property requirements for X6NiCrNbCe 32-27. Rp0.2:
Minimum 0.2% proof strength; Rp1.0: Minimum 1.0% proof strength; Rm: Tensile strength.
Section size Rp0.2
Rp1.0
Rm
Standard Std. No. Designation
Product
[mm]
[Nmm-2] [Nmm-2] [Nmm-2]
VdTÜV 497
1.4877
Plate
185
500-750
≤30
VdTÜV 497
1.4877
Bar
185
500-750
≤250
VdTÜV 497
1.4877
Tube
<180
185
215
500-750
EN
10095 X6NiCrNbCe32-27 Plate
180
500-750
≤75
EN
10095 X6NiCrNbCe32-27 Bar
180
500-750
≤160
Table 174. Minimum 0.2% and 1.0% proof strength values and tensile strength values at elevated
temperatures for X6NiCrNbCe 32-27
Std.
Minimum 0.2% proof strength, Rp0.2 [Nmm-2] at a temperature [°C] of
Designation
Standard
No
100 200 300 350 400 450 500 550 600 650 700
VdTÜV
497
1.4877
160 140 120 110 105 100 95
90
90
85
80
Standard
VdTÜV
Standard
VdTÜV
Std.
No
497
Std.
No
497
Designation
1.4877
Designation
1.4877
Minimum 1.0% proof strength, Rp1.0 [Nmm-2] at a temperature [°C] of
100 200 300 350 400 450 500 550 600 650 700
190 170 145 135 130 125 115 110 110 105 100
100
450
Minimum Rm [Nmm-2] at a temperature [°C] of
200 300 350 400 450 500 550 600 650
430 410 400 390 380 370 360 340 300
700
250
2.4.18.3 Creep rupture strength
Creep tests [3], [4] and [8] have been carried out on X6NiCrNbCe 32-27 (AC 66) in the temperature
range between 600 and 950 °C. The test results are represented in Fig. 397 in form of a Larson-Miller
plot. With a few exceptions all rupture points are included within a scatterband of ± 30 % in stress.
2.6
2.4
2.2
2.0
Log (σ)
1.8
1.6
1.4
1.2
Data
Mean Line
1.0
0.8
Fig. 397. Larson-Miller plot of creep rupture strength
of AC 66 (Data: 600 - 950 °C, 3.8 - 320 MPa;
scatterband: ±30 % in stress).
0.6
0.4
16
18
20
22
24
26
28
T (16.76 + log t R) / 1000
Landolt-Börnstein
New Series VIII/2B
Ref. p. 308]
2.4.18 27Cr-32Ni-Nb-Ce steel
303
Since the material originally was developed for application in coal gasification plants, most tests were
carried out between 850 and 950 °C (Fig. 398 to 400).
100
Stress (MPa)
850°C
10
Data
Mean Line
1
10
100
1000
10000
100000
100
Au, Z (%)
80
60
40
Fig. 398. Test results on creep
rupture strength and ductility
at 850 °C (Mean line acc.
Larson-Miller
evaluation;
scatterband: + 30 % in stress).
Au
Z
20
0
10
Landolt-Börnstein
New Series VIII/2B
100
1000
Time (h)
10000
100000
304
2.4 Austenitic stainless steels
100
Stress (MPa)
900°C
10
Data
Mean Line
1
10
100
1000
10000
100000
100
Au, Z (%)
80
Au
Z
60
40
20
0
10
100
1000
Time (h)
10000
100000
Fig. 399. Test results on creep
rupture strength and ductility
at 900 °C (Mean line acc.
Larson-Miller
evaluation;
scatterband: + 30 % in stress).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 308]
2.4.18 27Cr-32Ni-Nb-Ce steel
305
100
Stress (MPa)
950°C
10
Data
Mean Line
1
10
100
1000
10000
100000
100
Au, Z (%)
80
Au
Z
60
40
Fig. 400. Test results on creep
rupture strength and ductility
at 950 °C (Mean line acc.
Larson-Miller
evaluation;
scatterband: + 30 % in stress).
20
0
10
100
1000
Time (h)
10000
100000
The upper half of the figures shows creep rupture strength with respect to time, including the mean line
and scatterband from the Larson-Miller evaluation. In the lower half the accompanying rupture ductility
values are given, i.e. rupture elongation (Au) and reduction of area (Z). In most cases high ductility values
of more than 20 % are obtained. The tests have been performed both in air and in simulated coal
gasification atmosphere, but no difference with respect to the test atmosphere could be observed.
Only a limited number of tests have been carried out at lower temperatures (Fig. 401). The values
follow the trend as given by the Larson-Miller mean lines.
Landolt-Börnstein
New Series VIII/2B
306
2.4 Austenitic stainless steels
1000
600°C
650°C
Stress (MPa)
700°C
750°C
100
800°C
600°C
650°C
700°C
750°C
800°C
Fig. 401. Creep test results at
600, 650, 700, 750 and 800 °C
(Mean lines acc. Larson-Miller
evaluation).
10
10
100
1000
10000
100000
Time (h)
Fig. 402 shows the characteristic strength properties which are used for design. The mean line of the
mean 105 h creep rupture intersects the lines of minimum 0.2% and 1.0% proof strength in the range of
630 to 660 °C. A design below that temperature range uses strength values which are not time dependent.
180
160
5
Rm (10 h)
140
Strength (MPa)
120
100
Rp1.0 (Min)
80
Rp0.2 (Min)
60
40
20
0
500
Fig. 402. Design strength values of AC 66.
600
700
800
900
1000
Temperature (°C)
The creep rupture strength of X6NiCrNbCe 32-27 is shown in Fig. 403. The analysis from which the data
in the figure are derived was carried out as part of the activities of the European Creep Collaborative
Committee and additional details can be found in their published data sheets [9].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 308]
2.4.18 27Cr-32Ni-Nb-Ce steel
307
1000
10,000h
100,000h
Stress [MPa]
100
10
Fig. 403.
Creep rupture
strength of X6NiCrNbCe 3227.
1
500
600
700
800
900
1000
Temperature [°C]
The creep rupture properties have been obtained by comparative analysis of strength values as reported in
VdTÜV data sheet 497. Due to the comparative analysis assessment method, no master equation is
available.
The test data used for the original analyses were related to several casts tested at temperatures of 580 950 °C.
2.4.18.4 Estimated long term creep rupture strength
Based on the data shown in Fig. 403 the 100,000 h rupture strength values for a range of temperatures are
as follows:
Temperature
Stress
100,000 h rupture strengths [Nmm-2] at specified temperatures [°C]
580
590
600
610
620
630
640
650
160
150
140
130
120
111
101
92
660
83
670
74
Temperature
Stress
100,000 h rupture strengths [Nmm-2] at specified temperatures [°C]
680
690
700
710
720
730
740
750
66
58
52
45
39
34
30
27
760
24.5
770
22
Landolt-Börnstein
New Series VIII/2B
308
2.4 Austenitic stainless steels
Temperature
Stress
100,000 h rupture strengths [Nmm-2] at specified temperatures [°C]
780
790
800
810
820
830
840
850
20
18
16
14.5
13
11.5
10
9
860
8
Temperature
Stress
100,000 h rupture strengths [Nmm-2] at specified temperatures [°C]
870
880
890
900
910
920
930
7
6.3
5.6
5
4.5
4
3.6
950
3
940
3.3
All of the values have involved extended extrapolation
Other 104 h and 105 h creep rupture strength values between 580 and 1000 °C are listed for AC 66 steel in
[1].
2.4.18.5 References
[1] VdTÜV data sheet 497: High-Temperature Rolled and Forged Steel; X5 NiCrCeNb 32-27; GradeNo 1.4877 [in German], 1998.
[2] Lindemann, J., and Schendler, W.: Ein neuer Fe-Ni-Cr-Werkstoff (AC 66) für Komponenten in
Combined-cycle Kraftwerken sowie in modernen Müllverbrennungsanlagen. VGB
Kraftwerkstechnik 71 Jahrgang, Heft 8, August (1991), Seite 746-754 [in German].
[3] Final Report WE 998 B; 23rd. February (1989) of TÜV RHEINLAND.
[4] Bendick, W., Lindemann, J., Schendler, W.: Development of a New High-temperature Alloy for
Coal Gasification; Combined Cycle Processes and Waste Incinceration High-temperature Materials
for Power Engineering, Sept. 24-27, (1990); Liège, Belgium.
[5] Final Report Legierungsentwicklung für den Wärmetauscher der Wasserdampf-Kohlevergasung
01.01. (1982) - 31.12. (1988), Bergbauforschung, Mannesmann Forschungsinstitut [in German].
[6] R. Plür: Experience with and Measures for the Reduction of Corrosion and Flue Gas-side Deposition
on the Heating Surfaces of the Boilers of GMVA Niederrhein GmbH VGB Kraftwerkstechnik
Vol.70, No. 8, August (1990), 589-596
[7] AC 66: Iron-nickel-chrome alloy for high temperature service; MANNESMANN
EDELSTAHLROHR; information paper Edition (1992).
[8] Collected data of the MANNESMANN FORSCHUNGSINSTITUT GmbH, Duisburg.
[9] ECCC Data sheet for X6NiCrNbCe 32-27, European Creep Collaborative Committee, BRITE
EURAM Thematic Network BET2-0509 “Weld Creep”, (1999).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 311]
2.4.19 30Cr-50Ni-2Mo-Ti-Zr-B steel
309
2.4.19 30Cr-50Ni-2Mo-Ti-Zr-B steel
2.4.19.1 Introduction
30Cr-50Ni-2Mo-Ti-Zr-B steel (TEMPALOY CR30A, or CR30A for short) was developed by NKK,
Japan in the 1980s for heat exchanger tubings such as boiler superheaters and reheaters, steam generators,
chemical plants operated at very high temperatures of around 700 °C and in severely corrosive
environments. The CR30A steel features well-balanced hot corrosion resistance through increased Cr
content, high temperature creep rupture strength by alloying with Mo and improved fabricability and
weldability by means of the addition of Ti, Zr and B. The creep rupture strength of CR30A is comparable
to that of 17-14CuMo. The hot corrosion resistance of CR30A under the environmental conditions of 1 %
SO2 containing gas with equivalent amounts of 5 % O2, 15 % CO2 and N2, and molten ash with 34
%Na2SO4 - 41 % K2SO4 - 25 % Fe2O3 is much better than that of Alloy 617 and Alloy 800H at 700 °C.
2.4.19.2 Material standards, chemical and tensile requirements [1]
CR30A steel is not specified in any codes or standards except the internal specifications of NKK.
Table 175 and Table 176 respectively show the requirements for the chemical composition and tensile
properties specified by NKK. The tensile properties at room temperature are almost equivalent to those of
18Cr-8Ni stainless steels.
Fig. 404 shows the absorbed energies obtained by Charpy impact tests at 0 °C using full size specimens
with a 2 mm V-notch [2].
Table 175. Typical chemical composition of CR30A steel; [1].
Chemical composition [wt%]
Product
C
Si Mn Ni Cr Mo Al Ti Zr B
Fe
Tube
0.06 0.2 0.1 50 30 2
0.1 0.2 0.03 0.005 Bal.
Table 176. Typical tensile properties of CR30A steel at room temperature; [1].
Product Yield strenth (0.2% offset) Tensile strength Elongation
[MPa]
[MPa]
[%]
Tube
279
612
75
2.4.19.3 Creep properties [2], [3]
2.4.19.3.1 Creep rupture data and creep rupture strength
Table 177 lists the chemical compositions of the steels tested, which were melted as 150 kg ingots and
solution treated at 1180 °C for 0.5 h after rolling the plate to a 15 mm thickness.
Fig. 405 shows the creep rupture stress vs. time to rupture diagram for CR30A steel in the temperature
range of 600 °C to 1000 °C. The Larson-Miller plot of the creep rupture data for CR30A steel is shown in
Fig. 406. The creep rupture strength of CR30A is comparable to that of 17-14CuMo steel. The 100,000 h
rupture strengths obtained by linear extrapolation of the stress vs. time to rupture curves of CR30A steel
at 650 °C and 700 °C are 147 MPa and 108 MPa, respectively. However, the Larson-Miller plot yields
rather conservative estimations, i.e., 125 MPa for 650 °C and 843 MPa for 700 °C.
Landolt-Börnstein
New Series VIII/2B
310
2.4 Austenitic stainless steels
Temperature [°C ]
v E0 [J]
900
66
45
49
800
90
40
27
66
53
29
80
45
27
14
53
26
600
180
150
130
500
374
382
330
300
1000
3000
Aging time [h]
700
Fig. 404. Time-Temperature-Absorbed energy map at 0 °C
for CR30A steel; [2].
10000
Table 177. Chemical compositions of alloys tested; [3].
Chemical composition [wt%]
Alloy
C
Si
Mn P
S
Ni
Cr
Mo Al Ti
Zr
N
Fe
CR30A 0.058 0.27 0.20 0.001 0.0007 50.99 30.52 2.11 0.14 0.181 0.027 0.0016 Bal.
ZD3
0.004 0.32 0.22 0.001 0.0007 56.19 24.12 2.89 0.21 0.200 0.029 0.0022 Bal.
500
46
38
43
300
40
200
Stress (MPa)
36
38
31
32
20
44
18
19
3132
30
24
22
23
14
12
23 23
13
17 17
26
59
45
61
650℃
750℃
65
57
74
800℃
45
40
600℃
700℃
49
63
70
50
13
26
18
61
70
15
18
47
100
36
49
42
46
30
20
43
43 30
51
1000℃
32
102
101
900℃
41
Values designate
rupture elongation in %.
103
104
105
Fig. 405. Creep rupture strength
data of CR30A steel; [1].
Time to rupture (h）
500
550℃
600℃
650℃
700℃
750℃
800℃
900℃
1000℃
300
Stress (MPa)
200
100
70
50
30
20
10
18
20
22
24
26
28
30
Fig. 406. Creep rupture master curve of CR30A
steel by Larson-Miller parametric method; [3].
T (20＋log t) × 10- 3
Landolt-Börnstein
New Series VIII/2B
Ref. p. 311]
2.4.19 30Cr-50Ni-2Mo-Ti-Zr-B steel
311
2.4.19.3.2 Microstructural changes during creep
The excellent long-term creep rupture strength was found to be due to a high density and uniform
dispersion of intergranular precipitates in the bcc chromium phase. Very fine M23C6 carbide also
precipitates at the dislocation lines within the grains of austenite during creep, which contributes to
enhanced creep strength.
2.4.19.3.3 Creep deformation
Creep deformation was measured at 600 °C and 294 MPa for the as-solution treated steel, for steel aged at
700 °C for 1000 h to precipitate α' in the grains, and for a model alloy (ZD3 in Table 177) simulating a
single phase structure of γ which appears as a matrix of the aged material. Fig. 407 shows creep curves
for these three specimens. The solution-treated steel exhibited a 5 times longer rupture time against ZD3
with a single phase of γ, and the minimum creep rate was 3.5×10−4 h−1 which is 1/5 of that of ZD3. These
enhancements of creep in solution-treated steel are considered to be due to the dense precipitations of
M23C6 at the dislocation lines. Aged steel with precipitations of α' in the γ grain exhibited a rupture time
of twice as long as that of solution-treated steel. In this aged steel most of the M23C6 precipitated at the
grain boundaries.
30
Fe-30Cr-50Ni-0Mo
700 °C, 1000 h aged
25
Stress [MPa]
20
ZD3 solution
treated
Fe-30Cr-50Ni-2Mo
solution treated
15
10
5
0
100
200
300
400
Time [h]
500
600
Fig. 407. Creep curves at 650 °C - 294 MPa for ZD3
solution treated and Fe-30Cr-50Ni-2Mo steel solution
treated and aged at 700 °C for 1000 h; [3].
2.4.19.4 References
[1] Tamura, M., and Murase, S.: NKK Technical Review, No.56, (1989), 93.
[2] Tamura, M., Yamanouchi, N., Tanimura, M., and Murase, S.: Proc. (1985) Exposition and
Symposium on Industrial Heat Exchanger Technology, Pittsburgh, PA., (1985), 273.
[3] Yamanouchi, N., Shirada, T., Tamura, M., Matsuo, T., and Kikuchi, M.: Tetsu-to-Hagane, 76
(1990), 1179.
Landolt-Börnstein
New Series VIII/2B
312
2.4 Austenitic stainless steels
2.4.20 21Cr-11Ni-Si-N-Ce steel
2.4.20.1 Introduction
21Cr-11Ni-Si-N-Ce steel (EN 1.4835, X9CrNiSiNCe 21-11-2, UNS S30815) is an austenitic stainless
steel designed primarily for use at temperatures exceeding 550 °C, i.e. in the temperature range where
creep is controlling the mechanical strength and where oxidation resistance is of importance. EN 1.4835
is represented in the standard EN 10095 [1]. The trade name of EN 1.4835 is 253MA [2]. In comparison
with traditional stainless steels, EN 1.4835 has an increased nitrogen content and has been microalloyed
with rare earth metals (REM). The most suitable temperature range is 850 - 1100 °C, since structural
changes below 850 °C can lead to reduced impact toughness at room temperature. Since the steel is
optimized for high temperature service the resistance to aqueous corrosion is limited.
EN 1.4835 is used in a number of high temperature applications for example for components within
iron, steel, and non-ferrous industries; engineering industry; energy conversion plants, and cement
industry.
Like other austenitic steels, EN 1.4835 can be formed in hot and cold condition. However, as a result
of the relatively high nitrogen content, the mechanical strength is higher and consequently greater
deformation forces will be required. Also the ability to strain harden must be taken into consideration in
connection with machining. The steel has good weldability and can be welded using shielded metal arc
(SMA) welding with covered electrode or gas shielded welding. The latter method has given the best
creep properties for welds.
2.4.20.2 Material standards, chemical composition and tensile properties
Table 178. Chemical requirements of X9CrNiSiNCe 21-11-2.
Standard
Std. No. Designation
EN
10095
ECCC
Data
sheet
C
0.05
X9CrNiSiNCe
21-11-2
0.12
X9CrNiSiNCe 21-11-2 0.05
(X7CrNiSiNCe 21-11) 0.10
Chemical composition [wt%]
Si Mn P
S
Cr
Ni
1.40
20.00 10.00
≤1.00 ≤0.045 ≤0.015
22.00 12.00
2.50
1.40
20.00 10.00
≤0.80 ≤0.040 ≤0.030
22.00 12.00
2.00
N
0.12
0.20
0.14
0.20
Table 179. Room temperature mechanical property requirements for X9CrNiSiNCe 21-11-2.
StanHeat
Thickness Rp0.2
Rp1.0
Rm
Std. No.
Designation
dard
treat
[mm]
[MPa] [MPa] [Nmm-2]
X9CrNiSiNCe 21-11- Solution
650
EN
10095
310
350
≤75
2
annealed
850
X9CrNiSiNCe 21-11Solution
640
ECCC Data sheet 2
295
345
≤30
annealed
850
(X7CrNiSiNCe 21-11)
Ce
0.03
0.08
0.03
0.08
A
[%]
33
-
Table 180. Minimum 0.2% proof strength values at elevated temperatures for X9CrNiSiNCe 21-11-2.
Standard Designation
Thickness
Heat treat
[mm]
Avesta
1.4835
Polarit [2]
≤75
Solution
annealed
Minimum 0.2% proof strength, Rp0.2 [MPa] at a
temperature [°C] of
50
100 200 300 400 500 600 700
280
230
185
170
160
150
140
130
Landolt-Börnstein
New Series VIII/2B
Ref. p. 314]
2.4.20 21Cr-11Ni-Si-N-Ce steel
313
2.4.20.3 Creep rupture strength
The creep rupture strength of X9CrNiSiNCe 21-11-2 is shown in Fig. 408. The values have been derived
from experimental creep data using the free temperature time temperature parameter. This was carried out
as part of the activities of the European Creep Collaborative Committee [3] and [4].
Rupture strength, MPa
200
X9CrNiSiNCe 21-11-2
100
10000 h
100000 h
200000 h
50
20
10
5
500
Fig. 408. Creep rupture strength data
of X9CrNiSiNCe 21-11-2; [3], [4].
600
700
800
900
1000
1100
Temperature, °C
2.4.20.4 Estimated long term creep rupture strength
Values from Fig. 408 as well as from EN 10095 are summarized in Table 181 [1] and [3].
Table 181. Creep strength to rupture and to 1% strain of X9CrNiSiNCe 21-11-2
Standard
Std. No.
Criterion
ECCC
ECCC
ECCC
EN
Data sheet Data sheet Data sheet 10095
Rupture
Rupture
Rupture
Rupture
Temperature [°C]
550
600
650
700
750
800
850
900
950
1000
1050
1100
EN
10095
1% strain
EN
10095
1% strain
10,000 h
100,000 h 200,000 h 10,000 h 100,000 h 10,000 h
100,000 h
208.3
138.2
91.8
61.2
41.2
28.2
19.6
14
10.3
7.8
6.1
5.0
149.4*
88.2
54.4*
35.1*
23.5*
16.3
11.8
8.7
6.7*
5.3*
4.2*
3.5*
135.2*
77.0*
46.5*
29.6*
19.8*
13.9*
10.1*
7.6*
5.9*
4.7*
-
EN
10095
Rupture
157
88
126
80
63
35
45
26
27
15
19
11
13
8
10
6
7*
4*
5*
3*
* : Values which have involved extended time extrapolation (more than a factor of three in time)
- : Values which have involved extended stress extrapolation
Landolt-Börnstein
New Series VIII/2B
314
2.4 Austenitic stainless steels
The errors in the ECCC values in Table 181 due to the uncertainties in the evaluation procedure, are about
5 % of the values or 1 MPa whichever is the largest. In spite of this the values in Table 181 are given with
2 to 4 figures to avoid unnecessary loss in precision.
2.4.20.5 References
[1] European Standard EN 10095 “Heat-resisting steels and nickel alloys” (1999).
[2] High Temperature Stainless Steel, Data sheet, AvestaPolarit (2002).
[3] ECCC Data sheet for of X7CrNiSiNCe 21-11, European Creep Collaborative Committee, BRITE
EURAM Thematic Network BET2-0509 “Weld Creep”, (1999).
[4] Lindé, L, Sandström, R, Gommans, R, Spindler, M. W., An Evaluation of Creep Rupture Data for
the New European Standard for Stainless Steels-Edition 3, European Creep Collaborative
Committee, BRITE EURAM Thematic Network BET2-0509 “Weld Creep (1999).
[5] Lindé, L, Sandström, R, Creep rupture of the 21Cr 11Ni Si N Rem Stainless steel 253MA- Data
collation and assessment, Swedish Institute for Metals Research, Report IM-3572 (1998).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 318]
2.4.21 15Cr-10Ni-1Mo-Mn-Nb-V-B steel
315
2.4.21 15Cr-10Ni-1Mo-Mn-Nb-V-B steel
2.4.21.1 Introduction
15Cr-10Ni-1Mo-Mn-Nb-V-B steel (X10CrNiMoMnNbVB 15-10-1, Esshete 1250) is an austenitic steel
made by Corus in the UK. It is well established as superheater boiler tube material in UK power stations
[1] and has an approximately 30 % creep strength advantage over Type 316, which is a recognized
standard for similar applications. Esshete 1250 is readily welded with either inert gas welding with
wires/rods or metal arc welding with covered electrodes.
In addition to its application as boiler tubing Esshete 1250 has been used for piping and headers in
super critical plant [1]. Furthermore, warm working can increase the tensile and creep properties of
Esshete 1250. This enables the material to be used for bolting applications [2].
In the solution treated condition the microstructure of Esshete 1250 typically consists of dispersed
niobium carbides in an austenite matrix. The microstructure of boiler tubing is fine equiaxed austenite
grains, mainly of size ASTM 9 but some at ASTM 7 or 8. Long trails of Nb(CN) particles and a fine
background dispersion of carbides in broad bands aligned with the longitudinal direction of the tube wall
are also present.
2.4.21.2 Material standards and chemical composition and tensile requirements
Table 182. Chemical requirements of X10CrNiMoMnNbVB 15-10-1 (Esshete 1250) steel tubes for
pressure purposes concerning EN 10216-5, ASTM A213 and BS 3605 Part 1.
Chemical composition [wt%]
Stan- Std.
Designation
dard
No.
C
Si
Mn P
S
Cr Ni Mo Nb V
B
X10CrNiMoMn 0.06 0.20 5.50 ≤
14.0 9.0 0.80 0.75 0.15 0.003
≤
EN
10216-5
NbVB 15-10-1 0.15 1.00 7.00 0.035 0.015 16.0 11.0 1.20 1.25 0.40 0.009
0.06 0.20 5.50 ≤
14.0 9.0 0.80 0.75 0.15 0.003
≤
ASTM A213
UNS S 215000
0.15 1.00 7.00 0.040 0.030 16.0 11.0 1.20 1.25 0.40 0.009
3605
0.06 0.20 5.50 ≤
14.0 9.0 0.80 0.75 0.15 0.003
≤
BS
215S15
Part 1
0.15 1.00 7.00 0.040 0.030 16.0 11.0 1.20 1.25 0.40 0.009
X10CrNiMoMnNbVB 15-10-1 (Esshete 1250) tubes are usually solution heat treated at a temperature
range of 1050 to 1150 °C.
Table 183. Room temperature minimum mechanical property requirements for X10CrNiMoMnNbVB
15-10-1 (Esshete 1250).
R
Rp1.0
Rm
Standard Std. No. Designation p0.2 -2
[Nmm-2]
[Nmm ]
[Nmm-2]
540
3605
215S15
220
270
BS
740
Part 1
Section size ≤250 mm
Table 184. Minimum 0.2% and 1.0% proof strength values at elevated temperatures for
X10CrNiMoMnNbVB 15-10-1 (Esshete 1250).
DesigMinimum 0.2% proof strength, Rp0.2 [Nmm-2] at a temperature [°C] of
Standard Std. No.
nation
100 150 200 250 300 350 400 450 500 550 600 650
3605
215S15 188 171 161 153 148 145 144 141 139 136 133 130
BS
Part 1
Section size ≤250 mm
Landolt-Börnstein
New Series VIII/2B
316
2.4 Austenitic stainless steels
Table 184 cont.
Standard Std. No.
3605
Part 1
BS
Designation
Minimum 1.0% proof strength, Rp1.0 [Nmm-2] at a temperature [°C]
100 150 200 250 300 350 400 450 500 550 600 650
215S15
232 210 195 190 187 184 182 179 178 175 170 165
Section size ≤250 mm
2.4.21.3 Creep rupture strength
Creep rupture tests have been carried out on X10CrNiMoMnNbVB 15-10-1 (Esshete 1250) in the
temperature range between 550 and 950 °C. In addition to the data reported in [3], a considerable quantity
of long term creep rupture tests were performed by British Steel and ERA Technology Ltd in the UK. The
data are from 218 casts and the test durations extend to 178,988 h (at 700 °C), 143,109 h (at 650 °C) and
153,491 h (at 600 °C).
1000
Stress [MPa]
10,000h
30,000h
100,000h
200,000h
250,000h
100
Fig. 409. Creep rupture
strength data of
X10CrNiMoMnNbVB 15-10-1
(Esshete 1250).
10
550
600
650
700
750
800
Temperature [°C]
Not continuous line denotes extended time extrapolation.
The creep rupture test results at 600 to 800 °C have been analyzed in [4] using the standard ISO 6303
method. The extent of the data that were analyzed is shown in Table 4.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 318]
2.4.21 15Cr-10Ni-1Mo-Mn-Nb-V-B steel
Table 185. Distribution of test durations used to derive the stress rupture properties of
X10CrNiMoMnNbVB 15-10-1 (Esshete 1250).
Number of test points at the various test durations
10,000 20,001 30,001 50,001 70,001 <10,000 h
20,000 h
30,000 h
50,000 h
70,000 h
100,000 h
1227 (17)
102 (3)
58 (2)
51 (18)
17 (9)
13 (25)
317
>100,000h
9
( ) denotes unbroken tests
The master curve, that was derived with the ISO 6303 method, is described by:
P (σ ) = a + b (logσ ) + c (logσ ) 2 + d (logσ ) 3 + e (logσ ) 4 =
log t - log t a
(T - Ta ) r
where P(σ) is the creep rupture parameter; T is the temperature in Kelvin; t is the time to rupture in hours;
2
σ is the stress in N/mm ; r is a temperature exponent; and a, b, c, d, e, r, Ta and log ta are constants (see
the following table.
log( ta )
Ta r a
b
c
d
e
−18.431352615 0 −1 −336414.5313 799872.0625 −654687.0625 234517.4219 −31193.5293
2.4.21.4 Creep deformation behavior
The creep strength (1 % plastic strain) data for X10CrNiMoMnNbVB 15-10-1 have been assessed using
the Larson Miller parameter to predict isochronous curves for up to 100,000 h at 600 to 700 °C.
300
[h]
100
1000
3000
10000
20000
30000
50000
100000
Stress [N/mm 2 ]
200
100
80
60
40
20
600
Landolt-Börnstein
New Series VIII/2B
700
800
Temperature [°C]
900
Fig. 410.
Isochronous creep curves for X10
CrNiMoMnNbVB 15-10-1.
318
2.4 Austenitic stainless steels
2.4.21.5 Estimated long term creep rupture strength
Based on the data shown in Fig. 409 the 100,000 h, 200,000 h and 250,000 h rupture strength values for a
range of temperatures are as follows:
100,000 h rupture strengths at specified temperatures
Temperature [°C] 600 610 620 630 640 650 660 670
Stress [Nmm-2]
199 185 167 147 122 100 84 74
100,000 h rupture strengths at specified temperatures
Temperature [°C] 680 690 700 710 720 730 740 750
Stress [Nmm-2]
66 59 54 49 45 40* 36* 30*
* Values which have involved extended time extrapolation
200,000 h rupture strengths at specified temperatures
Temperature [°C] 600 610 620 630 640 650 660
Stress [Nmm-2]
183 165 143 118 97 82 72
200,000 h rupture strengths at specified temperatures
Temperature [°C] 670 680 690 700 710 720 730
Stress [Nmm-2]
64 58 52 48 43 39 35*
* Values which have involved extended time extrapolation
250,000 h rupture strengths at specified temperatures
Temperature [°C] 600 610 620 630 640 650 660
Stress [Nmm-2]
177 158 134 109* 90* 78* 69*
250,000 h rupture strengths at specified temperatures
Temperature [°C] 670 680 690 700 710 720
Stress [Nmm-2]
62* 56* 51* 46* 42* 37*
* Values which have involved extended time extrapolation
2.4.21.6 References
[1] Orr, J., Nileshwar, V. B., Esshete 1250: An advanced austenitic stainless steel for power station
tubes, piping and headers, Stainless steels ’84 Conference, Gottenburg, Sept (1994).
[2] Orr, J., Everson, H., Parkin, G., Warm Worked Esshete 1250: A high strength bolting steel, in
Performance of Bolting Materials in High Temperature Plant Applications, IOM London, UK,
(1995).
[3] Murray, J.D., Hacon, J. Wannell P. H., The high-temperature properties of a advances austenitic
steel: Esshete 1250, in High-Temperature Properties of Steels, ISI Publication 97, The Iron and
Steels Institute, London, UK, (1967).
[4] Burton, D., Orr, J., Dulieu, D., An Assessment of The Stress Rupture Properties of Esshete 1250,
British Steel Report No. S/RSC/S1122/2/88/E, (1988).
[5] ECCC Data sheet for X10CrNiMoMnNbVB 15-10-1, European Creep Collaborative Committee,
BRITE EURAM Thematic Network BET2-0509 “Weld Creep”, (1999).
Landolt-Börnstein
New Series VIII/2B
3 Superalloys - 3.1 Fe-base alloys
319
3 Creep and rupture data of superalloys
3.1 Fe-base alloys
3.1.1 Fe-15Cr-26Ni-Mo-Ti-V alloy
3.1.1.1 Introduction
Fe-15Cr-26Ni-Mo-Ti-V (A286, AISI660, JIS SUH660, ASTM S 66286) alloy is one of the gamma prime
precipitation strengthened type iron-base superalloys. This alloy was developed in the 1950s based on
German Tinidur alloy and has been widely used as jet engine and gas turbine components such as wheels,
disks, blades, shafts, casings and bolts up to about 650 °C because of its good strength, toughness and
oxidation resistance [1]. It contains 2 - 3 % Ti plus Al to precipitate gamma prime phase [Ni3(Al,Ti)], 25
% Ni to stabilize austenitic structure, 15 % Cr for corrosion resistance and Mo for improving solid
solution strengthening and stabilizing gamma prime phase [2]. Because A286 has also good strength and
toughness at cryogenic temperatures, it is used as a cryogenic structural material. It is primarily used as
wrought alloy and used after solid solution heat treatment at about 900 - 1000 °C and aging heat treatment
at about 700 - 760 °C for gamma prime precipitation [3].
3.1.1.2 Material standard and chemical composition
The chemical composition requirement in JIS G 4311 [3], which is nearly the same as for AMS 5805D,
and the mechanical property requirement [3] is given in Table 186. The mechanical properties of the
A286 alloy are given in [1], [4], and [5].
Table 186. Chemical composition and mechanical property requirement in JIS G 4311
(1) Chemical composition requirement
C
Si
Mn
≤0.08
≤1.00
≤2.00
Chemical composition [wt%]
P
S
Ni
Cr
Mo
24.00- 13.50- 1.00≤0.040 ≤0.030
27.00 16.00 1.50
V
0.100.50
Ti
1.902.35
Al
≤0.35
B
0.0010.010
(2) Mechanical property requirement
Proof stress Tensile strength Elongation Reduction of area
[N/mm2]
[N/mm2]
[%]
[%]
≥590
≥900
≥15
≥18
3.1.1.3 Tensile properties
0.2 % proof stress and tensile strength of A286 alloy are shown in Fig. 411 [4].
3.1.1.4 Creep and rupture properties
Creep rupture curves of iron-base A286 alloy at 550 to 700 °C are shown in Fig. 412 [4]. The effect of
grain-size on creep strength of this alloy is shown in Fig. 413 [1]. The rupture-stresses from 100 to
100,000 h as a function of temperatures are available in Fig. 414. Minimum creep rate for A286 alloy is
reported as shown in Fig. 415 [4].
Landolt-Börnstein
New Series VIII/2B
320
3.1 Fe-base alloys
Tensile strength
1200
1000
1000
800
800
Stress (MPa)
Stress (MPa)
0.2% proof stress
1200
600
400
200
0
600
400
200
0
100
200
300
400
500
600
700
0
800
0
100
200
Test temperature (℃)
300
400
500
600
700
800
Test temperature (℃)
Fig. 411. 0.2% proof stress and tensile strength of Fe base A286 superalloy; [4].
800
700
600
５５０℃
６００℃
６５０℃
７００℃
500
400
Stress (MPa)
300
200
100
90
80
70
60
Fig. 412.
Creep rupture
strength data of iron-based
A286 alloy of 3 heat turbine
discs; [4]. n indicates the
total number of data points.
n=79
1
10
102
103
104
105
106
Time to ruputure (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 322]
3.1.1 Fe-15Cr-26Ni-Mo-Ti-V alloy
321
100
A-286 sheet
100-hour rupture strength [ksi ]
60
40
1300 F
20
1400 F
10
6
1500 F
4
1600 F
2
1
1
2
4 6 10 20 40 60 100
Grain size [ µm ]
Fig. 413. Effect of grain size on creep strength of A286;
[1]. 1 ksi = 6.89476 MPa
200 400
2000
100 h
{
600
500
400
300
Tensile
strength
0.2% proof
stress
{
Stress [MPa]
1000
800
200
10000 h
100
80
60
500
550
600
a
700
650
Temperature [°C]
800
750
2000
600
500
400
300
{
200
100
80
60
500
b
Landolt-Börnstein
New Series VIII/2B
Tensile
strength
0.2% proof
stress
{
Stress [MPa]
1000
800
1000 h
Fig. 414. Temperature dependence of creep rupture strength
from 100 to 100000 h; [4].
100000 h
550
600
700
650
Temperature [°C]
750
800
322
3.1 Fe-base alloys
1000
800
Stress [MPa]
600
500
400
300 550 °C
200
600 °C
100
650 °C
700 °C
Fig. 415. Stress vs. minimum creep rate for A286 ironbased superalloy; [4]. n indicates the total number of data
points.
n = 27
80
60
10-7
550 °C
600 °C
650 °C
700 °C
10-6
10-5 10-4 10-3 10-2
Minimum creep rate [%/h]
10-1
1
3.1.1.5 References
[1]
[2]
[3]
[4]
[5]
Aerospace Structural Metals Handbook, vol.2, code 1601 (1987).
The superalloys, ed. by Sims, C. T., and Hagel, W. C., John Wiley & Sons (1972) pp. 20-21.
JIS G4311-1991, Heat-resisting steel bars (1991).
NRIM Creep Data Sheet, No.22B (1993).
Report on the mechanical properties of metals at elevated temperatures, Vol.IV Superalloys, The
Iron and Steel Institute of Japan (1979) pp. 1-21.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 326]
3.1.2 Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy
323
3.1.2 Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy
3.1.2.1 Introduction
Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy (S590, X40CoCrNi2020) is a member of the well-known
20Co-20Cr-20Ni group of superalloys, which were considered outstanding for high temperature service.
It is the strongest of these iron base alloys and is distinguished by its long time stability at high
temperature [1]. Its usage for wheels, shafts and buckets for gas turbines and for forgings requires high
strength up to 1500 °F and oxidation resistance up to 1800 °F. Flat products and castings are also
available; [2] and [3].
The alloy was introduced as S590 for gas turbine buckets in the 1940s [4]. The high creep strength of
S590 is obtained by additions of 4 % Mo, 4 % W, 4 % Nb for strengthening the matrix and 0.4 % C for
carbide precipitation [5]. The 20 % Cr content of the alloy improves oxidation resistance, while both 20
% Ni and Co additions help stabilize the austenite matrix [6].
3.1.2.2 Chemical composition and material preparation
The chemical compositions shown in Table 187 were reported by the manufacturers except for Al and N,
for which the analysis was carried out at NRIM [7]. The chemical requirements coincide with S590 and
AMS 5770B(revision 5770D).
Material preparation conditions are shown in Table 188. The bars were sampled in 1969.
Table 187. Chemical composition of Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy bars
NRIM
reference
code
Requirement
fBA
fBB
fBC
C
Si
Mn
P
S
Chemical composition [wt-%]
Ni
Cr
Mo Cu W
0.38~
0.48
0.43
0.43
0.40
≤
1.00
0.37
0.47
0.56
≤
2.00
1.27
1.51
1.18
≤
0.040
0.008
0.019
0.005
≤
0.030
0.013
0.023
0.003
18.50~
21.50
20.28
19.79
19.85
19.00~
22.00
20.26
19.98
20.25
3.50~
4.50
4.26
3.92
4.12
≤
0.50
0.06
0.06
0.05
3.50~
4.50
4.23
3.98
4.20
Co
Al
N
18.50~
3.50~
21.50
4.50
20.10 0.010 0.0083 4.48
20.09 0.030 0.0114 4.10
19.85 0.044 0.044 3.92
Table 188. Materials preparation of Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy bars.
NRIM
Type of Size of Deoxidation Product Dimen Processing and
Austenite
reference melting ingot
process
form
-sions thermal history
grain size
code
[t]
[mm]
number
fBA
0.15
Al-killed
20 D
Hot rolled
7.1
150 L 1204 °C/1 h WQ
760 °C/10 h AC
HFIVF
fBB
0.10
Al-Ca-Si20 D
Hot rolled
7.2
Bar
killed
580 L 1190 °C/1 h WQ
760 °C/10 h AC
fBC
ESR
1.28
20 D
Hot rolled
7.4
2000 L 1204 °C/1 h WQ
760 °C/10 h AC
HFIVF: high frequency induction vacuum furnace
Landolt-Börnstein
New Series VIII/2B
Nb+Ta
Rockwell
hardness
[HRC]
22
23
24
324
3.1 Fe-base alloys
3.1.2.3 Mechanical properties
The mechanical properties of Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy bars are shown in Fig. 416,
obtained from [7].
0.2 % proof stress
1200
800
800
Stress [MPa]
1000
Stress [MPa]
1000
600
400
upper 95% PI
average
lower 95% PI
0
200
400
600
Test temperature [°C]
200
800
0
1000
80
80
70
70
Reduction of area [%]
90
60
50
40
20
200
800
1000
1000
800
1000
Reduction of area
10
400
600
Test temperature [°C]
800
40
20
n = 38
400
600
Test temperature [°C]
50
30
0
200
n = 38
60
30
0
0
100
90
10
upper 95% PI
average
lower 95% PI
n = 38
Elongation
100
Elongation [%]
600
400
200
0
Tensile strength
1200
0
n = 38
0
200
400
600
Test temperature [°C]
Fig. 416. Short time tensile properties of Fe-20Cr-20Ni-20Co-W-Mo- (Nb+Ta) alloy bars. n indicates the total
number of data points.
3.1.2.4 Creep rupture properties
The creep rupture properties of Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy bars are shown in Fig. 417 to
419, which are obtained from [7].
Fig. 417 shows stress vs. time to rupture data at 650, 700, 750 and 800 °C for Fe-20Cr-20Ni-20Co-WMo-(Nb+Ta) alloy bars. The solid curves are based on the Larson-Miller parameter method for all
available data at various temperatures.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 326]
3.1.2 Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy
325
500
650 °C
700 °C
750 °C
800 °C
400
300
Stress [MPa]
200
100
80
60
50
40
30
10
Fig. 417. Creep rupture strength
data for Fe-20Cr-20Ni-20Co-WMo-(Nb+Ta) alloy bars. n indicates
the total number of data points.
n = 83
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 418 (next page) shows elongation and reduction of area data at 650, 700, 750 and 800 °C for Fe20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy bars. It shows the tendency that elongation and reduction of area
are decreasing with increasing temperature and rupture times.
Fig. 419 shows the master rupture curve obtained by the Larson-Miller parameter method for Fe-20Cr20Ni-20Co-W-Mo-(Nb+Ta) alloy bars. It should be noted that the creep rupture strength has some scatter
at high values of this parameter, corresponding to the scatter in creep rupture strength for long times at
750 and 800 °C in Fig. 417.
500
400
300
650 °C
700 °C
750 °C
800 °C
Stress [MPa]
200
100
80
60
average curve
40
n = 83 (88)1)
30
18
24
19 20
25 26
22
21
23
Larson-Miller-parameter TK (log t R -18.597) [×10 3 ]
Landolt-Börnstein
New Series VIII/2B
Fig. 419. Master rupture curve by the Larson-Miller
parameter method for Fe-20Cr-20Ni-20Co-W-Mo(Nb+Ta) alloy bars. n indicates the total number of data
points.
326
3.1 Fe-base alloys
100
100
650 °C
650 °C
80
80
60
60
40
40
20
20
0
n = 23
0
n = 23
100
100
700 °C
700 °C
80
60
60
40
40
Reduction of area [%]
80
Elongation [%]
20
0
n = 23
100
750 °C
80
20
0
100
750 °C
80
60
60
40
40
20
20
0
100
n = 23
n = 17
0
100
n = 17
800 °C
800 °C
80
80
60
60
40
40
20
20
0
10
n = 20
10
2
3
10
10
Time to rupture [h]
4
10
5
0
10
n = 20
10 2
10 3
10 4
Time to rupture [h]
10 5
Fig. 418. Elongation and reduction of area at 650, 700, 750 and 800 °C for
Fe-20Cr-20Ni-20Co-W-Mo-(Nb+Ta) alloy bars. n indicates the total number of data points in
each diagram.
3.1.2.5 References
[1]
[2]
[3]
[4]
[5]
[6]
[7]
Aerospace Structural Metals Handbook, Volume I, Ferrous Alloys, 1963, Code 1603, p. 1.
Metal Handbook, 8th Edition, Vol.1 Properties and Selection of Metals, AMS (1961), p. 487.
Aeronautical Material Specification, AMS5770B (1951).
Buergel, R.; Materialwise Werkstofftech , Vol.23, No.8, (1992) 287-292.
Matsunaga, Y.; Stainless Steel-Heat Resisting Alloy, Seibunndou Shinkousya (1963) p. 231.
Imai, Y.; High Temperature Materials Handbook, Asakura Shoten, (1965) p. 92.
NRIM Creep Data Sheet, No.23B, Fe based 20Cr-20Ni-20Co-W-Mo-(Nb+Ta)bars, (1989).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 330]
3.1.3 Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy
327
3.1.3 Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy
3.1.3.1 Introduction
Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy (N-155, JIS SUH 616) is an iron-based superalloy containing
20 % Co, 20 % Ni and 21 % Cr. It also contains smaller amounts of Mo, W, and Nb for improved
solution strengthening at elevated temperatures, but little or no precipitation strengthening is obtainalble.
N-155 is single phase austenitic and hardenable primarily by cold working or by hot-cold working and
has good forming, welding, and brazing characteristics. N-155 is used up to approximately 1500 °F (815
°C) and utilized primarily in low-stress applications up to approximately 2000 °F (1093°C), where
oxidation resistance is the most significant requirement [1]. The high contents of chromium and cobalt
contribute to good hot corrosion resistance. N-155 has been widely used in gas turbine engines, heat
exchangers, and other high temperature applications [2].
3.1.3.2 Chemical composition
The chemical compositions shown in Table 189 were reported by the manufacturers except for Al and B,
for which the analysis was carried out at NRIM [3]. The chemical requirements coincide with JIS G4311
SUH 661.
Materials preparation conditions are shown in Table 190. The bars were sampled in 1971.
Table 189. Chemical composition of Fe based 21Cr-20Ni-20Co-3Mo-2.5W-(Nb+Ta)-N alloy
NRIM
reference
code
Requirement
fFG
fFH
fFJ
C
Si
Mn
P
S
Chemical composition [wt-%]
Ni
Cr
Mo W
Co
0.080.16
0.14
0.12
0.13
≤
1.00
0.65
0.53
0.79
1.002.00
1.91
1.56
1.37
≤
0.040
0.005
0.006
0.012
≤
0.030
0.006
0.006
0.008
19.0021.00
20.16
20.25
20.77
20.0022.50
21.52
20.10
20.12
2.503.50
3.03
2.88
2.78
2.003.00
2.48
2.54
2.10
Al
B
N
18.500.1021.00
0.20
19.75 <0.01 0.001 0.161
19.35 0.049 <0.001 0.133
18.43 0.055 0.001 0.110
Table 190. Material preparation of Fe based 21Cr-20Ni-20Co-3Mo-2.5W-(Nb+Ta)-N alloy
NRIM
Type of Size of Deoxidation Product Dimen Processing and
Austenite
reference melting ingot
process
form
-sions thermal history
grain size
code
[kgf]
[mm]
number
fFG
150
Si-Al-killed
20 D
Forged
6.1
HFIF
150 L 1177 °C/1 h WQ
fFH
ESR
1100
25 D
Forged
5.8
5000 L 1150 °C/1 h WQ
Bar
800 °C/15 h AC
fFJ
HFIVF 100
C-killed
20 D
Forged
5.7
2740 - 1180 °C/0.5 h
3930 L WQ
816 °C/4 h AC
HFIF: high frequency induction furnace, ESR: electroslag remelting,
HFIVF: high frequency induction vacuum furnace
Landolt-Börnstein
New Series VIII/2B
Nb+
Ta
0.751.25
1.07
1.02
0.86
Rockwell
hardness
[HRB]
94
94
92
328
3.1 Fe-base alloys
3.1.3.3 Mechanical properties
The mechanical properties of Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings are shown in Fig. 420
[3].
0.2 % proof stress
900
900
800
800
700
700
600
upper 95% PI
average
lower 95% PI
500
400
500
400
300
200
200
100
100
0
200
400
600
Test temperature [°C]
800
0
1000
Elongation
100
90
90
80
80
70
70
60
50
40
20
10
10
400
600
Test temperature [°C]
800
1000
800
1000
Reduction of area
40
20
200
1000
50
30
0
800
60
30
0
0
upper 95% PI
average
lower 95% PI
200
400
600
Test temperature [°C]
100
Reduction of area [%]
Elongation [%]
600
300
0
Tensile strength
1000
Stress [MPa]
Stress [MPa]
1000
0
0
200
400
600
Test temperature [°C]
Fig. 420. Short-time tensile properties of Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings.
3.1.3.4 Creep and rupture properties
The creep rupture properties of Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings are shown in Fig. 421
to 424 [3].
Fig. 421 shows the stress vs. time to rupture at 550, 650, 750 and 850 °C for iron based 21Cr-20Ni20Co-Mo-(Nb+Ta)-N alloy forgings.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 330]
3.1.3 Fe-21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy
329
1000
500
Stress [MPa]
300
100
50
30
550 °C
650 °C
750 °C
850 °C
10
1
10
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 421. Creep rupture strength data for iron based 21Cr20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings.
Fig. 422 (next page) shows the elongation and reduction of area at 550, 650, 750 and 850 °C for iron
based 21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings.
Fig. 423 shows stress vs. minimum creep rate for iron based 21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy
forgings.
Fig. 424 shows time to rupture vs. minimum creep rate for iron based 21Cr-20Ni-20Co-Mo-(Nb+Ta)N alloy forgings.
1000
10 6
550 °C
650 °C
750 °C
850 °C
500
105
Time to rupture [h]
Stress [MPa]
300
550 °C
650 °C
750 °C
850 °C
100
50
30
104
103
102
10
10-6
10-5
10-4
10-3
10-2
10-1
1
Minimum creep rate [% / h ]
Fig. 423. Stress vs. minimum creep rate for iron based
21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings.
Landolt-Börnstein
New Series VIII/2B
10
10-6
10-5
10-4
10-3
10-2
10-1
1
Minimum creep rate [% /h]
Fig. 424. Time to rupture vs. minimum creep rate for iron
based 21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings.
330
3.1 Fe-base alloys
100
100
550 °C
550 °C
80
80
60
60
40
40
20
20
0
0
100
100
650 °C
650 °C
80
60
60
40
40
Elongation [%]
20
0
100
750 °C
80
Reduction of area [%]
80
20
0
100
750 °C
80
60
60
40
40
20
20
0
100
0
100
850 °C
850 °C
80
80
60
60
40
40
20
20
0
1
10
10 4
10 3
10 2
Time to rupture [h]
10 5
10 6
0
1
10
10 4
10 3
10 2
Time to rupture [h]
10 5
10 6
Fig. 422. Elongation and reduction of area at 550, 650, 750 and 850 °C for iron based
21Cr-20Ni-20Co-Mo-(Nb+Ta)-N alloy forgings.
3.1.3.5 References
[1] Superalloys II, High temperature materials for aerospace and industrial power (1987).
[2] Aerospace structural materials handbook, code 1602.
[3] NRIM Creep Data Sheet, No. 33A, Fe based 21Cr-20Ni-20Co-3Mo-2.5W-(Nb+Ta)-N Superalloy
for Gas Turbine Blades, 1999.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 335]
3.2.1 Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B alloy
331
3.2 Ni-base alloys
3.2.1 Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B alloy
3.2.1.1 Introduction
Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B (Inconel 713C, Alloy713C) alloy is one of the gamma prime
precipitation strengthened type nickel-base superalloys with most excellent high temperature creep
strength. Ti, Al, Nb and Ta which combine with Ni to form gamma prime contribute precipitation
strengthening and Cr and Mo which form carbides contribute to grain boundary strengthening [1]. This
alloy also has very good oxidation, hot corrosion and thermal shock resistance up to about 1000 °C.
Therefore it is in widespread use as turbine blades, vanes etc. [1, 2]. It is also used as connectors or pullrods in high temperature mechanical testing machines. Inconel 713C is usually produced by vacuum
melting and casting and used in as-cast condition.
3.2.1.2 Material standard and chemical composition
The chemical composition requirement given in ASM is shown in Table 191. The mechanical properties
of Inconel 713C reported here are given in [2], [3] and [4].
Table 191. Chemical composition requirement in AMS.
Chemical composition [wt%]
C
Mn
Si
S
Cr Mo Nb+Ta Ti Al
B
Zr
Min. 0.08
- 12.00 3.80 1.80 0.50 5.50 0.005 0.05
Max 0.20 0.25 0.50 0.015 14.00 5.20 2.80 1.00 6.50 0.015 0.15
AMS
Fe
2.5
Cu Co
Ni
balance
0.50 1.00
3.2.1.3 Tensile properties
0.2% proof stress and tensile strength of Inconel 713C alloy are shown in Fig. 425 [3, 4] and Fig. 426
[2-4].
3.2.1.4 Creep and rupture properties
Creep rupture curves of Inconel 713C alloy are shown in Fig. 427 [3], Fig. 428 [4] and Fig. 429 [2].
1,000, 10,000 and 30,000 h rupture-stresses as a function of temperatures are available in Fig. 430 [3, 4].
Minimum creep rate for Inconel 713C is reported as shown in Fig. 431 and Fig. 432 [2, 3].
Landolt-Börnstein
New Series VIII/2B
332
3.2 Ni-base alloys
Tensile strength
1000
800
800
600
600
Stress (MPa)
Stress (MPa)
0.2% proof stress
1000
400
200
0
400
200
0
200
400
600
800
1000
0
1200
0
200
Test temperature (℃)
400
600
800
1000
1200
Test temperature (℃)
Fig. 425. 0.2% proof stress and tensile strength of Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B superalloy; [3].
Fig. 427, see next page
Ni-13Cr-6Al-4Mo-2Cb-0.7Ti
as cast bar
80
70
60
Tensile stress
50
120
80
40
0
40
Stress [kg f / mm 2 ]
80
Yield stress [ksi]
Tensile stress [ksi]
120
750 °C
850 °C
40
30
20
Yield stress
1200
0
1400
1600
1800
Temperature [°F]
2000
Fig. 426. Tensile (circles) and yield stress (squares) of
Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B superalloy;
[2].
1 ksi = 6.89476 MPa
10
1
10
10 3
10 2
Time to rupture [h]
10 4
Fig. 428. Creep rupture curve of Ni-13Cr-4.5Mo-0.75Ti6Al-(Nb+Ta)-Zr-B; [4].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 335]
3.2.1 Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B alloy
333
500
850 °C
900 °C
950 °C
1000 °C
400
300
Stress [MPa]
200
100
80
850 °C
60
50
40
Fig. 427. Creep rupture strength
data of Ni-13Cr-4.5Mo-0.75Ti6Al-(Nb + Ta)-Zr-B superalloy
including 8 heats; [3]. n indicates
the total number of data points.
900 °C
950 °C
30
20
10
1000 °C
n = 187
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Ni-13 Cr-6 Al-4 Mo-2 Cb-0.7 Ti
vac melt, vac cast bar
100
Stress [ksi ]
1350 F
50
1500 F
20
1700 F
1800 F
10
5
10
2000 F
2
10 2
2
5 10 3 2
Rupture time [h]
5
5
10 4
Ni-13 Cr-6 Al-4 Mo-2 Cb-0.7 Ti
as-cast, various sources, vac melt, cast in vac or inert atm
100
Stress [ksi ]
1350 F
50
1500 F
1600 F
20
1700 F
1900 F
10
1800 F
Fig. 429. Creep rupture curves of Ni-13Cr-4.5Mo-0.75Ti6Al-(Nb+Ta)-Zr-B superalloy; [2]. 1 ksi = 6.89476 MPa
2000 F
5
10 2
Landolt-Börnstein
New Series VIII/2B
5 10 2 2 5 10 3 2
Rupture time [h]
5 10 4
334
3.2 Ni-base alloys
500
400
300
Stress [MPa]
200
100 h
100
80
60
50
40
10000 h
30
20
800
850
a
1000
900
950
Temperature [°C]
1050
500
400
300
Stress [MPa]
200
100
80
1000 h
60
50
40
30
Fig. 430. Temperature dependence of creep rupture strength
from 100 to 100,000 h; [3].
100000 h
20
800
850
b
1000
900
950
Temperature [°C]
1050
300
200
Stress [MPa]
100
100
80
60
50
40
30
850 °C
900 °C
950 °C 1000 °C
850 °C
900 °C
950 °C
1000 °C
n = 65
20
10 -7 10 -6 10 -5 10 -4 10 -3 10 -2 10 -1
Minimum creep rate [% / h]
Fig. 431. Stress vs. minimum creep rate for Ni-13Cr4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B superalloy; [3].
1
Landolt-Börnstein
New Series VIII/2B
Ref. p. 335]
3.2.1 Ni-13Cr-4.5Mo-0.75Ti-6Al-(Nb+Ta)-Zr-B alloy
335
Ni-13 Cr-6 Al-4 Mo-2 Cb-0.7 Ti
100 as-cast bar, vac melt, vac cast
1350 F
1500 F
Stress [ksi ]
50
1700 F
20
Minimum creep rate
10
10 − 4 2
5
10 − 3 2
5 10 − 2 2
Minimum creep rate [% / h ]
5
10 − 1
Fig. 432. Stress vs. minimum creep rate for Ni-13Cr 4.5Mo - 0.75Ti-6Al-(Nb + Ta) - Zr-B superalloy; [2]. 1 ksi
= 6.89476 MPa
3.2.1.5 Creep crack growth properties
Creep crack growth properties of Ni-10Cr-15Co-3Mo-5Ti-6Al-V-B (IN100) and Ni-13Cr-4.5Mo-0.75Ti-6Al(Nb+Ta)-Zr-B (Inconel 713C) superalloys obtained in NIMS using CT specimen [5] are shown in Fig. 433.
1
d a /d t [mm /h ]
10 1
10 2
IN 100, 732 °C
IN 100, 850 °C
Inconel 713 C, 900 °C
Inconel 713 C, 1000 °C
10 3
10 4
10 2
10 1
C * [kJ /m 2 h ]
1
10
Fig. 433. Creep crack growth rate vs. C* relations of Ni10Cr-15Co-3Mo-5Ti-6Al-V-B (IN100) and 13Cr-4.5Mo0.75Ti-6Al-(Nb+Ta)-Zr-B (Inconel 713C) superalloys;
[5].
3.2.1.6 References
[1]
[2]
[3]
[4]
The superalloys, ed. by Sims, C. T., and Hagel, W. C., John Wiley & Sons (1972).
Aerospace Structural Metals Handbook, vol.5, code 4119 (1976).
NRIM Creep Data Sheet, No.29B (1990).
Report on the mechanical properties of metals at elevated temperatures, Vol. IV Superalloys, The
Iron and Steel Institute of Japan (1979).
[5] Tabuchi, M., Kubo, K., Yagi, K., Yokobori, A. T. Jr., and Fuji, A., Eng. Fract. Mech., 62 (1999), 4760.
Landolt-Börnstein
New Series VIII/2B
336
3.2 Ni-base alloys
3.2.2 Ni-10Cr-15Co-3Mo-5Ti-6Al-V-B alloy
3.2.2.1 Introduction
Ni-10Cr-15Co-3Mo-5Ti-6Al-V-B alloy (IN-100 alloy) is a nickel-base precipitation hardenable, vacuum
cast superalloy possessing high rupture strength up to about 1040 °C. The high percentage of aluminum
and titanium and the low refractory metal content make IN-100 alloy particularly attractive on strength to
density basis. The alloy has been successfully cast and utilized in a variety of shapes from turbine blades,
vanes and nozzles to integral wheels.
Because of the high content of gamma prime precipitate that constitutes one of the strengthening
components of the alloy, the equilibrium solution temperature approaches the solidus, so the alloy is
usually used in the as-cast condition.
IN-100 alloy has been widely known as a cast superalloy, because the alloy was not hot worked due to
its high content of gamma prime precipitate. Recently, however, there has been considerable development
of a powder metallurgy product which permits working of the alloy. At high temperatures the powderconsolidated product has been found to show superplasticity, thus many possibilities in fabrication-toshape of wrought complex components can be available. However, because no public specification is
available for the powder-consolidated IN-100 alloy, only data for as-cast condition are contained in this
sheet.
3.2.2.2 Material standards, chemical and mechanical requirements
As with other high strength superalloys, deterioration of strength after long time exposure at elevated
temperatures is of major concern with IN-100 alloy. It has been widely accepted that electron vacancy
concentration is useful in indicating the susceptibility of an alloy to form σ phase [1]. The electron
vacancy number, Nv, of IN-100 alloy of nominal composition is 2.46. Nv values over 2.50 generally
indicate that an alloy is susceptible to σ phase formation.
When IN-100 alloy was originally introduced, the suggested range for titanium extended from 4.5 to a
maximum of 5.5 %. Compositions toward the top side of this range exhibited σ phase formation. For
example, a 5.3Ti alloy with Nv 2.70, contained σ phase which detracted from rupture life. The maximum
titanium level then was reduced to the AMS specification value of 5.0 %. This change eliminated the
deleterious effects of σ phase formation on material properties without sacrificing any properties desired
[1]. Also, because of the same reason, GE’s new spec. C50T77C specifies that the maximum titanium
content is 4.4 % while their original spec. C50T77A specifies it as 5.5 % [2].
Although AMS5397B was only a public specification for IN-100 alloys, please note that the
specification has been declared “NONCURRENT” by the Aerospace Materials Division, SAE, in
November 1995 [3].
Table 192. Chemical requirement of AMS5397B
Chemical composition [wt%]
C
Mn Si
P
S
Cr
Co
Mo Ti
Al
Ti+Al B
V
Zr
Fe
Ni
max. 0.20 0.10 0.10 0.015 0.015 11.00 17.00 4.00 5.00 6.00 11.00 0.02 1.20 0.09 1.00 remainder
min. 0.15
8.00 13.00 2.00 4.50 5.00 10.00 0.01 0.70 0.03
Table 193. Mechanical requirements of AMS5397B
Tensile properties
Condition
Min. tensile
Min. yield stress at
Min. elongation
strength
0.2 % offset
in 4D
as-cast
795 MPa
655 MPa
5%
Stress-rupture properties
Min. rupture Min. elongation
life
in 4D
23 h
4%
Landolt-Börnstein
New Series VIII/2B
Ref. p. 340]
3.2.2 Ni-10Cr-15Co-3Mo-5Ti-6Al-V-B alloy
337
3.2.2.3 Mechanical properties
Tensile properties of IN-100 alloy are shown in Table 194 (raw data) and Fig. 434 [1].
Table 194. Tensile properties of IN-100 alloy
Test temp. 0.2% yield strength Tensile strength
[°C]
[MPa]
[MPa]
21
849
1014
538
883
1090
649
890
1111
732
876
1097
816
814
994
927
504
738
1038
283
442
1200
Elongation
[%]
9.0
9.0
6.0
6.5
6.0
6.0
6.0
Reduction of area
[%]
11.0
11.0
7.0
7.2
7.2
7.2
8.0
20
Tensile Strength
18
1000
16
Reduction of Area
Ductility (%)
Stress (MPa)
14
0.2% Yield Strength
800
600
400
12
10
8
6
Elongation
4
200
2
0
0
0
200
400
600
800
1000
1200
Test temperature (OC)
0
200
400
600
800
1000
1200
Test temperature (OC)
Fig. 434. Typical tensile properties of as-cast IN-100 alloy.
3.2.2.4 Creep and rupture properties
Stress-rupture data for IN-100 alloy are shown in Fig. 435 and Table 195 (raw data) [1]. Coarse grain
(>1/8”) and fine grain (<1/16”) castings have exhibited nearly identical creep rupture lives. 10, 100 and
1000 h rupture-stresses as function of temperature are shown in Fig. 436 [1].
Landolt-Börnstein
New Series VIII/2B
338
3.2 Ni-base alloys
Table 195. Stress-rupture data on as-cast IN-100 alloy
Coarse grain (>1/8”)
Temp
Stress
Life [h] Elong. R. o. a.
Temp.
[°C]
[MPa]
[%]
[%]
[°C]
104
115
9
13
982
1038
83
278
9
17
62
705
11
14
927
200
45
11
12
173
82
11
12
899
982
90
1779
9
15
816
83
2440
12
23
345
15
7
7
927
131
3468
13
18
899
207
1322
6
11
587
16
6
10
816
518
82
5
6
380
806
9
9
760
587
202
6
690
52
2
5
732
621
355
3
6
552
1468
3
8
R.o.a. : Reduction of area
Fine grain (<1/16”)
Stress
Life [h] Elong.
[MPa]
[%]
200
43
14
124
526
12
242
183
11
173
1013
11
207
970
7
621
9
7
414
771
7
345
1548
8
R. o. a.
[%]
16
19
12
16
14
6
10
14
1000
732OC
Stress (MPa)
816OC
927OC
982OC
100
1038OC
IN-100 alloy
4.80% Ti
: Fine grain
: Coarse grain
Fig. 435. Creep rupture strength
data for as-cast IN-100 alloy.
10
1
10
100
1000
10000
Time (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 340]
3.2.2 Ni-10Cr-15Co-3Mo-5Ti-6Al-V-B alloy
339
1000
IN-100 alloy
4.80% Ti
800
Stress (MPa)
100h
10h
600
1000h
400
200
Fig. 436. Typical stress-rupture
properties of as-cast IN-100 alloy
0
700
800
900
1000
1100
Temperature ( C)
O
Long-time creep data on as-cast IN-100 alloy are shown in Table 196 (raw data) and Fig. 437 [1].
Table 196. Long-time creep data on as-cast IN-100 alloy
Coarse grain (>1/8”)
Temp. Stress Time for total creep strain of
[°C]
[MPa] [h]
0.1% 0.2% 0.5% 1.0%
732
816
927
982
1038
690
621
552
587
518
380
345
131
200
173
90
83
104
83
62
Landolt-Börnstein
New Series VIII/2B
Min.
creep
rate
[%/h]
1.0 2.0
8.0
26.0
0.0291
5.0 10.0 44.0
135.0 0.0050
25.0 175.0 500.0 800.0 0.0014
0.2
1.9
5.6
0.1300
2.0 2.5
7.7
28.0
0.0185
50.0 170.0 395.0 560.0 0.0010
0.4 1.8
2.9
5.8
0.0320
80.0 200.0 1000.0 1760.0 0.0004
1.5 7.0
16.0
25.0
0.0310
2.5 7.0
19.0
35.0
0.0350
60.0 175.0 625.0 1175.0 0.0006
75.0 195.0 750.0 1380.0 0.0005
2.7 6.0
25.0
58.0
0.0144
8.0 20.0 62.0
140.0 0.0060
4.0 15.0 90.0
270.0 0.0028
Fine grain (<1/16”)
Temp. Stress Time for total creep strain of
[°C]
[MPa] [h]
0.1% 0.2% 0.5% 1.0%
816
899
927
982
621
414
207
242
173
124
0.10
7.50
55.00
5.00
60.00
30.00
0.15
40.00
190.00
15.00
120.00
88.00
0.40
170.00
440.00
45.00
260.00
165.00
Min.
creep
rate
[%/h]
1.85
0.2800
340.00 0.0025
570.00 0.00091
79.00 0.0010
480.00 2850.00 0.0030
340
3.2 Ni-base alloys
750
0.5% creep strain
600
1.0% creep strain
816OC
732OC
700
550
Stress-rupture
Stress (MPa)
Stress (MPa)
650
600
550
0.1% creep strain
500
Stress-rupture
500
450
400
0.1% creep strain
0.2% creep strain
350
450
400
0.2% creep strain
0.5% creep strain 1.0% creep strain
300
1
10
100
1000
10000
1
10
Time (h)
400
927OC
10000
982OC
230
1.0% creep
210
190
300
Stress (MPa)
Stress (MPa)
1000
250
0.5% creep strain
1.0% creep strain
350
100
Time (h)
Stress-rupture
250
200
0.1% creep strain
150
1
Stress-rupture
150
130
110
0.1% creep strain
90
70
0.2% creep strain
100
170
10
100
0.2% creep strain 0.5% creep strain
50
1000
10000
Time (h)
1
10
100
1000
Time (h)
120
1038 C
110
1038OC
Stress (MPa)
100
90
Stress-rupture
80
70
60
0.5% creep strain
50
1.0% creep strain
40
1
10
100
1000
10000
Time (h)
Fig. 437. Creep curves for as-cast IN-100 alloy
3.2.2.5 References
[1] International Nickel Co., Engineering Properties of IN-100 Alloy.
[2] Aerospace Structural Metals Handbook, vol.5, code 4212 (1978).
[3] Aerospace Materials Division, SAE, Aerospace Material Specification, AMS5397C (1995).
Landolt-Börnstein
New Series VIII/2B
10000
Ref. p. 344]
3.2.3 Ni-15.5Cr-8Fe alloy
341
3.2.3 Ni-15.5Cr-8Fe alloy
3.2.3.1 Introduction
Ni-15.5Cr-8Fe (Inconel 600) alloy is one of the oxidation-resistant type superalloys based on the Ni-Cr
alloy system [1]. Because Inconel 600 has good workability, weldability and corrosion resistance, it is
extensively used for heat exchanger tubes in chemical and nuclear power plants. Because this alloy is
used as the steam generator tubing in nuclear power plants, there are many researches concerning not only
creep but also stress corrosion cracking (SCC) and fatigue properties. The protective Cr2O3 scale
contributes to corrosion-resistance and Fe is a solid solution strengthening element. This wrought alloy is
used after annealing at about 1000 °C.
3.2.3.2 Material standard and chemical composition
The chemical composition requirement of JIS [2-4] is shown in Table 197, which is nearly the same as in
ISO4955. The mechanical properties of Inconel 600 are reported in [5] for bars, plates and tubes, in [6]
and [7].
Table 197. Chemical composition requirement in JIS G 4904
Chemical composition [wt%]
C
Si
Mn P
S
Ni
Cr
F
Cu
NCF600TB ≤0.15 ≤0.50 ≤1.00 ≤0.030 ≤0.015 ≥72.00 14.00-17.00 6.00-10.00 ≤0.50
JIS
G 4904
3.2.3.3 Tensile properties
0.2% proof stress and tensile strength of Inconel 600 alloy are shown in Fig. 438 [5]. The yield strength
and tensile strength are shown in Fig. 439 [6].
Tensile strength
800
700
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
800
400
300
400
300
200
200
100
100
0
0
200
400
600
800
1000
1200
0
0
200
Test temperature (℃)
Fig. 438. 0.2% proof stress and tensile strength of Ni-15.5Cr-8Fe alloy; [5].
Landolt-Börnstein
New Series VIII/2B
400
600
800
Test temperature (℃)
1000
1200
342
3.2 Ni-base alloys
120
100
Tensile stress
60
40
Yield stress
20
120
80
0
40
Elongation
Elongation [%]
Stress [ksi]
80
0
-400
0
800
1200
400
Temperature [°F]
Fig. 439. Yield and tensile strength of Inconel 600 alloy;
[6]. 1 ksi = 6.89476 MPa
1600
3.2.3.4 Creep and rupture properties
Stress vs. creep rupture time relations of Inconel 600 alloy are shown in Fig. 440 [5]. Creep strain and
rupture data is shown in Fig. 441 [6]. 1,000, 10,000 and 30,000 h rupture-stresses as functions of
temperature are shown in Fig. 442 [5].
500
400
300
600 °C
700 °C
800 °C
900 °C
1000 °C
Stress [MPa]
200
100
80
60
50
40
30
20
10
8
10
n = 103
10 2
10 3
10 4
Time to rupture [h]
10 5
10 6
Fig. 440. Creep rupture strength data of Inconel 600
alloys including bars, plates and tubes; [5]. n indicates the
total number of data points.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 344]
3.2.3 Ni-15.5Cr-8Fe alloy
343
Fig. 441, see next page
800
600
500
400
300
100
80
60
50
40
30
{
{
Stress [MPa]
200
Tensile
strength
0.2% proof
stress
20
1000 h
10
8
a
6
800
600
500
400
300
100
80
60
50
40
30
{
Tensile
strength
{
0.2% proof
stress
{
Stress [MPa]
200
Tensile
strength
20
10
8
b
6
800
600
500
400
300
10000 h
100
80
60
50
40
30
{
Stress [MPa]
200
0.2% proof
stress
20
10
8
6
500
c
Landolt-Börnstein
New Series VIII/2B
Fig. 442. Temperature dependence of creep rupture strength
from 1,000 to 30,000 h; [5].
30000 h
600
700
800
Temperature [°C]
900
1000
1100
344
3.2 Ni-base alloys
20
Inconel alloy 600
0.060 in sheet
2050 F, 2 h
Stress [ksi ]
10
5
1500 F
2
1
10
0.5
5 10
Rupture
10
5
1300 F
2
1
0.5
5
Elongation%
Rupture
1650 F
2
Total creep and
rupture
Tested in
argon
1
20
0.5
1
Inconel alloy 600
0.060 in sheet
CW 20 % + 1900 F,
41/2 min
1300 F
Stress [ksi ]
10
5
2
0.5 1 2
0.5 1
2
5
2
10
R
0.5 1
2 5
10
Rupture
e, percent
5
1500 F
1650 F
Total creep and
rupture
10
Tested in
argon
Rupture
1
10
10 2
Time [h]
10 3
Fig. 441. Creep strain and rupture curves of Inconel 600
alloy; [6]. 1 ksi = 6.89476 MPa
10 4
3.2.3.5 References
[1]
[2]
[3]
[4]
[5]
[6]
[7]
The superalloys, ed. by Sims, C. T., and Hagel, W. C., John Wiley & Sons (1972) p. 294.
JIS G4901-1991, Corrosion-resisting and heat-resisting superalloy bars.
JIS G4902-1991, Corrosion-resisting and heat-resisting superalloy plates and sheets.
JIS G4904-1991, Seamless nickel-chromium-iron alloy heat exchanger tubes.
NRIM Creep Data Sheet, No.41A (1999).
Aerospace Structural Metals Handbook, vol.4, code 4101 (1967).
Report on the mechanical properties of metals at elevated temperatures, Vol. IV Superalloys, The
Iron and Steel Institute of Japan (1979).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 347]
3.2.4 Ni-15.5Cr-2.5Ti-0.7Al-1Nb-7Fe alloy
345
3.2.4 Ni-15.5Cr-2.5Ti-0.7Al-1Nb-7Fe alloy
3.2.4.1 Introduction
Ni-15.5Cr-2.5Ti-0.7Al-1Nb-7Fe (Inconel X-750, JIS NCF750, ISO NW7750) is one of the precipitation
hardened nickel chromium alloys. These alloy families were developed from the solid-solution
strengthened Ni-15.5Cr-8Fe alloy (Inconel 600) by addition of titanium and aluminum to form gamma
prime phase for precipitation strengthening [1, 2]. Addition of 1 % Nb gives Inconel X-750. Because
Inconel X-750 has not only high creep strength but also good formability, weldability and corrosion
resistance, it is widely used for aerospace applications such as rotors, blades, wheels and bolts etc. up to
about 1000 °C [2]. It has also excellent properties at cryogenic temperatures, therefore it is used for bolts,
springs etc. at various temperatures. This wrought alloy is normally used in the precipitation hardened
condition after solid-solution at about 1150 °C and ageing heat treatment whose condition is dependent
on desired properties of products.
3.2.4.2 Material standard and chemical composition
The chemical composition requirement given in JIS G 4901 [3] is shown in Table 1, which is nearly the
same as ISO 9723. The mechanical properties of Inconel X-750 alloy can be obtained in NRIM Creep
Data Sheet 39A [4], Aerospace structural metals handbook [2] and the report on the mechanical
properties of metals at elevated temperatures [5].
Table 198. Chemical composition requirement in JIS G 4901
Standard
Chemical composition [wt%]
C
Si
Mn P
S
Ni
Cr
Cu
Ti
Al
Nb+Ta
JIS G 4901
14.002.25- 0.40- 0.70≤0.50
≤0.08 ≤0.50 ≤1.00 ≤0.030 ≤0.015 ≥70.00
17.00
2.75 1.00 1.20
Fe
5.009.00
3.2.4.3 Tensile properties
0.2% proof stress and tensile strength of Inconel X-750 are shown in Fig. 443 [4].
Tensile strength
1200
1000
1000
800
800
Stress (MPa)
Stress (MPa)
0.2% proof stress
1200
600
400
400
200
200
0
600
0
200
400
600
800
1000
0
0
Test temperature (℃)
Fig. 443. 0.2% proof stress and tensile strength of Inconel X-750 alloy; [4].
Landolt-Börnstein
New Series VIII/2B
200
400
600
Test temperature (℃)
800
1000
346
3.2 Ni-base alloys
3.2.4.4 Creep and rupture properties
Stress vs. creep rupture relations of Inconel X-750 alloy are shown in Fig. 444 [4] and Fig 445 [2]. The
creep rupture stresses from 100 to 30,000 h as functions of temperature are given in Fig. 446 [4]. The
relation between stress and minimum creep rate is given in Fig. 447 [2].
Stress [MPa]
900
700
600
500
400
300
600 °C
650 °C
700 °C
750 °C
800 °C
850 °C
900 °C
200
100
70
60
50
40
30
Fig. 444. Creep rupture strength
data of Ni based 15.5Cr-2.5Ti0.7Al-1Nb-7Fe superalloy bars;
[4]. n indicates the total number of
data points.
n = 100
20
1
10 2
10
10 3
10 4
Time to rupture [h]
10 5
10 6
10 7
Fig. 446, see next page
IN X-750 bar
2100 F, 2h, AC+1550 F,
24 h, Ac+1300 F, 20 h, AC
2
1000 F
1100 F
1000 F
1100 F
100
1350 F
Stress [ksi ]
10
Stress [ksi ]
IN X-750 bar
1625 F, 24 h, AC+1300 F,
20 h, Ac
1200 F
50
10
1500 F
1800 F
1
10 -1
1700 F
1350 F
1600 F
1500 F
10
1
10 3
10 2
10
Rupture time [h]
10 4
10 5
Fig. 445. Creep rupture curves of Inconel X-750
alloy bar; [2]. 1 ksi = 6.89476 MPa
1
10
10 2
3
Creep rate [ %/10 h]
Fig. 447. Stress vs. minimum creep rate for Inconel X-750
superalloy; [2]. 1 ksi = 6.89476 MPa
Landolt-Börnstein
New Series VIII/2B
Ref. p. 347]
3.2.4 Ni-15.5Cr-2.5Ti-0.7Al-1Nb-7Fe alloy
347
2000
Tensile
strength
{
100 h
0.2%
proof
stress
{
100
80
60
50
40
30
a 20
Tensile
strength
{
200
{
Stress [MPa]
1000
800
600
500
400
300
0.2%
proof
stress
10000 h
2000
Stress [MPa]
1000
800
600
500
400
300
200
100
80
60
50
40
30
20
550
1000 h
30000 h
600
650
b
700
750
800
Temperature [°C]
850
900
950
1000
Fig. 446. Temperature dependence of creep rupture stress from
100 to 30000 h; [4].
3.2.4.5 References
[1]
[2]
[3]
[4]
[5]
The superalloys, ed. by Sims, C. T., and Hagel, W. C., John Wiley & Sons (1972) p. 16.
Aerospace Structural Metals Handbook, vol.4, code 4105 (1981).
JIS G 4901-1999, Corrosion-resisting and heat-resisting superalloy bars.
NRIM Creep Data Sheet, No. 39A (1992).
Report on the mechanical properties of metals at elevated temperatures, Vol.IV Superalloys, The
Iron and Steel Institute of Japan (1979) pp. 219-228.
Landolt-Börnstein
New Series VIII/2B
348
3.2 Ni-base alloys
3.2.5 Ni-15Cr-28Co-Mo-Ti-Al alloy
3.2.5.1 Introduction
Nickel based 15Cr-28Co-4Mo-2.5Ti-3Al alloy is one of the γ’ strengthened superalloys. This alloy has
been developed for high temperature service in the 1940s. The commercial designation of this alloy is
“Inconel Alloy 700”. The alloy is available in form of bars and forgings. Chemical properties are similar
to Waspaloy [1].
3.2.5.2 Chemical composition and material preparation
The conditions of material preparation are shown in Table 199. The chemical compositions shown in
Table 200 were analyzed by NRIM [2] and reported by the manufactures. The chemical requirements
coincide with Inconel 700.
Table 199. Details of Nickel based 15Cr-28Co-4Mo-2.5Ti-3Al superalloy bars (1)
NRIM
Type of Size of ingot Product Dimensions(3) Processing and
reference melting(2)
thermal history
form
code
[mm]
[t]
iBA
CE
0.3
20D
Forged
1600L
Bar
1200 °C/2 h AC
iBB
HFIVF
0.5
20D
870 °C/24 h AC
140L
Austenite
grain size
number(4)
2.4
Rockwell
hardness
[HRC]
36
2.4
36
(1) The bars were sampled in 1969. Details other than grain size number and hardness are as reported by the
manufactures.
(2) CE: consumable electrode vacuum melt, HFIVF: high frequency induction vacuum furnace.
(3) D: diameter, L: length
(4) JIS G 0551-1977, “Method of Austenite Grain Size Test for Steel’’
Table 200. Chemical composition (product analysis) of Nickel based 15Cr-28Co-4Mo-2.5Ti-3Al
superalloy bars
Chemical composition [wt-%](1)
S
Ni
Cr
Mo Cu
Co Ti
(3)
≤0.015 bal 13.0 1.0 ≤0.50 24.0 1.75
~17.0 ~4.5
~34.0 ~2.75
0.13 0.09 0.09 0.002 0.003 46.90 14.40 3.74 0.01 28.88 2.36
0.14 0.12 0.08 0.001 <0.005 bal(3) 15.13 3.63 tr(3) 27.91 2.26
NRIM
reference code C
Si
Mn P
Requirement(2) ≤0.20 ≤1.00 ≤2.00
iBA
iBB
Al
B
N*
2.50
~3.50
3.07 0.0059 0.0039
2.99 0.005 0.0030
Fe
≤4.0
0.24
0.72
(1) The chemical composition given above was reported by the manufactures except for the elements marked with an
asterisk, for which the analysis was carried out at NRIM.
(2) The chemical requirements coincide with Inconel 700 and with AISI 687.
(3) bal: balance, tr: trace
3.2.5.3 Mechanical properties
The tensile properties of Nickel based 15Cr-28Co-4Mo-2.5Ti-3Al alloy are shown in Fig. 448 which is
obtained from [3].
The solid curves represent the average; the broken curves are the upper and lower 95% PI (prediction
intervals).
Landolt-Börnstein
New Series VIII/2B
Ref. p. 350]
3.2.5 Ni-15Cr-28Co-Mo-Ti-Al alloy
Tensile strength
1600
1400
1400
1200
1200
1000
1000
Stress [MPa ]
Stress [MPa ]
0.2 % proof stress
1600
800
600
Upper 95 % PI
Average
Lower 95 % PI
400
200
0
600
0
Elongation
100
70
60
50
40
30
n = 24
Reduction of area
90
Reduction of area [%]
80
Upper 95 % PI
Average
Lower 95 % PI
200
90
Elongation [%]
800
400
n = 24
100
349
80
70
60
50
40
30
20
20
10
n = 24
0
0
200 400 600 800 1000
Test temperature [°C]
10
n = 24
0
0
200 400 600 800 1000
Test temperature [°C]
Fig. 448. Short-time tensile
properties of Nickel based
15Cr-28Co-4Mo-2.5Ti-3Al
superalloy bars.
3.2.5.4 Creep rupture properties
The creep rupture properties of Nickel based 15Cr-28Co-4Mo-2.5Ti-3Al alloy bars are shown in Fig. 449
and Fig. 450 which are obtained from [2].
Fig. 449 shows stress vs. time to rupture data at 700, 725, 750, 800, 825 and 850 °C for nickel based
15Cr-28Co-4Mo-2.5Ti-3Al alloy bars. Fig. 450 shows the master rupture curve assessed by MansonHaferd parameter method.
Landolt-Börnstein
New Series VIII/2B
350
3.2 Ni-base alloys
800
300
700 °C
725 °C
750 °C
800 °C
825 °C
850 °C
200
700 °C
Stress [MPa]
600
500
400
750 °C
100
80
60
50
40
10
Fig. 449. Creep rupture strebgth
data for Nickel based 15Cr- 28Co
- 4Mo - 2.5Ti - 3Al superalloy
bars. n indicates the total number
of data points.
800 °C
850 °C
n = 62
10 2
10 3
10 4
Time to rupture [h]
825 °C
10 5
10 6
800
600
500
400
Stress [MPa]
300
700 °C
725 °C
750 °C
800 °C
825 °C
850 °C
200
100
80
60
average
n = 62
50
40
-2.7
-2.5
-1.9
-1.7
-2.3
-2.1
Manson-Haferd parameter [( log tR -26.696)/( TK )] [×10 - 2 ]
Fig. 450. Master rupture curve by Manson-Haferd
parameter method for nickel based 15Cr-28Co-4Mo2.5Ti-3Al superalloy bars. n indicates the total number of
data points.
The creep data including time to reach specific total strain and minimum creep rate of nickel based 15Cr28Co-4Mo-2.5Ti-3Al alloy bars are available from [2].
3.2.5.5 References
[1] Aerospace Structural Metals Handbook, Volume I, Non-ferrous Alloys, 1963, Code 4201, p.1
[2] National Research Institute for Metals: NRIM Creep Data Sheet, No.24B, (1989).
[3] NRIM Creep Data Sheet, No.23B, 1989
Landolt-Börnstein
New Series VIII/2B
Ref. p. 355]
3.2.6 Ni-19Cr-18Co-4Mo-3Ti-3Al-B alloy
351
3.2.6 Ni-19Cr-18Co-4Mo-3Ti-3Al-B alloy
3.2.6.1 Introduction
This material is a Nickel based U500 superalloy (19Cr-18Co-4Mo-3Ti-3Al-B-Bal. Ni) which is applied
for gas turbine blades etc. and is manufactured in casting or forging form. It has shown excellent
resistance to hot corrosion attack and structural stability for long time service requirements. It is used at
high temperatures up to 1000 °C.
3.2.6.2 Material standards, chemical compositions and tensile properties
Information can be obtained from [1].
3.2.6.2.1 Nickel based 19Cr-18Co-4Mo-3Ti-3Al-B superalloy castings
Table 201 shows a requirement range for the chemical composition of Nickel based 19Cr-18Co-4Mo-3Ti3Al-B superalloy castings (U500) which coincides with the requirements of AMS5384 and U500.
Moreover, the chemical composition analysis results of the three heats used for data treatment are shown
in Table 201. These analysis results are almost equivalent. Table 202 shows the heat treatment conditions
of each heat. The treatment conditions of solid solution, high temperature aging and final aging are almost
equal for each heat.
Table 201. Requirement of chemical composition and analysis results for U500 casting alloy.
Chemical composition [wt %]
S
Cr
Mo Cu
Co
≤0.015 16.0- 3.0- ≤0.10 16.020.0 5.0
20.0
≤0.002 ≤0.005 19.51 4.12 trace 18.89
0.002 0.008 18.5 4.14 0.03 18.15
18.57 4.4 ≤0.05 18.81
≤0.002 0.01
C
Si
Mn
P
≤0.10 ≤0.30 ≤0.20 -
Requirement*
Heat 1
0.07
Heat 2
0.09
Heat 3
0.09
0.02
0.12
0.13
trace
0.07
0.12
Ti
2.503.25
2.88
3.16
2.93
Al
2.503.25
3.13
2.96
3.04
B
0.0030.010
0.005
0.009
0.01
N
-
Fe
≤2.0
0.004 0.2
0.006 0.23
0.001 0.12
* The chemical requirements coincide with AMS5384
Table 202. Heat treatment history for U500 casting alloy.
Thermal history
Heat 1 1149 °C × 4 h AC, 1080 °C × 4 h AC, 760 °C × 16 h AC
Heat 2 1149 °C × 4 h AC, 1079 °C × 4 h AC, 760 °C × 16 h AC
Heat 3 1149 °C × 4 h AC, 1080 °C × 4 h AC, 760 °C × 16 h AC
Table 203 shows the requirements of tensile strength and tensile elongation at 649 °C. Fig. 451 (a)-(d)
show the tensile properties from room temperature up to 1000 °C. The tensile strength has the tendency to
slightly decrease up to about 600 °C and then to increase near 700 °C. After that the strength falls rapidly
at high temperatures. 0.2% proof stress shows a similar tendency.
The tensile elongation ranges between 5 and 20 % from room temperature up to 900 °C and the
temperature dependence is not large. However, the elongation at 1000 °C is 15 to 20 %. The reduction of
area differs largely between the test materials and a remarkable variation is seen. Although there is a
slight decrease of reduction of area between 700 and 800 °C, the reduction of area indicates rapid
increasing at temperatures above 800 °C. This reversely corresponds well with the tendency for tensile
strength.
Landolt-Börnstein
New Series VIII/2B
352
3.2 Ni-base alloys
1400
1400
1200
1200
0.2% Proof stress (MPa)
Tensile strength (MPa)
Table 203. Requirement of tensile properties at 649 °C for U500 casting alloy.
Tensile properties (at 649 °C)
Tensile strength 0.2% proof stress Elongation Reduction of area
[MPa]
[MPa]
[%]
[%]
Requirement ≥827
≥7
1000
800
600
1000
800
600
400
400
200
200
0
0
0
200
400
600
800
0
1000
200
70
70
60
60
Reduction of area (%)
Elongation (%)
600
800
1000
800
1000
b
a
50
40
30
50
40
30
20
20
10
10
0
0
200
400
600
800
0
1000
0
200
Temperatura (℃)
c
400
Temperature (℃)
Temperature (℃)
400
600
Temperature (℃)
d
Fig. 451. Tensile strength (a), 0.2% proof stress (b), tensile elongation (c) and reduction of area (d) for U500 casting
alloy.
3.2.6.2.2 Nickel based 19Cr-18Co-4Mo-3Ti-3Al-B superalloy forgings
Table 204 shows a requirement range for the chemical composition of Nickel based 19Cr-18Co-4Mo-3Ti3Al-B superalloy forgings (U500) which coincides with the requirements of AMS5751A and U500.
Moreover, the chemical composition analysis results of the three heats used for data treatment are shown
in Table 204. These analysis results are almost equivalent except for the higher Fe content in heat 3
compared to the other two heats. There is no large difference of chemical composition between castings
and forgings.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 355]
3.2.6 Ni-19Cr-18Co-4Mo-3Ti-3Al-B alloy
353
Table 204. Requirement of chemical composition and the analysis results for U500 forging alloy.
Requirement*
Heat 1
Heat 2
Heat 3
Chemical composition [wt %]
S
Cr
Mo Cu
Co
≤0.015 15.0- 3.0- ≤0.15 13.020.0 5.0
20.0
0.004 0.002 0.01
19.23 4.13 trace 18.47
0.02 0.001 0.003 16.86 3.78 trace 18.2
0.002 0.002 0.009 19.25 4.12 0.01 18.8
C
Si
Mn
P
≤0.15 ≤0.75 ≤0.75 0.07
0.06
0.08
0.01
0.01
0.02
Ti
2.503.25
3.02
2.76
2.86
Al
2.503.25
3.1
2.66
2.81
B
0.0030.010
0.009
0.007
0.0065
N
-
Fe
≤4.0
0.005 0.22
0.006 0.73
0.007 2.32
* The chemical requirements coincide with AMS5384
Table 205 shows the heat treatment conditions of each heat. The treatment conditions of solid solution,
high temperature aging and final aging are almost equal for each heat. Compared with castings, a 170 °C
lower solid solution temperature is applied for forgings.
Table 205. Heat treatment history for U500 forging alloy.
Thermal history
Heat 1 1180 °C × 4 h AC, 843 °C × 24 h AC, 760 °C × 16 h AC
Heat 2 1179 °C × 4 h AC, 843 °C × 24 h AC, 760 °C × 16 h AC
Heat 3 1179 °C × 4 h AC, 843 °C × 24 h AC, 760 °C × 16 h AC
Table 206 shows the requirements of tensile properties at 649 °C. Compared with the castings in Table
203, the minimum requirement of tensile strength is about 350 MPa higher. Fig. 452 (a)-(d) show the
tensile properties of forgings from room temperature up to 1000 °C. Although the tensile strength
decreases slightly from room temperature up to 200 °C, it is almost constant up to near 700 °C. At higher
temperatures, the tensile strength decreases rapidly. Although the forgings are 200 to 300 MPa higher in
tensile strength than the castings from room temperature up to about 800 °C, at temperatures above
900 °C, a big difference is no longer seen between castings and forgings.
The 0.2% proof stress stays almost constant from room temperature up to near 700 °C and rapidly
decreases at temperatures above 700°C. The difference of proof stress between castings and forgings is
smaller than that of tensile strength. However, 0.2% proof stress of the forgings is about 100 MPa higher
than that of the castings up to near 700 °C.
Although the tensile elongation is about 10 %, like for castings, up to near 600 °C, it rapidly increases
at higher temperatures. The reduction of area shows a similar tendency.
Table 206. Requirement of tensile properties at 649 °C for U500 forging alloy.
Tensile properties (at 649 °C)
Tensile strength 0.2% proof stress Elongation Reduction of area
[MPa]
[MPa]
[%]
[%]
Requirement ≥1172
≥758
≥6
≥10
3.2.6.3 Creep rupture properties of Ni-19Cr-18Co-4Mo-3Ti-3Al-B alloy
3.2.6.3.1 Nickel based 19Cr-18Co-4Mo-3Ti-3Al-B superalloy castings
Fig. 453 shows the creep rupture test results of superalloy U500 castings. The creep tests were carried out
in the temperature range from 700 °C to 900 °C in 50 °C pitch and 1000 °C was also selected. The creep
stresses were adopted between 29 MPa and 608 MPa.
Landolt-Börnstein
New Series VIII/2B
354
3.2 Ni-base alloys
1400
1200
1200
0.2 % proof stress [MPa]
Tensile strength (MPa)
1400
1000
800
600
400
1000
800
600
400
200
0
200
0
200
400
600
800
0
1000
200
70
70
60
60
Reduction of area (%)
Elongation (%)
600
800
1000
800
1000
b
a
50
40
30
50
40
30
20
20
10
10
0
0
0
200
400
600
800
1000
0
200
Temperature (℃)
c
400
Temperature (℃)
Temperature (℃)
400
600
Temperature (℃)
d
Fig. 452. Tensile strength (a), 0.2% proof stress (b), tensile elongation (c) and reduction of area (d) for U500 forging
alloy.
3.2.6.3.2 Nickel based 19Cr-18Co-4Mo-3Ti-3Al-B superalloy forgings
Fig. 454 shows the creep rupture test results of superalloy U500 forgings. The creep tests were carried out
in the temperature range from 700 °C to 900 °C in 50 °C pitch and 1000 °C was also selected. The creep
stresses were adopted between 24 MPa and 608 MPa.
Comparison of Fig. 453 and Fig. 454 shows that the difference of creep rupture strength for forgings
and castings is low at temperatures of 850 °C or less. However, the decrease of the creep rupture strength
of forgings is larger than that of castings at temperatures above 900 °C.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 355]
3.2.6 Ni-19Cr-18Co-4Mo-3Ti-3Al-B alloy
355
Stress (MPa)
1000
100
T=700℃
T=750℃
T=800℃
T=850℃
T=900℃
T=1000℃
10
1
10
100
1000
104
105
Fig. 453. Creep rupture strength
data for U500 casting alloy.
Time to rupture (h)
Stress (MPa)
1000
100
T=700℃
T=750℃
T=800℃
T=850℃
T=900℃
T=1000℃
Fig. 454. Creep rupture strength
data for U500 forging alloy.
10
1
10
100
1000
104
105
Time to rupture (h)
3.2.6.4 Reference
[1] National Research Institute for Metals : NRIM Creep Data Sheet, No.34B, (1993).
Landolt-Börnstein
New Series VIII/2B
356
3.3 Co-base alloys
3.3 Co-base alloys
3.3.1 Co-20Cr-20Ni-Mo-W-Ni alloy
3.3.1.1 Introduction
This material is Co-based S-816 superalloy (20Cr-20Ni-4Mo-4W-4Nb-4 Fe-Bal. Co) which is applied to
blades, bolts, springs etc. for gas turbines. It is superior in corrosion resistance and is used in the high
temperature region up to about 900 °C. Although it has been manufactured by casting or rolling
processes, only data of rolling material is dealt with in this data book. The following heat treatments have
been performed: solid solution treatment or solid solution treatment with subsequent aging treatment.
3.3.1.2 Material standards and chemical compositions
This information on this alloy is obtained from [1].
Table 207 shows the typical chemical composition and the analysis results for two heats of S-816
superalloy (AISI 671). Both heats were almost equivalent in chemical composition. Table 208 shows the
typical heat treatment conditions and the thermal history of the two heats. Although the solid solution
treatment was performed at 1180 °C for all test materials, for the aging process, either no aging, 760 °C or
816 °C has been applied.
Table 207. Typical chemical composition and the analysis results for S-816 alloy.
Chemical composition [wt%]
Standard/
heat
C
Si
Mn P
S
Ni Cr Mo Co Nb W
AISI (S-816) 0.38 0.7 1.5 20 20 4
43 4 4
Nominal
Heat 1
0.37 0.37 1.31 0.006 0.019 19.6 20.4 4.3 42.7 3.6 4.2
Heat 2
0.35 0.64 1.58 0.004 0.017 19.7 20.1 4
42.2 3.9 4.3
Table 208. Typical heat treatment conditions and the thermal history for S-816 alloy.
Heat treatment
Standard/
heat
Solution treatment
Aging
AISI (S-816) 1180 °C × 1 h WQ 760 °C × 16 h AC
Nominal
1180 °C × 1 h WQ Heat 1
1180 °C × 1 h WQ 816 °C × 14 h AC
1180 °C × 1 h WQ 760 °C × 14 h AC
Heat 2
1180 °C × 1.5 h WQ 760 °C × 16 h AC
3.3.1.3 Creep rupture properties of alloy S-816
The influence of heat and aging condition on the creep rupture strength is shown in this chapter.
Fig. 455 shows all the creep rupture test results of superalloy S-816 dealt with in this data book. The
creep tests were carried out at temperatures from 600 °C to 850 °C in 50 °C pitch and at stresses which
were selected between 78 MPa and 785 MPa.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 361]
3.3.1 Co-20Cr-20Ni-Mo-W-Ni alloy
357
Stress (MPa)
1000
100
T=600℃
T=700℃
T=750℃
T=800℃
T=850℃
Fig. 455. Creep rupture strength
data for Co-20Cr-20Ni-Mo-W-Ni
alloy S-816.
10
10
100
1000
104
Time to rupture (h)
Fig. 456 (a) - (c) show the creep rupture strength of Heat 1 which is either not aged or aged at 760 °C or
816 °C. Fig. 457 shows creep rupture strength of Heat 2. As for Heat 1, the creep rupture data are
obtained only from the 760 °C aged material.
Stress (MPa)
1000
100
Fig. 456 a.
Creep rupture
strength data of Heat 1 of Co20Cr-20Ni-Mo-W-Ni alloy S-816
without aging treatment.
T=600℃
T=700℃
T=800℃
T=850℃
10
100
1000
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
104
358
3.3 Co-base alloys
Stress (MPa)
1000
100
T=600℃
T=700℃
T=800℃
T=850℃
10
Fig. 456 b.
Creep rupture
strength data of Heat 1 of Co20Cr-20Ni-Mo-W-Ni alloy S-816
aged at 760 °C.
100
1000
104
Time to rupture (h)
Stress (MPa)
1000
100
T=600℃
T=700℃
T=800℃
T=850℃
10
Fig. 456 c. Creep rupture strength
of Heat 1 data of Co-20Cr-20NiMo-W-Ni alloy S-816 aged at
816 °C.
100
1000
104
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
Ref. p. 361]
3.3.1 Co-20Cr-20Ni-Mo-W-Ni alloy
359
Stress (MPa)
1000
100
T=700℃
T=750℃
T=800℃
10
Fig. 457. Creep rupture strength
data of Heat 2 of Co-20Cr-20NiMo-W-Ni alloy S-816.
100
1000
104
Time to rupture (h)
All creep rupture test results were compared and the influence of heat and aging treatment condition were
summarized in Fig. 458. Fig. 458 a shows the influence of aging treatment condition on test results at
600 °C for Heat 1. Although there is some uncertainty due to data scatter, the decrease in strength for
longer times becomes smaller by aging treatment. Moreover, by increasing aging temperature from
760 °C to 816 °C, the decrease in creep rupture strength for longer times becomes smaller.
[
]
1000
900
Stress (MPa)
800
700
600
500
Heat1, S.T.=1180℃, Aging Less
Heat1, S.T.=1180℃, Aging=816℃
Heat1, S.T.=1180℃, Aging=760℃
400
Fig. 458 a. Influence of heat and
aging treatment condition on test
results for Co-20Cr-20Ni-Mo-WNi alloy S-816 at 600 °C.
300
10
100
1000
104
Time to rupture (h)
Fig. 458 b shows the influence of aging treatment condition on test results at 700 °C for Heat 1 and Heat
2. Because data has large scatter, the influence of aging treatment and aging temperature on the creep
rupture strength is not clear. The creep rupture strength is influenced by the difference of heat, and Heat 2
is lower than Heat 1 in creep rupture strength.
Landolt-Börnstein
New Series VIII/2B
360
3.3 Co-base alloys
700
600
Stress (MPa)
500
400
300
200
Heat1,
Heat1,
Heat1,
Heat2,
S.T.=1180℃,
S.T.=1180℃,
S.T.=1180℃,
S.T.=1180℃,
Aging Less
Aging=816℃
Aging=760℃
Aging=760℃
Fig. 458 b. Influence of heat and
aging treatment condition on test
results for Co-20Cr-20Ni-Mo-WNi alloy S-816 at 700 °C.
100
10
100
1000
104
Time to rupture (h)
Fig. 458 c similarly shows the results of creep rupture strength at 800 °C. The same tendency as in the
case of 700 °C can be seen.
Stress (MPa)
500
300
100
80
Heat1,
Heat1,
Heat1,
Heat2,
60
S.T.=1180℃,
S.T.=1180℃,
S.T.=1180℃,
S.T.=1180℃,
Aging Less
Aging=816℃
Aging=760℃
Aging=760℃
Fig. 458 c. Influence of heat and
aging treatment condition on test
results for Co-20Cr-20Ni-Mo-WNi alloy S-816 at 800 °C.
40
10
100
1000
104
Time to rupture (h)
Fig. 458 d is the result of creep rupture strength at 850 °C. Also in this case, the existence of aging
treatment and the effect of aging temperature on creep rupture strength is difficult to assess.
Landolt-Börnstein
New Series VIII/2B
Ref. p. 361]
3.3.1 Co-20Cr-20Ni-Mo-W-Ni alloy
361
Stress (MPa)
300
100
80
60
40
Heat1, S.T.=1180℃, Aging Less
Heat1, S.T.=1180℃, Aging=816℃
Heat1, S.T.=1180℃, Aging=760℃
Fig. 458 d. Influence of heat and
aging treatment condition on test
results for Co-20Cr-20Ni-Mo-WNi alloy S-816 at 850 °C.
20
10
100
1000
104
Time to rupture (h)
3.3.1.4 Reference
[1] Report on The Mechanical Properties of Metals at Elevated Temperatures: The Iron and Steel
Institute of Japan, (1975).
Landolt-Börnstein
New Series VIII/2B
362
3.3 Co-base alloys
3.3.2 Co-25Cr-10Ni-7.5W-B alloy
3.3.2.1 Introduction
Although the creep strength of Co-base alloys is smaller than that of Ni-base alloys, recently they are
used in turbine applications in a secondary position to Ni-base alloys because of their advantages in hotcorrosion resistance and thermal-shock resistance [1, 2]. Co-25Cr-10Ni-7.5W-B (X-45) is a daughter of
X-40, which was first introduced in the 1940s. These alloys have higher carbon content, because they are
basically strengthened by carbide precipitation. The carbides are mainly M23C6 because of high Cr content
[1]. Tungsten is a solid solution strengthening element [2]. Although the creep strength of X-45 is lower
than that of X-40 due to the lower carbon content, weldability and phase stability is improved. X-45 is
used for nozzle vane partitions in industrial turbines and some aircraft engines. X-45 is normally used in
as cast condition, but sometimes solution heat treatment is applied.
3.3.2.2 Materials standard and chemical composition
The chemical composition requirement of ASTM A567 Grade 13 (discontinued 1987) is shown in Table
209. The mechanical properties of X-45 are reported in [2], [3], and [4].
Table 209. Chemical composition requirement of Co-25Cr-10Ni-7.5W-B (X-45) superalloy castings
ASTM
Chemical composition [wt%]
A567
Si
Mn P
S
Ni
Cr
W
B
Fe
Grade13 C
0.20- 0.75- 0.409.50- 24.5- 7.00- 0.005X-45
≤0.04 ≤0.04
≤2.00
0.30 1.00 1.00
11.50 26.5 8.00 0.015
3.3.2.3 Tensile properties
The 0.2% proof stress and tensile strength of X-45 alloy are shown in Fig. 459 [3]. The yield strength and
tesnsile strength of X-45 alloy are shown in Fig. 460 [2].
Tensile strength
800
700
700
600
600
500
500
Stress (MPa)
Stress (MPa)
0.2% proof stress
800
400
300
400
300
200
200
100
100
0
0
200
400
600
800
1000
Test temperature (℃)
0
0
200
400
600
800
1000
Test temperature (℃)
Fig. 459. 0.2% proof stress and tensile strength of X-45 alloy [3].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 365]
3.3.2 Co-25Cr-10Ni-7.5W-B alloy
363
Ultimate strength, Ftu , and Yield strength, Fty [ksi ]
150
X-45, As cast
Ftu
100
80
60
Fty
40
20
Electroslag remelted
Vacuum-arc remelted
0
400
1200
800
Temperature [F ]
1600
2000
Fig. 460. Yield strength and tensile strength of X-45
alloy [2]. 1 ksi = 6.89476 MPa
3.3.2.4 Creep and rupture properties
Stress vs. creep rupture time relations of X-45 alloy are shown in Fig. 461 [3]. 10, 100, 1,000, 10,000 and
100,000 h rupture-stresses as a function of temperatures are available in Fig. 462 [3] and 463 [2].
Minimum creep rate for X-45 alloy castings are reported as shown in Fig. 464 [3].
g
500
７５０℃
８００℃
８５０℃
９００℃
９５０℃
300
100
80
60
40
Fig. 461. Creep rupture strength
data of X-45 alloy; [3]. n
indicates the total number of data
points.
n=103
20
10
1
10
2
10
3
10
4
Time to rupture (h)
Landolt-Börnstein
New Series VIII/2B
10
5
10
6
364
3.3 Co-base alloys
800
{
{
Stress [MPa]
600
500
400
300
200
100
80
Tensile
strength
0.2%
proof
stress
100 h
60
50
40
30
a
10000 h
20
800
{
{
Stress [MPa]
600
500
400
300
200
100
80
60
50
40
30
20
600
650
700
100000 h
750
850
900
800
Temperature [°C]
Stress [MPa]
Rupture stress [ksi]
1000
200
10 h
100 h
1000 h
30
20
15
100
80 750 °C
60
50 800 °C
40
30
10
8
20
6
4
1000
950
Fig. 462. Temperature dependence of creep rupture strength
from 100 to 100,000 h; [3].
300
100
80
40
0.2%
proof
stress
1000 h
b
60
Tensile
strength
1200
1400
Temperature [F]
1600
1800
Fig. 463. Temperature dependence of creep rupture
strength from 10 to 1,000 h; [2]. 1 ksi = 6.89476 MPa
850 °C
900 °C
950 °C
750 °C
800 °C
850 °C
900 °C
950 °C
n = 48
10
10 -7 10 -6 10 -5 10 -4 10 -3 10 -2 10 -1
Minimum creep rate [% / h]
1
Fig. 464. Stress vs. minimum creep rate for X-45
superalloy castings; [3].
Landolt-Börnstein
New Series VIII/2B
Ref. p. 365]
3.3.2 Co-25Cr-10Ni-7.5W-B alloy
365
3.3.2.5 Referecnes
[1]
[2]
[3]
[4]
The superalloys, ed. by Sims, C. T., and Hagel, W. C., John Wiley & Sons (1972) p.145.
Aerospace Structural Metals Handbook, vol.5, code 4305 (1985).
NRIM Creep Data Sheet, No.30B (1988).
Report on the mechanical properties of metals at elevated temperatures, Vol. IV Superalloys, The
Iron and Steel Institute of Japan (1979).
Landolt-Börnstein
New Series VIII/2B