91钢与奥氏体钢焊接指南-EPRI


    2017 TECHNICAL REPORT
    Electric Power Research Institute
    3420 Hillview Avenue Palo Alto California 943041338 • PO Box 10412 Palo Alto California 943030813 USA
    8003133774 • 6508552121 • askepri@epricom • wwwepricom
    Program on Technology Innovation Guidelines
    and Specifications for High Reliability Fossil
    Power Plants—Best Practice Guideline for
    Manufacturing and Construction of Grade 91
    Steel to Austenitic Stainless Steel Dissimilar
    Metal Welds

    98421449842144EPRI Project Manager
    J Siefert
    J Parker
    ELECTRIC POWER RESEARCH INSTITUTE
    3420 Hillview Avenue Palo Alto California 943041338 ▪ PO Box 10412 Palo Alto California 943030813 ▪ USA
    8003133774 ▪ 6508552121 ▪ askepri@epricom ▪ wwwepricom
    Program on Technology Innovation
    Guidelines and Specifications for
    HighReliability Fossil Power
    Plants—Best Practice Guideline for
    Manufacturing and Construction of
    Grade 91 Steel to Austenitic
    Stainless Steel Dissimilar Metal
    Welds
    3002007221
    Final Report December 2017
    9842144
    DISCLAIMER OF WARRANTIES AND LIMITATION OF LIABILITIES
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    NOTE
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    Electric Power Research Institute EPRI and TOGETHERSHAPING THE FUTURE OF ELECTRICITY
    are registered service marks of the Electric Power Research Institute Inc
    Copyright © 2017 Electric Power Research Institute Inc All rights reserved

    9842144This publication is a corporate document that should be cited in the literature in the following
    manner
    Program on Technology Innovation Guidelines and Specifications for HighReliability Fossil
    Power Plants—Best Practice Guideline for Manufacturing and Construction of Grade 91 Steel
    to Austenitic Stainless Steel Dissimilar Metal Welds EPRI Palo Alto CA 2017 3002007221
    iii
    ACKNOWLEDGMENTS
    The following organizations prepared this report
    Electric Power Research Institute (EPRI)
    1300 West WT Harris Blvd
    Charlotte NC 28262
    Principal Investigators
    J Siefert
    J Parker
    Under contract to EPRI
    Structural Integrity Associates Inc
    11515 Vanstory Dr Suite 125
    Huntersville NC 28078
    Principal Investigators
    T Totemeier
    I Perrin
    This report describes research sponsored by EPRI EPRI acknowledges George Galanes
    Diamond Technical Services for providing feedback comments and input to this report

    98421449842144
    v
    ABSTRACT
    This report provides important guidelines regarding the design fabrication and construction of
    dissimilar metal welds (DMWs) between 9Cr creepstrengthenhanced ferritic steel and
    austenitic stainless steel components designed to transport or collect steam This report does not
    address details specific to the following overlay welding such as for corrosion resistance repair
    welding such as might be addressed by the American Society of Mechanical Engineers (ASME)
    Post Construction Code or the National Board Inspection Code prescriptive guidance including
    how to use particular welding processes or qualification guidelines addressed by ASME Boiler
    and Pressure Vessel Code Section IX
    Several practical considerations exist regarding the specification and fabrication of DMWs and
    it is not possible to address all potential concerns regarding this type of connection in this report
    Instead this report provides basic guidelines that are important to the goal of promoting better
    practices and providing equipment owners or operators information to consider regarding
    additional requirements that are relevant to purchasing specifications Specific factors that affect
    performance in the context of design fabrication operation and metallurgical risk are addressed
    Keywords
    Dissimilar metal welds (DMWs)
    Filler metals
    Metallurgical risk
    Postweld heat treatment
    Weld geometries
    Weld repair methods

    98421449842144

    EXECUTIVE SUMMARY

    vii
    Deliverable Number 3002007221
    Product Type Technical Report
    Product Title Program on Technology Innovation Guidelines and Specifications for
    HighReliability Fossil Power Plants—Best Practice Guideline for Manufacturing and
    Construction of Grade 91 Steel to Austenitic Stainless Steel Dissimilar Metal Welds

    PRIMARY AUDIENCE Research engineers
    SECONDARY AUDIENCE Plant corporate and design engineers
    KEY RESEARCH QUESTION
    There is limited information in the open literature regarding the fabrication and performance of dissimilar metal
    welds (DMWs) between 9Cr creepstrengthenhanced ferritic (CSEF) steel and austenitic stainless steel
    components Furthermore the results from the available studies have not established the factors that
    contribute to the fusion line cracking reported for some inservice failures This report provides examples of
    industry failures in several enduse components and links them to the key contributions from design
    operation fabrication and metallurgy of DMWs
    RESEARCH OVERVIEW
    This report provides important guidelines regarding the design fabrication and construction of DMWs
    between 9Cr CSEF steel and austenitic stainless steel components that are designed to transport or collect
    steam This report does not address details specific to the following overlay welding such as for corrosion
    resistance repair welding such as might be addressed by the American Society of Mechanical Engineers
    (ASME) Post Construction Code or the National Board Inspection Code prescriptive guidance including how
    to use particular welding processes or qualification guidelines addressed by the ASME Boiler and Pressure
    Vessel Code Section IX Numerous practical considerations exist regarding the specification and fabrication
    of DMWs and it is not possible to address all potential concerns regarding this type of connection in this
    report Instead this report provides basic guidelines that are important to the goal of promoting better
    practice and providing equipment owners and operators with information to consider regarding additional
    requirements that are relevant to purchasing specifications Specific factors that affect performance in the
    context of design fabrication operation and metallurgical risk are addressed
    KEY FINDINGS
    The summary of recommendations in this report provides a concise set of guidelines regarding key aspects
    of DMW construction including restrictions on the enduse application weld joint geometry welding process
    postweld heat treatment transition pieces operation and special considerations
    WHY THIS MATTERS
    Every steamraising power boiler in a power generating plant is made with a range of different steels The
    pressure boundary joints necessitate welding of the steels Despite the number of joints that exist there has
    not been an extensive evaluation of fabrication factors that affect performance in crossweld creep for DMWs
    constructed between 9Cr CSEF steel and austenitic stainless steel components This report provides insight
    into the factors that contribute to performance with respect to when a weldment fails and how it fails (that is
    damage tolerance)
    9842144

    EXECUTIVE SUMMARY

    TogetherShaping the Future of Electricity®

    Electric Power Research Institute
    3420 Hillview Avenue Palo Alto California 943041338 • PO Box 10412 Palo Alto California 943030813 USA
    8003133774 • 6508552121 • askepri@epricom • wwwepricom
    © 2017 Electric Power Research Institute (EPRI) Inc All rights reserved Electric Power Research Institute EPRI and
    TOGETHERSHAPING THE FUTURE OF ELECTRICITY are registered service marks of the Electric Power Research Institute Inc
    HOW TO APPLY RESULTS
    This report can be used to improve existing purchase specifications for new DMWs identify potential issues
    regarding installed DMWs or inform the discussion regarding review of bid packages for new units that will
    contain some of these types of welded connections This report should be considered a summary document
    that pulls key information from the following related reports
     Cracking in ThickSection Dissimilar Metal Welds—Case Studies EPRI Palo Alto CA 2015
    3002006759
     Program on Technology Innovation Mechanical Analysis of Dissimilar Metal Welds Part I Insight into
    Potential Failure Modes EPRI Palo Alto CA 2016 3002007215
     Factors Affecting Performance of Dissimilar Metal Welds Creep Performance of Screening Dissimilar
    Metal Welds Between Grade 91 Steel and Stainless Steel 347H EPRI Palo Alto CA 2016
    3002007216
     Factors Affecting Performance of Dissimilar Metal Welds Residual Stress Analysis of Welds Between
    Grade 91 Steel and Stainless Steel 347H EPRI Palo Alto CA 2016 3002007217
     Factors Affecting Performance of Dissimilar Metal Welds Fabrication and Metallurgical Assessment
    of Screening Dissimilar Metal Welds between Grade 91 Steel and Stainless Steel 347H EPRI Palo
    Alto CA 2016 3002007218
     Factors Affecting Performance of Dissimilar Metal Welds Mechanical Analysis of Dissimilar Metal
    Welds Part II Detailed Assessment to Support Best Practice Guidance EPRI Palo Alto CA 2017
    3002007220
     Factors Affecting Performance of Dissimilar Metal Welds Microstructural Characterization and
    Modeling of InService Failures Involving Welds Between Grade 91 Steel and Austenitic Stainless
    Steel EPRI Palo Alto CA 2016 3002007222
    EPRI CONTACTS John Siefert Principal Technical Leader jsiefert@epricom
    PROGRAM Technology Innovation

    9842144
    ix
    TERMS AND DEFINITIONS
    ASME American Society of Mechanical Engineers
    AWS American Welding Society
    B&PV Boiler and Pressure Vessel (Code)
    CSEF creepstrengthenhanced ferritic steel a class of lowalloy steels with minor
    additions of elements (such as vanadium niobium and nitrogen) that are intended
    to enhance elevated temperature strength
    DMW dissimilar metal weld in this report specifically between a 9Cr type CSEF steel
    and an austenitic stainless steel
    HRSG heat recovery steam generator
    MRF metallurgical risk factor which involves an assessment of contributions from steel
    making processing and fabrication to the creep strength and creep ductility
    NBIC National Board Inspection Code
    RCA root cause analysis

    98421449842144
    xi
    CONTENTS
    ABSTRACT V
    EXECUTIVE SUMMARY VII
    1 INTRODUCTION 11
    2 BACKGROUND 21
    Low Alloy Steel DMWs 21
    9Cr DMWs 25
    3 BEST PRACTICE GUIDELINES FOR DESIGN AND FABRICATION 31
    Location 31
    Fabrication 35
    Filler Metals 35
    Transition Pieces 38
    Welding Technique 314
    PostWeld Heat Treatment 315
    PostWeld Machining 316
    Welding Procedure Qualification 316
    4 EXAMPLE APPLICATIONS AND INSERVICE OPERATION EXPERIENCE 41
    TubetoTube Butt Welds 41
    ThickSection Welds 42
    TubetoHeader Connections 44
    Attachments to Piping and Headers 46
    Attachments to Tubing 47
    Thermowells and Steam Sampling Nozzles 413
    Steam Flow Elements 414
    Drains 416
    9842144

    Introduction
    xii
    5 METALLURGICAL RISK 51
    9Cr CSEF Steels 51
    Austenitic Materials 51
    6 INSPECTION 61
    7 LONGTERM PERFORMANCE MICROSTRUCTURAL EVOLUTION AND LIFE
    MANAGEMENT 71
    LongTerm Performance 71
    Microstructural Evolution 75
    Life Management 712
    8 REPAIR 81
    9 CONCLUSION 91
    10 SUMMARY OF RECOMMENDATIONS 101
    11 REFERENCES 111

    9842144
    xiii
    LIST OF FIGURES
    Figure 21 Example of a DMW failure in T22 to 347H showing the macro appearance (A)
    and the distinct fracture surface detail (B) 22
    Figure 31 HRSG configuration highlighting potential DMW locations 32
    Figure 32 Difference in mean coefficient of thermal expansion for Grade 91 and
    common nickelbase filler metals 33
    Figure 33 Headerstubtube arrangement 34
    Figure 34 Effect of composition on the formation of carbide and embrittling phases at
    the fusion line between ferritic steels and commonly specified nickelbase filler
    metals 37
    Figure 35 Example of inadequate control of welding consumables in a girth weld made
    between a butter P91 main steam pipe and an austenitic stainless steel 321H outlet
    header 38
    Figure 36 Example of a poor transition between a 9Cr CSEF steel and an austenitic
    stainless steel 310
    Figure 37 Example of a proper use of a transition piece between a low alloy steel and
    an austenitic stainless steel 311
    Figure 38 Summary of acceptable and unacceptable design practices when installing
    transition pieces in 9Cr CSEF steel to austenitic stainless steel DMWs 311
    Figure 39 Examples of better practice joint designs for DMW fabrication between a
    stainless steel to either a CrMo or 9Cr CSEF steel 313
    Figure 310 Schematic representation of the formation of damage in a step weld where
    damage forms in the upper portion of the weld and is forced to propagate through
    parent material 313
    Figure 311 Feature crossweld creep test example in Grade 91 steel defect formed in
    the HAZ is being driven through more deformationresistant and creep ductile
    parent metal and as evidenced by the large deformation 314
    Figure 312 Failed bend test coupon where in the lower half of the weldment a manual
    GTAW process with 40mmdiameter (0156indiameter) solid wire was used [31] 315
    Figure 313 Measured ultimate tensile strength values at elevated temperature (500°C
    [930°F]) for common nickelbase filler materials (blue) Grade 91 steel plate (black)
    and 347H plate (red) and as compared to minimum values for T91 and T92 317
    Figure 41 Premature failure of a DMW where the DMW was placed at the stubtotube
    weld in a location of poor flexibility 41
    Figure 42 Local throughwall steam leak in 103000 hours in a superheater tube DMW
    between T91 and austenitic stainless steel 347H with 548mm (215in) OD and a
    894mm (0352in) nominal wall thickness 42
    9842144

    Introduction
    xiv
    Figure 43 Premature failure of a DMW in a thicksection 321H outlet header to P91
    piping system 43
    Figure 44 Premature failure of a DMW in a thicksection 304H to P91 piping system 43
    Figure 45 Examples of cracking in a DMWs between a P91 hot reheat header and
    austenitic stainless steel 347H 45
    Figure 46 Premature failure of a DMW in a stainless steel 304H attachment to P91 46
    Figure 47 Examples of thicksection weldments in P91 to Alloy 800A and stainless steel
    TP347H made with filler metals nominally matching in composition to P91 and free
    of welding defects 48
    Figure 48 Bead on plate measurements for nickelbase filler metal ENiCrFe2 deposited
    on a Grade 91 steel plate 48
    Figure 49 Examples of poorly controlled welding procedures premature failure of a
    DMW in a stainless steel sliding spacer to T91 in a final superheater (A) (failure
    mode in this example is attributed to a local seam weld effect) location of
    hardness map through the T91 thickness at the attachment weld (B) and color
    hardness plot showing the extension of the HAZ from the attachment weld through
    the T91 (C) 49
    Figure 410 Example of multiple damage modes in attachment welds in a Grade 91 steel
    hot reheat system 410
    Figure 411 Examples of unacceptable fabrication practices in T91 attachment welds 411
    Figure 412 ASME B&PV Code Section I guidance for welding lugs hangers and
    brackets to shells drums and headers 412
    Figure 413 Summary schematic of important details and notes for welded attachments
    on 9Cr CSEF steel tubing 412
    Figure 414 Example of a DMW failure in a Grade 91 boss to stainless steel 316
    thermowell 414
    Figure 415 Failure of a steam sampling nozzle in a Grade 91 hot reheat HRSG system 414
    Figure 416 Catastrophic failure in a DMW between a Grade 91 main steam piping
    system to a stainless steel flow nozzle 415
    Figure 417 Macro view (A) and a crosssection (B) of a stainless steel 316 flow element
    welded in a P91 main steam HRSG piping system 415
    Figure 61 Example of damage through the wall thickness of a low alloy steel to stainless
    steel DMW (nickelbase filler metal) In Figure 61 note that depending on the type
    and quality of performed inspection the results may indicate that this weld is at end
    of life a stated life fraction or contain no damage aside from the oxide notch at the
    surface 63
    Figure 71 Comparison of reported fusion line failures in Grade 91 DMWs to Grade 91
    HAZ failure database and meanminimum relationships for the database 71
    Figure 72 Fabrication of screening DMWs for evaluation in creepdominated test
    conditions 72
    Figure 73 Identified regions in a DMW adapted from the classic description by Nippes
    and as modified and reported in 75
    Figure 74 Asfabricated condition of the fusion line between Grade 91 steel and a
    nickelbase filler metal 77
    9842144

    Introduction
    xv
    Figure 75 Microstructural regions and the phase(s) present composition details and
    association with inservice damage 78
    Figure 76 Location of damage in the ferrite band adjacent to the fusion line 78
    Figure 77 Location of damage in the ferrite band adjacent to the fusion line and
    composition mapping for nickel as determined using SEMEDS 79
    Figure 78 Location of damage in the ferrite band adjacent to the fusion line and
    composition mapping for chromium as determined using SEMEDS 79
    Figure 79 Schematic of potential explanations for the evolution of ferrite in the unmixed
    zone 711
    Figure 710 Results of computational analysis which show the sensitivity of the nickel
    base weld to Grade 91 fusion line region to carbon migration 712
    Figure 81 Simulated weld repair made using EPRI P87 filler metal and in a P91 to
    stainless steel 316 DMW (the original weld metal is ENiCrFe3) 81
    Figure 91 Example of recent failures in DMWs in CSEF steel T91 to austenitic stainless
    steel 347H in a wingwall panel of a natural circulation fluidized bed boiler 92

    98421449842144
    xvii
    LIST OF TABLES
    Table 21 Predicted Life as a Function of Stress and Temperature for an Austenitic
    Stainless Steel 309type Filler Metal and Nickelbase Filler Metal (ENiCrFe3) 23
    Table 22 Summary of Screening Criterion for Candidate Dissimilar Metal Weld Filler
    Metals 24
    Table 23 Comparison of Factors Contributing to Dissimilar Metal Weld Failures in
    Ferritic to Austenitic Stainless Steel Dissimilar Metal Welds 25
    Table 24 Composition ranges for commonly specified base metals used in the
    fabrication of DMW between 9Cr CSEF steels and austenitic stainless steels 29
    Table 31 Pro and cons for placing DMWs in small bore tubing or large bore piping
    components 31
    Table 32 Challenges regarding the location of DMWs in an HRSG system 32
    Table 33 Composition ranges for commonly specified nickelbase filler metals used in
    the fabrication of DMWs between 9Cr CSEF steels and austenitic stainless steels 36
    Table 34 Descriptions for the DMW simulations in Figure 34 37
    Table 35 Minimum ultimate tensile strength values for materials commonly utilized in
    DMWs 317
    Table 61 Validity and comments for common inspection techniques that may be applied
    to DMWs 62
    Table 71 Results of uniaxial feature crossweld creep testing for DMWs manufactured
    in P91 steel to the approaches provided in Figure 72 73
    Table 72 Comparison of results at 625°C (1157°F) and 80 MPa (116 ksi) for time to
    failure (such as when the sample fails) and the life fraction spent in tertiary creep
    (such as how the sample fails) 74
    Table 73 Descriptions of the regions in Figure 73 and specific to a DMW in a
    9Cr CSEF steel where the weld metal is nickelbase 76

    98421449842144
    11
    1
    INTRODUCTION
    Currently 9Cr creepstrengthenhanced ferritic (CSEF) steels are being specified in a wide
    variety of power plant applications including a number of unique geometries and in conditions
    that can be classified as lowtemperature (<550°C [1022°F]) or hightemperature (≥550°C
    [1022°F]) with respect to this family of steels Increased demand for efficient power plants has
    led to these materials being used in designs approaching or exceeding 600°C (1112°F) and
    resulted in a need to transition from these materials to suitable austenitic stainless steel grades
    (note that the currently allowable stress values in American Society of Mechanical Engineers
    [ASME] Boiler and Pressure Vessel [B&PV] Code Section IID allow 9Cr steels to be utilized
    to 650°C [1200°F]) The transition from a 9Cr steel to stainless steel generally requires that a
    dissimilar metal weld (DMW) be made although in some cases it is possible to use a
    mechanical connection such as a bolted flange
    The performance (that is time to and mode of cracking and failure) of DMWs operating at high
    temperature is highly variable and sensitive to details in design fabrication operation and
    metallurgical risk factors associated with the base metals being joined It should be emphasized
    that these factors introduce significantly greater complexity than is the case in fusion welds
    between similar’ steels Thus there is no rule of thumb approach to overcome the complexities
    involved with the transitions introduced Similarly it is not the case that always performing a
    post weld heat treatment (PWHT) during or after fabrication will mitigate problems Indeed
    sometimes the act of performing PWHT can introduce or at least exaggerate the variables which
    must be considered
    The very large number of potential DMWs is because typically there are only outline guidelines
    available in most Construction Codes For example there are a number of different types of
    DMWs commonly present in power plant systems allowed by conventional construction rules
    within ASME B&PV Code Section I or ASME B311 However there are no specific rules
    or recommendations that address all DMW joint combinations of the steels allowed by design
    [1 2] Because of the very large number of possible configurations that are currently allowed for
    DMW construction it is important to emphasize the need to address DMW issues on a caseby
    case basis and engineer these connections to the unique requirements of each specific
    application
    Historically the observed inservice performance regarding the time to failure for DMWs has
    been highly variable This is problematic since postconstruction life management activities
    must base inspection plans on a reasonable expectation of inservice performance (that is when a
    DMW will fail) The second consideration in performance is damage tolerance (that is how the
    DMW will fail) which is equally important as the time to failure Optimization of the eventual
    failure is vital not only from a primary safety perspective but also to minimize the potential for
    collateral damage and reduce the complexity of any repairs that must be made The variability in
    DMW performance can be significantly reduced by application of a set of consistent design and
    fabrication practices Focusing engineering efforts on a smaller number of joint type would have
    9842144

    Introduction
    12
    significant benefits across industry as it would be possible to use the results of inservice
    experience as a guide to defining the factors affecting life This document provides information
    and recommendations on these practices that are aimed at not only increasing lifetime but also
    addressing how the DMW will fail in service Adopting these recommendations offers major
    benefit to all stakeholders in the electricity supply industry

    9842144
    21
    2
    BACKGROUND
    Low Alloy Steel DMWs
    The need for DMWs between ferritic steels and austenitic stainless steels was documented as
    early as 1935 [3] and sufficient information was generated such that a series of initial reviews on
    the topic of DMWs was authored starting in the early 1940s [4–6] Initial studies of DMWs were
    concerned with the effect of dilution and other fundamental firstprinciple issues regarding
    material compatibility With regard to the power generation industry—and more specifically in
    coalfired power boilers in the United States—DMWs between ferritic steels and austenitic
    stainless steels have been widely utilized since the 1950s The need for DMWs arose from the
    fact that the new generation of coalfired units being put into service were operating with steam
    temperatures in the range of 565–593°C (1050–1100°F) [7] requiring the use of austenitic
    stainless steels and therefore a number of potential DMWs In this period investigations were
    conducted on DMWs between low alloy (1 to 9Cr) ferritic steels and 300 series (nominally
    18Cr10Ni) austenitic stainless steels A number of welding consumables were investigated
    including austenitic stainless steel filler metals low alloy steel filler metals and nickelbase
    alloy filler metals [8 9] It is emphasized that in the 1950s and 1960s there was a large volume
    of research investigating DMWs in componentspecific scenarios such as for tubes attachments
    and piping [10 11] and it was widely recognized that the issues concerning DMWs were
    potentially unique to the endcomponent application
    The widespread and unpredictable nature of failures in DMWs through the 1970s (see Figure 2
    1) led to the formation of a Dissimilar Metal Weld Task Group that was jointly led by ASME
    ASTM International and the Materials Property Council (MPC) As a consequence of the
    outcomes prescribed by this task group the Electric Power Research Institute (EPRI) initiated a
    collaborative research program to address the challenges facing the industry This research
    project RP1874 began in 1980 with six key tasks and ended when the last report was published
    in 1989 The coordinated effort involved at least 27 individual corporations and no fewer than 65
    key industry contacts including the direct involvement of the stakeholders in the manufacturing
    supply chain and on a global basis such as endusers original equipment manufacturers
    (OEMs) material suppliers research labs and universities [12] EPRI presented the results from
    RP1874 in a series of reports listed below
     DissimilarWeld Failure Analysis and Development Program Volume 1 Executive
    Summary EPRI Palo Alto CA 1985 CS4252V1
     DissimilarWeld Failure Analysis and Development Program Volume 2 Metallurgical
    Characteristics EPRI Palo Alto CA 1985 CS4252V2
     DissimilarWeld Failure Analysis and Development Program Volume 3 Accelerated
    Discriminatory Tests EPRI Palo Alto CA 1986 CS4252V3
    9842144

    Background
    22
     DissimilarWeld Failure Analysis and Development Program Volume 4 Utility Plant
    Results EPRI Palo Alto CA 1985 CS4252V4
     DissimilarWeld Failure Analysis and Development Program Volume 5 Evaluation of
    Acoustic Emission and Enhanced Radiography EPRI Palo Alto CA 1985 CS4252V5
     DissimilarWeld Failure Analysis and Development Program Volume 6 Weld Condition
    and Remaining Life Assessment Manual EPRI Palo Alto CA 1988 CS4252V6
     DissimilarWeld Failure Analysis and Development Program Volume 8 Design and
    Procedure Guide for Improved Welds EPRI Palo Alto CA 1989 CS4252V8
    Figure 21
    Example of a DMW failure in T22 to 347H showing the macro appearance (A) and the
    distinct fracture surface detail (B)
    Key outcomes from this research still relevant to DMWs manufactured today included the
    following design recommendations
     DMWs should be placed away from highrestraint locations (for example fixed locations
    such as the roof header or attachments)
     Minimize thermal cycling which includes both inservice cycling within the boiler and the
    unit operating mode (startups and shutdowns)
     Thermocouples should be utilized to monitor both operating temperature and cycling
     Tube wall thickness transitions should be made with a transition piece or an upset stainless
    steel tube
    An additional set of guidelines were provided with respect to fabrication
     Preference (and eventual widespread adoption) for nickelbase filler materials over stainless
    steel filler metals and recognition that not all nickelbase filler metals are the same This led
    to informal standardization within major OEMs to use ERNiCr3 for the gas tungsten arc
    welding (GTAW) process and either ENiCrFe2 or ENiCrFe3 for the shielded metal arc
    welding (SMAW) process
     Optimization of welding practice (proper consideration of backing rings avoiding defects
    and so on)
     Avoid excessive weld reinforcement
    9842144

    Background
    23
     Benefit of postweld heat treatment (PWHT) was not clear (and for CrMo steels PWHT is
    undesirable)
     Consideration of a wider bevel or wide cap in joint design
    The widespread adoption of nickelbase filler materials for joining DMWs was predicated on
    both the observed differences regarding inservice performance and the life predicted from
    laboratory testing as shown in Table 21 A review of available filler materials at that time
    (summarized in Table 22) suggested that in general nickelbase filler metals would be the best
    option and there was merit in optimizing a new filler material designed for the purpose of joining
    low alloy steels to austenitic stainless steels This filler material was identified as HFS6 but was
    never commercialized due to gross microfissuring issues A summary of concerns regarding
    other filler metal options is provided below
     The poorest performance was observed in austenitic stainless steel filler metals such as E309
    or similar
     Nickelbase filler metals with Al and Ti additions such as Inconel 61 Inconel 92 and
    Hastelloy1 W will ageharden in service reducing creep ductility and increasing the strength
    mismatch across the fusion line
     Martensitictype filler metals such as X20 E410 9Cr1Mo (for example –B8 type filler
    metals) would require PWHT This was generally not desirable especially since for
    common wall thicknesses in DMWs T22 was not required to be given a PWHT according to
    ASME B&PV Code Section I (Table PW394)
    Table 21
    Predicted Life as a Function of Stress and Temperature for an Austenitic Stainless Steel
    309type Filler Metal and Nickelbase Filler Metal (ENiCrFe3)
    Note that the Nickelbase filler metal consistently achieves a life that is 23X greater

    1 INCONEL and INCOWELD are registered trademarks of the Special Metals Corporation Hastelloy is a
    registered trademark of Haynes International
    Temperature Stress Stainless Steel Filler Metal ENiCrFe3 Filler Metal
    °C °F MPa ksi Life (hours) Life (hours)
    625 1157 620 90 350 775
    585 1085 620 90 2750 6100
    550 1022 620 90 20000 43000
    625 1157 345 50 2880 6370
    585 1085 345 50 22600 50000
    550 1022 345 50 161800 358000
    625 1157 173 25 35400 78400
    585 1085 173 25 278000 615000
    550 1022 173 25 > 1000000 > 1000000
    Scenario Tube with OD 508 mm (20 inch) WT 95 mm (0375 inch) OD
    Operating Pressure 172 barg (2500 psig) Axial stress 221 MPa (32 ksi)
    9842144

    Background
    24
    Table 22
    Summary of Screening Criterion for Candidate Dissimilar Metal Weld Filler Metals [66]
    Filler Metal
    Difference in Coefficient
    of Thermal Expansion
    (CTE) with 2¼Cr1Mo
    (RT1000F°)
    Tendency to
    Form Type I
    Carbide PWHT
    DMW Performance
    Observations
    E309 27 higher None No Gives poorest performance
    INCONEL 92 5 higher Slight No Shows little tendency to
    interfacial failure
    INCONEL 132 7 higher Marked No Significantly better than E309
    but shows interfacial failure
    INCONEL 182 10 higher Marked No Better than E309 in most
    cases and widely used Some
    tendency to ageharden
    Filler Metal 82 3 higher Marked No Better than E309 and widely
    used with GTAW
    INCOWELD A 3 higher Marked No Fairly stable
    INCONEL 61 3 higher Slight No Marked agehardening
    Hastelloy W 9 lower Slight No Very marked agehardening
    X20CMoV121 18 lower None Yes Thermally softens
    E410 19 lower None Yes Thermally softens
    9Cr1Mo 14 lower None Yes Thermally softens
    HFS6 7 higher None No No tendency to interfacial
    failure shows best life
    It should be recognized that DMWs in a simple low alloy steel such as Grade 22 can exhibit
    markedly different behavior than for DMWs in a CSEF steel such as Grade 91 steel Important
    differences are highlighted in Table 23 for Grade 22 and Grade 91 welded to an austenitic
    stainless steel It is clear from this comparison that although general guidelines may be applied
    across the different alloys in DMWs the details are unique to the materials requiring welding
    and special considerations may need to be taken into account In the balance of the document
    the challenges regarding the fabrication of 9Cr CSEF steels to austenitic stainless steels will be
    reviewed

    9842144

    Background
    25
    Table 23
    Comparison of Factors Contributing to Dissimilar Metal Weld Failures in Ferritic to
    Austenitic Stainless Steel Dissimilar Metal Welds
    Factor Grade 22 (Low Alloy CrMo Steel) Grade 91 (CSEF Steel)
    Creep range (parent
    metal) ≥ ~480°C (900°F) ≥ ~540°C (1000°F)
    Typical DMW use
    temperature
    ≥ ~540°C (1000°F) ≥ ~575°C (1065°F) (tubing)
    ≥ ~540°C (1000°F) (piping)
    Relative failure time
    for DMW compared
    to a similar weld
    (creepdominated)
     Creepdominated failures in
    laboratory tests show a consistent
    reduction from documented
    behavior in similar welds
     Creepdominated failures in
    laboratory tests fall into database
    for HAZ failures in similar welds
     Service failures show variable
    behavior that in some cases
    cannot be explained by laboratory
    test data
    Attributed
    metallurgical risk
    factor for fusion line
    failure
     Coarsegrain HAZ (for SS filler
    metal)
     Type I carbides (for nickelbase
    filler metal)
     Formation of ferrite at the fusion
    line
     Possibly others research ongoing
    Common
    components
     Tubes attachments and
    thermowells
     Few examples in piping (See
    CEGB experience and early US
    experience in the 1950s and
    1960s)
     Tubes attachments thermowells
     Piping in stateoftheart heat
    recovery steam generators
    (HRSGs) where the wall thickness
    is >050 in (127 mm)
    Observed failure
    modes
     Fusion line failure ~1–2 grains
    from the fusion line in the Grade
    22 coarsegrained HAZ for DMWs
    with stainless steel filler materials
     Fusion line failure adjacent to the
    fusion line in the Grade 22 for
    DMWs with nickelbase filler
    materials Damage accumulates
    at Type I carbides
     Oxide notching in thinsection
    components andor where a
    source of thermal cycling is
    present
     Fusion line failure with damage
    concentrated in the ferrite band
     Oxide notching
     Cavitation in the Grade 91 HAZ
     Mixed mode (HAZ + fusion line)
     Failure on the stainless side due to
    stress relaxation cracking or
    thermal fatigue
    9Cr DMWs
    Grade 91 steel was first adopted by the ASME B&PV Code in 1983 The general use of Grade
    91 steel in initial applications in the United States and Europe was as a replacement material for
    either low alloy steels (such as replacement headers due to ligament cracking in Grade 22) or
    damaged or temperamental steels such as CrMoV (UK) or X20 (Germany) Widespread use in
    new construction of supercritical ultrasupercritical and HRSGs throughout the world started in
    9842144

    Background
    26
    the late 1990s and early 2000s Through the 1980s and 1990s there was little research conducted
    specific to DMWs in either Grade 91 or the newer Grade 92 welded to austenitic stainless steels
    (note that Grade 92 was introduced to ASME B&PV Code in 1995 as Code Case 2179) In light
    of this the stateoftheart manufacturing of DMWs in 9Cr steels to austenitic stainless steels
    remains heavily influenced by previous learnings from low alloy steels with little consideration
    given to the unique aspects of these newer more complex materials
    In the early 2000s the increase in new construction prompted EPRI to initiate research to revisit
    the HFS6 composition Eventually the microfissuring issue that plagued the original HFS6
    composition was eliminated and a patent was filed in 2004 The new modified composition
    known as EPRI P87 filler metal possessed an optimized NbC ratio where Nb was added to
    ~13 wt and C was added to ~010 wt The patent for this filler material was later
    published in 2009 two detailed EPRI reports document its development and assessment [14 15]
    In the mid2000s a number of catastrophic failures in DMWs in Grade 91 steel components
    occurred including flow nozzles (predominantly made from 316H) installed in main steam
    piping HRSGs thermowells and steam sampling probes The industry response was relatively
    limited and in many cases rootcause or failure analysis was not performed In stark contrast to
    the legacy of DMWs in low alloy steels no coordinated response was organized by EPRI or
    other entities to investigate and evaluate the specific challenges in DMW fabrication and
    operation between 9Cr CSEF steels and austenitic stainless steels More recently and starting
    in ~2015 the first failures in tubetotube DMWs between 9Cr steels and austenitic stainless
    steels were reported to EPRI including failures after ~50000 to ~125000 hours of service The
    range in types of failures and specific features will be reviewed in subsequent sections of this
    report
    In response to a review of failures [16] and particularly in thicksection DMWs EPRI launched
    an integrated program to evaluate and assess the installation of DMWs in cyclic and thick
    section components This project initiated in late 2014 targeted the application of DMWs in
    stateoftheart HRSGs where the drive for increased efficiency has pushed steam parameters
    (for the first time in HRSGs) into a regime that would require a material transition from 9Cr
    CSEF steel to an austenitic stainless steel This research has culminated in a number of
    technology transfer initiatives including proposed nonmandatory appendices for incorporation
    into ASME B&PV Code Section I and ASME Power Piping Code B311 workshops hosted by
    EPRI and ASME Pressure Vessels and Piping Conference through 2017 and seven reports
     Program on Technology Innovation Mechanical Analysis of Dissimilar Metal Welds Part I
    Insight into Potential Failure Modes EPRI Palo Alto CA 2016 3002007215
     Factors Affecting Performance of Dissimilar Metal Welds Creep Performance of Screening
    Dissimilar Metal Welds Between Grade 91 Steel and Stainless Steel 347H EPRI Palo Alto
    CA 2016 3002007216
     Factors Affecting Performance of Dissimilar Metal Welds Residual Stress Analysis of Welds
    Between Grade 91 Steel and Stainless Steel 347H EPRI Palo Alto CA 2016 3002007217
     Factors Affecting Performance of Dissimilar Metal Welds Fabrication and Metallurgical
    Assessment of Screening Dissimilar Metal Welds Between Grade 91 Steel and Stainless Steel
    347H EPRI Palo Alto CA 2017 3002007218
    9842144

    Background
    27
     Factors Affecting Performance of Dissimilar Metal Welds Mechanical Analysis of
    Dissimilar Metal Welds Part II Detailed Assessment to Support Best Practice Guidance
    EPRI Palo Alto CA 2017 3002007220
     Guidelines and Specifications for HighReliability Fossil Power Plants Best Practice
    Guideline for Manufacturing and Construction of Grade 91 Steel to Austenitic Stainless Steel
    Dissimilar Metal Welds EPRI Palo Alto CA 2017 3002007221
     Factors Affecting Performance of Dissimilar Metal Welds Microstructural Characterization
    and Modeling of Inservice Failures Involving Welds between Grade 91 Steel and Austenitic
    Stainless Steel EPRI Palo Alto CA 2016 3002007222
    As noted above failures during service have been reported in DMWs between 9Cr CSEF
    steels and austenitic stainless steels joined using a nickelbase filler metal and the standard use
    of nickelbase filler metals for joining of DMWs between ferritic and austenitic stainless steels is
    an outcome of the best practice provided through a series of legacy EPRI reports specific to
    DMWs between low alloy steels and austenitic stainless steels [12 17] Commonly utilized
    parent materials for DMWs are provided in Table 24 The scope and variety of recent failures in
    9Cr type CSEF steel to austenitic stainless steel DMWs include the following [16 18 19]
     Catastrophic failures in stainless steel flow element nozzles and stainless steel thermowells
    welded into Grade 91 steel piping systems in HRSG systems
     Repeat failures in the form of leaks in thicksection Grade 91 steel to stainless steel girth
    welds at terminal points and material thickness transitions in main steam piping systems
     Failures in stainless steel attachment and slip spacer welds to 9Cr tubing (failures
    numbering in the tens of thousands)
     Failures in tubetotube butt welds in superheater pendants and HRSG harps
     Failures of stainless steel warming lines connected to Grade 91 valve bodies
    As noted above DMWs have failed in service both catastrophically (for example rupture) or
    with leakbeforebreak behavior (such as a local steam leak) DMW failures have occurred in a
    range of component applications in lowtemperature service (<550°C [1022°F]) or high
    temperature operation well within the creep regime for 9Cr steels (≥550°C [1022°F])
    Documented failure modes include but may not be limited to (mechanism or metallurgical root
    cause of failure underlined)
     T91T92 to stainless steel sliding spacers seam weld effect
     T91T92 to stainless steel sliding spacers poor quality
     T91 to stainless steel reheater attachments HAZ damage oxide notching and fusion line
    damage
     T91 to stainless steel tube to tube weld (superheater application) oxide notching in T91
     P91 to stainless steel thick section girth weld fusion line failure
     P91 to stainless steel thick section girth weld HAZ failure
    9842144

    Background
    28
     P91 to stainless steel thick section girth weld mixed mode failure (that is HAZ and fusion
    line)
     P92 – IN617 transition piece – HR3C tube stress relaxation cracking in HR3C
    Although traditional failure analysis is typically sufficient to classify a failure as being DMW
    related it is emphasized that EPRI recommends that a meaningful root cause analysis be
    performed to better define the specific failure mechanism and the design metallurgical or
    operational root cause of the observed failure In many cases there may be several contributing
    factors and thus it is vitally important that expert RCA is carried out for DMWs (the list above
    recognizes no less than eight types of failure) Moreover the characteristics of the component
    and its operation such as service temperature section thickness local stiffness or flexibility
    details of the fabrication and so on significantly contribute to a given documented failure in a
    DMW Damage mechanisms responsible for these failures include creep fatigue corrosion or
    an interaction of these
    Historically the power generation industry has relied on internal specifications to capture the
    lessons learned from plant failures restrict poor practice and reduce potential issues with
    DMWs The background of many of these specifications came from analysis of research results
    such as the series of studies completed by the Dissimilar Metal Weld Task Group described
    above Although substantial information was generated from investigating inservice failures and
    from these and other engineering studies no formal rules regarding DMW design or use were
    introduced into ASME B&PV Section I or B311 Codes
    Today the only reference that exists in Section I with respect to the fabrication of DMWs are the
    PWHT considerations described in Note 2 of Table PW 395 To meet the intent expressed in the
    foreword to Section I—namely to afford reasonably certain protection of life and property and
    to provide a margin for deterioration in service to give a reasonably long safe period of
    usefulness—it is the principal objective of the present document to record and preserve the
    lessons learned to date This knowledge will thus be available to future generations of engineers
    and key learnings can be incorporated into specifications regarding the design fabrication
    construction and operation of DMWs

    9842144

    Background
    29
    Table 24
    Composition ranges for commonly specified base metals used in the fabrication of DMW between 9Cr CSEF steels and
    austenitic stainless steels [20–23]
    Element T91
    (Type I)
    T91 Type II
    (CC 2864) T92 Grade 93
    (CC 2839) 304H 347H 347HFG Super 304H
    (CC 2328)
    TP310HCbN
    (HR3C)
    C 007–014 008–012 007–013 005–010 004–010 004–010 006–010 007–013 004–010
    Mn 030–060 030–050 030–060 020–070 200 200 200 100 200
    P 0020 0020 0020 0020 0045 0045 0045 0040 0045
    S 0010 0005 0010 0008 0030 0030 0030 0010 0030
    Si 020–050 020–040 050 005–050 100 100 100 030 100
    Ni 040 020 040 020 80–110 90–130 90–130 75–105 190–220
    Cr 80–95 80–95 85–95 850–950 180–200 170–190 170–190 170–190 240–260
    Mo 085–105 085–105 030–060
    V 018–025 018–025 015–025 015–030
    B 0001 0001–
    0006
    0007–
    0015 0001–0010
    Nb 006–010 006–010 004–009 005–012 10xC to 110 10xC to 110 030–060 020–060
    N 0030–0070 0035–0070 0030–
    0070
    0005–
    0015 005–012 015–035
    Al 0020 0020 002 0030 0003–0030
    W 005 15–200 25–35
    Ti 001 001 001
    Zr 001 001 001
    Cu 010 25–35
    As 0010
    Sb 0003

    9842144

    Background
    210
    Table 24 (Continued)
    Composition ranges for commonly specified base metals used in the fabrication of DMW between 9Cr CSEF steels and
    austenitic stainless steels [20–23]
    Element T91
    (Type I)
    T91 Type II
    (CC 2864) T92 Grade 93
    (CC 2839) 304H 347H 347HFG Super 304H
    (CC 2328)
    TP310HCbN
    (HR3C)
    Sn 0010
    Co 25–35
    Nd 0010–006
    O 00050
    NAl
    Ratio ≥40


    9842144
    31
    3
    BEST PRACTICE GUIDELINES FOR DESIGN AND
    FABRICATION
    Location
    The initial aspects of design that must be considered include the location of DMWs within a
    system whether the DMWs are in thin or thicksection components and basic requirements
    regarding fabrication (for example ensuring thickness transitions are properly accommodated)
    The location of the DMW in the system must address potential future concerns about safety and
    consequence of failure (that is proximity to highly trafficked areas of the plant and risk of
    collateral damage) Particularly in larger bore piping systems DMWs should be located in a
    location amenable to nondestructive examination and should be provided with sufficient
    surrounding space to facilitate such inspection
    An important set of decisions may need to be made in the design phase regarding placement of
    DMWs and whether DMWs should be fabricated in thinsection tubing or thicksection piping
    material Some considerations for this decision are highlighted in Table 31 A review of
    challenges particularly regarding the placement of DMWs in an HRSG system (Figure 31) are
    detailed in Table 32
    Table 31
    Pro and cons for placing DMWs in small bore tubing or large bore piping components
    Tubes (Small bore) Pipes (Large Bore)
    Pros

     Few failures in tubing where DMW is
    located away from header
     DMWs receive PWHT
     Low failure consequence (safety)
     DMW can be placed in a location for easy
    access
     Can be shopfabricated to better control and
    manage the welding and associated
    fabrication processes
    Cons

     Many (hundreds) of welds
     Often poor access for inspection
     Potentially difficult to repair
     Lack of criteria for life management
     Potentially high consequence of failure
    (separation of pipe)
     Potentially frequent repairinspection and
    replacement
     Lack of criteria for life management


    9842144

    Best Practice Guidelines for Design and Fabrication
    32

    Figure 31
    HRSG configuration highlighting potential DMW locations (1 tubing internal to the HRSG
    setting 2 link piping 3 outlet piping)
    Table 32
    Challenges regarding the location of DMWs in an HRSG system
    Location Challenges
    1 Tube to
    Tube
     Many to inspect and potentially uninspectable
     Lower consequence of failure
     Can locate away from header
     Requires CSEF stub tubes
    2 Link Pipe  Moderate number and can be located in inspectable region
     Moderate failure consequence (safety)
     Can use piping stress analysis to identify low stress location
     Requires stainless steel headers
    3 Outlet Piping  One or two welds can be located in accessible area with low stress
     Major consequence if failure does occur (large steam release pipe whip)
     Maybe at scope interface and not properly engineered
    Note A Stainless
    Steel
    Headers
     Thermal fatigue risk of header due to unfavorable thermal properties
     Welding challenges risk of relaxation cracking
     Sensitization of stainless steel
    9842144

    Best Practice Guidelines for Design and Fabrication
    33
    DMWs should be located taking into account a desire to minimize the secondary system loads as
    well as the risks of personnel injury and collateral damage in case of failure The difference in
    primary and secondary loads are described below in reference to DMW performance [12]
     Primary system loads or stresses – continuous loads applied during typical service that are
    not relaxed by deformation This type of loading includes deadweight and steampressure
    loads Although primary system loads are generally not excessive in typical DMW
    applications (most DMWs are girth welds so that crossweld pressure loading is low) local
    failure of an attachment may increase deadweight loading
     Secondary system loads or stresses – loads generated during typical service from thermal
    expansion mismatch or thermal displacements These loads are straincontrolled and can be
    relaxed by inelastic deformation at elevated temperature but they may be regenerated by
    temperature cycling Some longrange thermal stresses cannot be fully relaxed by small local
    inelastic strains and must be classified as primary stresses A subset of secondary system
    loads referred to as selfstresses are sometimes used to describe the stresses that result from
    differential thermal expansion between the parent metal and filler metal constituents in a
    DMW The magnitude of the selfstress is determined by the difference in thermal expansion
    coefficient between the ferritic parent metal and selected filler metal Coefficients for
    Grade 91 steel and typical Codeapproved F No 43 nickelbase filler metals are shown in
    Figure 32

    Figure 32
    Difference in mean coefficient of thermal expansion for Grade 91 and common nickelbase
    filler metals (Note ENiCrCoMo1 equivalent to Filler Metal 617 ENiCrFe2 commonly
    referred to as INCOWELD A ENiCrMo3 equivalent to Filler Metal 625 and EPRI P87
    incorporated into ASME as Code Cases 2733 and 2734)

    9842144

    Best Practice Guidelines for Design and Fabrication
    34
    DMWs operating in the creep range have experienced inservice failures in the past One key
    factor identified in these failures is the influence of secondary system loads There are a number
    of examples where these loads have contributed to failure some of them are presented in the
    following sections The following guidelines should be followed in locating DMWs to minimize
    system loads
     The DMW should not be located in a highrestraint location A highrestraint location is
    generally defined as one in which the position is fixed such as at a roof penetration the
    location of a mechanical restraint or attachment or in close proximity to a large component
    such as a header Attachments may exist in a welded or mechanical form including welded
    attachments and sliding spacers in tubes or lugshangers in a piping system Some specific
    guidelines are
    – DMWs should be located away from a fixed location such as roof penetration [12]
    – DMWs should never be placed at the stub to header weld (Figure 33) DMWs may be
    placed at the stub to tube weld provided it has been demonstrated that there is sufficient
    flexibility in the system [19]
    – DMWs should not be placed at an attachment weld in thicksection piping and the DMW
    should be as far as possible from any type of rigid restraint in a tube bundle or in a piping
    system [12]
     DMWs should be located where the component has sufficient flexibility to avoid
    intensification of secondary stresses Routine failures occur for example where DMWs are
    placed at terminal locations (such as a pipetoturbine stop valve or an outlet header–topipe
    weld)
     DMWs should be placed away from local sources of thermal cycling for example
    sootblowers or attemperators
    Figure 33
    Headerstubtube arrangement (Note DMWs should not be placed at the stubtoheader
    weld and the stub should be sufficiently long to reduce restraint if the DMW is to be
    placed at the stubtotube location)
    9842144

    Best Practice Guidelines for Design and Fabrication
    35
    Fabrication
    Filler Metals
    Filler metals used to weld 9Cr CSEF steels to austenitic stainless steels should comply with
    applicable approved codes Filler materials commonly specified in the construction of DMWs in
    9Cr CSEF steels are provided in Table 33 Preferred consumables have the following AWS
    classifications ERNiCr3 ENiCrFe3 ENiCrFe2 ENiFeCr4 (Code Case 2734) and
    ERNiFeCr4 (Code Case 2733) Results from a recent report detailing links between specific
    consumables and the formation of carbides and other potentially embrittling phases are shown in
    Figure 34 [24] the evaluation of the simulated conditions is described in Table 34 The study
    showed that filler metals consistent with AWS classifications ERNiCrMo3 ENiCrMo3
    ERNiCrCoMo1 and ENiCrCoMo1 may form embrittling phases at the weld fusion line as a
    consequence of PWHT andor longterm operation The development of brittle constituents
    significantly increases the risk of rapid crack growth and these filler materials should not be
    used in 9Cr to austenitic stainless steel DMWs
    Although widespread failures have not been documented at the time of writing where these
    undesirable filler materials have been utilized it should be noted that most supercritical and
    ultrasupercritical units in the world have not yet achieved 100000 hours of operation It remains
    to be seen whether the evolution of deleterious phases at the fusion line will contribute to
    accelerated failure in the same manner as observed for low alloy steel DMWs However since a
    potential issue exists and there is no particular offsetting benefit there is no basis to specify these
    filler metals (ERNiCrMo3 ENiCrMo3 ERNiCrCoMo1 and ENiCrCoMo1) for newly
    fabricated DMWs

    9842144

    Best Practice Guidelines for Design and Fabrication
    36
    Table 33
    Composition ranges for commonly specified nickelbase filler metals used in the fabrication of DMWs between 9Cr CSEF
    steels and austenitic stainless steels [25–27]
    Element ENiCrFe3
    (IN 182)
    ERNiCr3
    (FM 82)
    ENiCrFe2
    (INCO A)
    ENiCrMo3
    (IN 112)
    ERNiCrMo3
    (FM 625)
    ENiCrCoMo1
    (IN 117)
    ERNiCrCoMo1
    (FM 617)
    ENiFeCr4
    ERNiFeCr4
    (EPRI P87)
    C 010 max 010 max 010 max 010 max 010 max 005–015 005–015 008–014
    Si 10 max 050 max 075 max 075 max 050 max 075 max 10 max 005–050
    Mn 50–95 25–35 10–35 10 max 050 max 03–25 10 max 12–18
    Cr 130–170 180–220 130–170 200–230 200–230 210–260 200–240 85–95
    Ni 59 min 67 min 62 min 55 min 58 min Bal Bal 54 max
    Nb+Ta 10–25 20–30 05–30 315–415 315–415 10 max 09–14
    Fe 100 max 30 max 120 max 70 max 10 max 50 max 30 max 38–42
    Ti 10 max 075 max 040 max 060 max 005 max
    P 003 max 003 max 003 max 002 max 002 max 003 max 003 max 001 max
    S 0015 max 0015 max 002 max 003 max 0015 max 0015 max 0015 max 001 max
    Mo 05–25 80–100 80–100 80–100 80–100 18–22
    Co 90–150 100–150
    Cu 050 max 050 max 050 max 050 max 050 max 050 max 050 max 025 max
    B 00005–002
    Al 040 max 08 to 15 010–020
    N 002 max

    9842144

    Best Practice Guidelines for Design and Fabrication
    37
    Figure 34
    Effect of composition on the formation of carbide and embrittling phases at the fusion line
    between ferritic steels and commonly specified nickelbase filler metals (simulation
    conditions listed in Table 34) [24]
    Table 34
    Descriptions for the DMW simulations in Figure 34 [24]
    Simulation Ferritic Material Weld Metal PWHT12 Service Simulation3
    1
    Grade 91
    ENiCrFe2
    None 625°C (1157°F)
    2 675°C (1247°F) 625°C (1157°F)
    3
    760°C (1400°F)
    625°C (1157°F)
    4 550°C (1022°F)
    5 9Cr1Mo 625°C (1157°F)
    6 Grade 22 730°C (1346°F) 550°C (1022°F)
    7
    Grade 91
    ENiCrFe3
    760°C (1400°F)
    550°C (1022°F)
    8 625°C (1157°F)
    9
    ENiCrMo3
    550°C (1022°F)
    10 625°C (1157°F)
    11
    ENiCrCoMo1
    550°C (1022°F)
    12 625°C (1157°F)
    13
    ENiFeCr4
    (Code Case 2734)
    None 625°C (1157°F)
    14 675°C (1247°F) 625°C (1157°F)
    15
    760°C (1400°F)
    625°C (1157°F)
    16 550°C (1022°F)
    1 PWHT Post Weld Heat Treatment
    2 All PWHT simulations 4hour duration
    3 All service simulations 50000hour duration
    9842144

    Best Practice Guidelines for Design and Fabrication
    38
    Maintaining proper control through adequate quality assurance and inspection is often a
    particular challenge in the fabrication of DMWs An example of a DMW between Grade 91 and
    austenitic stainless steel 321H is shown in Figure 35 In this particular configuration the P91
    main steam pipe was buttered in the fabrication shop under wellcontrolled conditions using
    ENiCrFe2 as evidenced by the consistency of the composition In contrast the fieldwelded
    girth weld between the buttered P91 and 321H outlet header contains at least three different
    types of filler materials including compositions consistent with ERNiCrMo3 in the root and
    ENiCrFe2 and ENiCrFe3 in the fill passes Although the use of nonspecified filler materials
    did not contribute to the root cause of the failure in this case it does highlight the importance of
    quality control in ensuring that welding procedures are closely followed because in other cases
    even such minor seemingly unimportant variances can result in premature failure

    Figure 35
    Example of inadequate control of welding consumables in a girth weld made between a
    butter P91 main steam pipe and an austenitic stainless steel 321H outlet header
    A particular challenge in filler metal selection is associated with the Code of Construction and
    the emphasis on the phrase overmatching With respect to longterm behavior the time to
    failure of the filler material (Codes often recognize only strength) needs to exceed only that of
    the weakest constituent in the DMW Failures are often associated with the HAZ in the 9Cr
    CSEF steel or at the fusion line between the nickelbase filler material and the 9Cr CSEF steel
    In this regard the filler material performance does not need to exceed that of both parent metal
    constituents as might be typically specified Particular concern exists regarding the development
    of stress relaxation cracking in the stainless steel HAZ where the filler material exceeds the
    strength of both parent materials this is a particularly important consideration for highly
    restrained components
    Transition Pieces
    There are two types of transition pieces that have been used in DMWs pieces used to provide a
    transition between the differing thermal expansion coefficients of the 9Cr CSEF steels and
    austenitic stainless steels (these transition pieces are typically nickelbase or nickelironbase
    alloys) and pieces used to transition between the differing wall thicknesses in the 9Cr CSEF
    steels and austenitic stainless steel sections (as is common in tube applications)
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    39
    An austenitic nickelbase or nickelironbase transition piece of intermediate thermal expansion
    behavior between the 9Crtype steel and the austenitic stainless steel is neither required nor
    prohibited Where concerns exist regarding the potential risk of sensitization in the stainless steel
    section the use of a transition piece that does not form deleterious phases in the PWHT range
    (that is 700–780°C) is recommended for welds made in 9Cr CSEF steels Suitable alloys that
    are not susceptible to sensitization include alloys 800 800H or 617 However it should be noted
    that the use of these nickelbase alloys can introduce other complexities that need to be
    considered Of primary concern is the potential risk for stress relaxation cracking in service For
    alloy 617 there is a risk of agehardening during PWHT and service Alloy 625 will certainly
    harden and embrittle at typical service temperatures and should not be allowed for use in this
    application
    A transition piece between dissimilar thicknesses should be properly designed to accommodate
    the different wall thicknesses of the 9Crtype and austenitic stainless steels In the welding of
    an austenitic stainless steel to a 9Crtype steel the transition piece must be fabricated from
    either an austenitic stainless steel or an austenitic nickelbase or nickelironbase material that
    exceeds the strength of the 9Crtype steel Matching of dissimilar thicknesses in the minimum
    must conform to the requirements in ASME B&PV Code Section I PG421 [28]
    An example of an improper transition between dissimilar thickness piping in a main steam
    HRSG system is provided in Figure 36 In this example the expected service duration to
    damage was reduced by at least a factor of 3 due to poor design In contrast and shown in
    Figure 37 is an example of an acceptable transition between a low alloy steel and an austenitic
    stainless steel Because the fabrication of DMWs often involves a thickness transition due to the
    differences in the allowable stress values at a given temperature the historical practice was to
    use a transition such as shown in Figure 37 As this practice increases the number of welds
    OEMs have recently moved away from it since it is not mandated by ASME B&PV Code (or
    other construction codes) A summary of acceptable and unacceptable practices regarding the
    installation of a transition piece in a 9Cr CSEF steel to austenitic stainless steel DMW is
    presented in
    Figure 38

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    Wall Thickness
    (mm in)
    Hoop Stress
    (MPa ksi)
    Gr 91 Crossweld
    Minimum Life (hours)
    Reduction in Performance
    (Relative to 38 mm Wall Thickness)
    38 150 46 67 226000 1
    30 118 60 87 77000 3X
    29 110 62 90 70000 325X
    Figure 36
    Example of a poor transition between a 9Cr CSEF steel and an austenitic stainless steel
    (Note the provided estimates on lifetime were calculated using the EPRI Life Calculator
    use of the hoop stress in this evaluation is illustrative [29])
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    Figure 37
    Example of a proper use of a transition piece between a low alloy steel and an austenitic
    stainless steel (Note there was no observed damage after >137000 hours of operation at
    an outlet steam temperature of 540°C [1000°F])
    Figure 38
    Summary of acceptable and unacceptable design practices when installing transition
    pieces in 9Cr CSEF steel to austenitic stainless steel DMWs

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    Alternative Weld Geometries for DMWs
    Design and fabrication of DMWs should consider the benefits of using improved joint geometry
    For smalldiameter thinwall connections the use of a wider bevel or a wide weld cap depicted
    in Figure 39 is expected to improve performance when the inservice failure is creep
    dominated As best practice a wide weld cap should be used to offset the location of the oxide
    notch from the main portion of the fusion line and the HAZ where the longterm evolution of
    damage may localize The edge of the wide weld cap should thus be sufficiently removed from
    the outer regions of the HAZ for example a minimum of 050 in (127 mm) in width with
    respect to the original machined fusion line For butt DMWs in thinwall tubes (generally
    defined as ≤025 in [635 mm]) a wide (beyond the basic preparation) weld cap should be
    specified because a larger bevel angle or step weld will typically not offer an improvement in
    performance
    In thickwall components (generally defined as >050 in [127 mm] in wall thickness) where
    there is a severe consequence of a rupture a step weld configuration with a wide cap (Figure 39)
    should be considered to improve the damage tolerance of the connection The step should be
    located at a depth in the component that is onehalf of the wall thickness The intent of the step
    and widecap is to move the surface damage resulting from an oxide notch away from the fusion
    line and HAZ regions and offset subsurface macro damage that may initiate at the fusion line or
    HAZ The step effectively forces damage to propagate through stronger and more ductile parent
    material providing an increased potential for detection and should cracking occur a greater
    resistance to catastrophic failure and promotion of leakbeforebreak behavior This is
    summarized schematically in Figure 310 and for a feature crossweld creep test in Grade 91
    steel where damage is formed in the HAZ in Figure 311
    Although the alternative joint designs shown in Figure 39 are for full penetration welds backing
    rings are not prohibited in the welding of DMWs For operating modes where there is a risk of
    thermal fatigue backing rings should be avoided because they create a local stress concentration
    at the inside diameter Thickness transitions where backing rings are installed should be carefully
    designed and fabricated to eliminate the potential for local and excessive reductions in wall
    thickness
    It should be noted that for situations in which fatigue is the dominant damage mechanism stress
    concentrations linked to the weld profile (such as a wide cap pass) have been shown to reduce
    performance Thus in these cases the joint design must be selected to complement the operating
    characteristics of the component and account for the complexities that may exist in longterm
    operation
    Good practice also requires that the geometric transition associated with ID counterbore is offset
    from the weld fusion line by a minimum of 050 in (127 mm) to avoid the local stress
    enhancement from the counterbore influencing the stress at the fusion line

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    Figure 39
    Examples of better practice joint designs for DMW fabrication between a stainless steel to
    either a CrMo or 9Cr CSEF steel
    Note WCL weld centerline
    Figure 310
    Schematic representation of the formation of damage in a step weld where damage forms
    in the upper portion of the weld and is forced to propagate through parent material
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    Figure 311
    Feature crossweld creep test example in Grade 91 steel defect formed in the HAZ is
    being driven through more deformationresistant and creep ductile parent metal and as
    evidenced by the large deformation
    Welding Technique
    For thinsection weldments especially those in reheat or HRSG applications where the tube wall
    thickness is generally <5 mm (020 in) a minimum number of weld fill passes or limitation in
    the heat input should be considered to avoid excessive heat input General guidelines may be
    used to place a minimum limit on the number of layers (3 for example) or limit the size of the
    electrode (no larger than 32 mm [0125in] diameter for SMAW process or 24 mm [0093 in]
    for manual GTAW process)
    For tubetotube weldments there has been an increase in the use of automated GTAW processes
    to improve quality and control over the welding of DMWs in the fabrication shop However the
    process must be controlled to reduce the interpass temperature and risk of excessive ferrite
    formation in the HAZ of the 9Cr steel In many cases the interpass temperature has not been
    well controlled leading to excessively slow cooling rates ferrite formation and very wide
    HAZs Recommended interpass temperature control for 9Cr CSEF steels is 315°C (600°F)
    Where the manual GTAW process is used defects have been noted in the qualification of
    welding procedures using weld rod ≥32 mm (0125 in) [30] in diameter For qualifications of
    similar weldments in Grade 91 steel (such as ER90SB9type filler material) use of a manual
    GTAW process and a rod diameter of 40 mm (0156 in) has introduced difficulties in passing
    procedure qualification records see Figure 312 [31] The same 40mm (0156in) GTAW cut
    rod that produced the failures in Figure 312 was given to a highly skilled welder This highly
    skilled welder also failed the side bend tests These observations reinforce that even with using
    highly skilled welders attentive diligence when welding with nickelbase filler materials must
    always be in place The basic technique disciplines of avoiding continuous GTAW filler metal
    feeding (aka stuffing) interpass cleaning to bright metal even if only heat tint is apparent and
    using pencil grinders (rear exhaust and nail point burr tips) to clean up weld bead toe edges serve
    to minimize many of the bend test issues found during the PQR tests Such attention to detail
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    should be regarded as best practice and mandatory when executing procedure qualifications
    welds in the fabrication shop and welds in the field To reduce the risk of introducing defects
    into the weld metal it is recommended to limit the diameter of manual GTAW weld rod to 25
    mm (0093 in) This precaution will reduce or eliminate the potential for lackoffusion defects
    through a reduction in the amount of filler material that can be fed into the weld pool

    Figure 312
    Failed bend test coupon where in the lower half of the weldment a manual GTAW process
    with 40mmdiameter (0156indiameter) solid wire was used [31]
    For thicksection weldments where the final layer at the cap of the weld contains multiple fill
    passes the general welding sequence where stringer beads are utilized should include an
    outsidein approach For the final layer at the cap of the weldment the last weld bead should
    not be placed on the fusion line of either parent material but rather in the center of the weldment
    in the nickelbase weld
    PostWeld Heat Treatment
    The application of PWHT in the fabrication of DMWs will reduce the peak welding residual
    stresses that inevitably result from arc welding processes It is recommended that PWHT be
    performed near the minimum of the allowable range and minimum allowable time for P15E
    material and as described in ASME B&PV Code Section I Table PW 395 [1]
    Note (2) For dissimilar metal welds (ie welds made between a PNo 15E Group 1 and
    another lower chromium ferritic austenitic or nickelbased steel) if the filler metal
    chromium content is less than 30 or if the filler metal is nickelbased or austenitic the
    minimum holding temperature shall be 1300°F (705°C)
    Although some Codes may allow a maximum PWHT temperature of 790°C or 800°C (1450°F or
    1470°F) the maximum recommended PWHT temperature is 770°C (1420°F) and in accordance
    with guidance in [32]
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    PostWeld Machining
    After welding the weld profile should be machined to remove excessive reinforcement at the cap
    of the weld and to eliminate excessive penetration at the root of the weld Removing these local
    features will reduce local stress concentrations that may exacerbate thermal fatigue damage in
    service Additional benefits include an increase in the time to oxide notch initiation and an
    improvement in performing future inspections using ultrasonicbased techniques
    Welding Procedure Qualification
    Qualification of welding procedures to ASME B&PV Code Section IX can introduce difficulties
    when the austenitic stainless steel is an advanced grade such as Super 304H or 310HCbN (see
    Table 24 for compositions) Because qualification of procedures in part relies on the ambient
    temperature tensile strength of the materials if the parent materials requiring welding possess
    minimum ultimate tensile strengths ≥585 MPa (85 ksi) or ≥620 MPa (90 ksi) the failure in the
    tensile tests may occur in the weld metal at stress levels below these minimum requirements A
    comparison of minimum ultimate tensile strength for common parent materials and nickelbase
    filler materials is provided in Table 35
    Pertaining to DMW qualification ASME B&PV Code Section IX QW1531 states that in
    procedure qualifications the test specimen
    shall have a tensile strength that is not less than … (b) the minimum specific tensile
    strength of the weaker of the two if base metals of different minimum tensile strengths
    are used or (c) the minimum specified tensile strength of the weld metal when the
    applicable Section provides for the use of weld metal having lower room temperature
    strength than the base metal (d) if the specimen breaks in the base metal outside of the
    weld or weld interface the test shall be accepted as meeting the requirements provided
    the strength is not more than 5 below the minimum specified tensile strength of the
    base metal
    Note that ASME B&PV Code Section I does allow for clause QW1531 (c) to be invoked
    The results of elevatedtemperature tensile tests of all weld metal pads for commonly used
    nickelbase filler metals at 500°C (930°F) are presented in Figure 313 This graph illustrates that
    at servicerelevant temperatures the observed values in the weld metal are well above that of the
    representative parent materials requiring welding (for both minimumspecified and typically
    measured values) Hence the issue regarding ambient temperature ultimate tensile strength is
    simply procedurerelated and not performancerelated
    Where issues are encountered in the qualification of nickelbase filler materials for DMW
    applications the current Code rules will allow for the following solution In highstrength
    combinations where both parent materials exhibit a minimum ultimate tensile strength value
    ≥585 MPa (85 ksi) a higherstrength filler material such as ENiCrMo3 or ERNiCrMo3 can be
    utilized for the purposes of the procedure qualification record For shop or field fabrication a
    different F No 43 filler material can be substituted in manufacturing of the component or weld
    and in accordance with Code rules (for example use of ERNiCr3 or ENiCrFe2 or other)

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    Table 35
    Minimum ultimate tensile strength values for materials commonly utilized in DMWs [20 25]
    Material (SA213) Minimum Ultimate Tensile Strength
    (MPa ksi)
    CSEF Steels
    T91 585 (85)
    T92 620 (90)
    T93 620 (90)
    Austenitic Stainless
    Steels
    300 Series Hgrade 520 (75)
    347HFG 550 (80)
    Super 304H (Code Case 2328) 585 (85)
    310HCbN (HR3C) 655 (95)
    NickelBase Filler
    Metals
    ERNiCr3
    550 (80)
    ENiCrFe2
    ENiCrFe3
    EPRI P87
    ENiCrMo3 750 (110)
    ENiCrMoCo1 620 (90)
    Figure 313
    Measured ultimate tensile strength values at elevated temperature (500°C [930°F]) for
    common nickelbase filler materials (blue) Grade 91 steel plate (black) and 347H plate
    (red) and as compared to minimum values for T91 and T92

    T91 T92 347H Gr 91 ENiCrFe2 EPRI P87 ENiCrMo3 ENiCrMoCo1
    250
    300
    350
    400
    450
    500
    550
    600
    650
    Measured ValuesTensile Strength (MPa) at 500

    C
    (
    930

    F
    )
    Minimum Values
    98421449842144
    41
    4
    EXAMPLE APPLICATIONS AND INSERVICE
    OPERATION EXPERIENCE
    This section presents considerations for specific component applications where DMWs have
    historically been used in boiler applications An emphasis in this report has been placed on
    documenting previous failures to provide examples of situations that should be avoided
    Consideration of the recommendations for improved methods reinforce the guidelines given in
    Section 3
    TubetoTube Butt Welds
    The use of a tubetotube DMW in a location with insufficient flexibility is likely to enhance
    stress at the weld and thus result in premature failure a typical example is shown in Figure 41
    A second failure in a tubetotube DMW is provided in Figure 42 In this second example the
    failure was due to oxide notching in the leading edge tubes nearest to the flame As the tube is
    sufficiently thick (such as installed in a superheater pendant) the propagation of the oxide notch
    through the wall thickness would have required a source of thermal cycling This cycling was
    later determined to be due to soot blower operation in the vicinity of the DMW
    Figure 41
    Premature failure of a DMW where the DMW was placed at the stubtotube weld in a
    location of poor flexibility [18] (A – location of failure in the component B – fracture
    surface C – polished macro image in etched condition showing fusion line dominated
    damage Reference operating conditions 560°C [1040°F] 270 bar [3915 psi] and failure in
    ~50000 hours)

    9842144

    Example Applications and InService Operation Experience
    42
    Figure 42
    Local throughwall steam leak in 103000 hours in a superheater tube DMW between T91
    and austenitic stainless steel 347H with 548mm (215in) OD and a 894mm (0352in)
    nominal wall thickness
    In Figure 42 note that there is matching of the ID of the butt weld locally reducing the wall
    thickness in the T91 The advance of the oxide notch through the wall thickness was estimated
    through oxide dating and is shown in the righthand image The reference operating conditions for
    the failure in Figure 42 include 538°C (1000°F) steam temperature 286 barg (4147 psig)
    estimated tube ID metal temperature 580°C (1075°F) estimated tube outside diameter metal
    temperature 650°C (1200°F)
    ThickSection Welds
    Operating temperature and cycling are identified as important factors in establishing the in
    service behavior of thicksection DMWs (generally defined here as in components where the
    wall thickness is >050 in [127 mm]) Detailed modeling has shown that a concentration of
    strain can develop at the fusion line between the nickelbase weld and the Grade 91 steel This
    concentration is exacerbated under operating conditions where the temperature of the component
    is <550°C (1022°F) This occurs due to the elevated temperature characteristics of 9Cr steels
    which exhibit very low creep rates at 540°C (1000°F) as recognized by ASME B&PV Code
    Section IID where timedependent properties are first mandated at this stated design
    temperature
    A number of unexpected and premature failures have been reported for components operating in
    this lower temperature regime such as documented in Figure 43 [16 33] The strain localization
    in the component shown in Figure 43 was not only enhanced by the operating temperature but
    also by the fact that this DMW was placed at a terminal point in the piping system a location of
    inherently low flexibility
    An example of a DMW in a thicksection piping system where the operating temperature was
    >550°C (1022°F) is provided in Figure 44 [16] In this case cracking occurred after a
    considerably longer period of service and creep damage was present both at the fusion line and
    in the HAZ
    9842144

    Example Applications and InService Operation Experience
    43
    Figure 43
    Premature failure of a DMW in a thicksection 321H outlet header to P91 piping system
    (AB – polished macro image in etched condition C – extent of OD connected cracking
    D – micro image of fusion line dominated damage ahead of the main crack) [16]
    Note Figure 43 is also an example where the P91 was buttered in the fabrication shop the
    DMW was placed at a terminal point and operation was at relatively low temperature Reference
    operating conditions 540°C (1004°F) operating pressure unknown and failure in ~40000
    hours
    Figure 44
    Premature failure of a DMW in a thicksection 304H to P91 piping system [16] (A – polished
    macro image in aspolished condition B – micro image of fusion line and HAZ damage)
    Note Figure 44 is also an example of a poor consideration in design where the transition
    between inside diameters should not have involved machining the P91 Reference operating
    conditions 566°C (1050°F) 109 bar (1580 psi) and failure in ~125000 hours
    9842144

    Example Applications and InService Operation Experience
    44
    TubetoHeader Connections
    DMWs in small bore connections to thicksection components (such as stubtoheader welds or
    drain lines) should never be placed at the weld joining the small bore tube to the largediameter
    pipe A case study in [19] is reviewed in Figure 45 detailing cracking observed on the stainless
    side of a P91 reheater outlet header stub tube weld to TP347H In this example hundreds of
    leaks and flaws were found in similar welds in a oncethrough supercritical boiler outlet header
    (main steam operating conditions of 570°C [1060°F] and 250 barg [3625 psig]) A significant
    contributing issue to the failures was the design of the roof penetration In this unit the tube
    collar penetration through the boiler roof membrane was welded to the tubes resulting in the
    component expanding and contracting as the roof membrane moved with thermal transients This
    type of arrangement creates very complex loading distributions across the header and any
    temperature imbalance across the tubes will accelerate the development of damage
    Due to the extent of damage the repair process required a major outage lasting ~45 days It was
    decided to remove and modify the stubtoheader weld using a T91 transition piece between the
    P91 header and TP347H tubing This required repair of all 1070 stub tubes To reduce the
    restraint caused by the rigid welding of the tubes to the tube collar roof penetration these collars
    were cut to allow the tube to float and removable roof seal blocks were installed A number of
    similar failures have also been documented in recently constructed units where DMWs were
    placed at stubtoheader connections between P91 and P92 to austenitic stainless steels

    9842144

    Example Applications and InService Operation Experience
    45
    Figure 45
    Examples of cracking in a DMWs between a P91 hot reheat header and austenitic stainless
    steel 347H (Note that the cracking was observed at the weld toes on the TP347H side of
    the DMW The TP347H measured 508 mm [2 in] OD and 4 mm (0157 in) wall thickness
    Reference operating conditions 570°C [1060°F] 50 barg [725 psig])

    9842144

    Example Applications and InService Operation Experience
    46
    Attachments to Piping and Headers
    The welding of stainless steel lugs or hangers to ferritic or CSEF headers or piping systems
    should be prohibited Under no circumstance should this type of DMW be permitted in
    fabrication or construction In some applications the attachment can be expected to operate at
    elevated temperature due to the vestibule gas temperature However these attachments are
    typically sufficiently thick that oxidation is not an issue and the attachments will experience
    cooling due to heat conduction to the header or piping system In this regard the design
    emphasis should be making the weld and component sufficiently thick to accommodate in
    service corrosion and fabrication should include a sufficiently sized full penetration fillet weld
    (that is full penetration with respect to the throughthickness orientation of the attachment)
    An example of a failure of a 304H stainless steel attachment to a Grade 91 steel pipe in an HRSG
    system is provided in Figure 46 [16] Note that although the loading on this type of attachment
    should have been very low the combination of the load and operating characteristics of the unit
    (for example thermal cycling) was sufficient to cause failure in <100000 hours of operation at
    540°C (1000°F)
    Figure 46
    Premature failure of a DMW in a stainless steel 304H attachment to P91 [16] (A – macro
    image of failure B – failed stainless steel attachments C – polished macro sample
    showing fusion line dominated failure between the nickelbase weld and P91 Reference
    operating conditions 540°C [1004°F] and failure after 12 years of service)

    9842144

    Example Applications and InService Operation Experience
    47
    Attachments to Tubing
    There are three key considerations when stainless or nickelbase attachments are required to be
    welded onto a 9Cr CSEF tube The first involves the expected loading and failure mode that
    may occur at the attachment Past experience for low alloy steels has shown that the use of a
    weak filler material that is more similar to the ferritic tube in composition may provide a benefit
    with respect to the failure location by moving the failure location away from the pressure
    boundary [34] While little or no improvement in the lifetime of the connection may be realized
    forcing a failure to occur in the weld metal or attachment rather than in the pressure boundary of
    the tube is of significant benefit with respect to both boiler operation and the complexity of
    repairs A common misconception in welding dissimilar materials is that that gross cracking can
    develop as a consequence of dilution In this respect a higheralloy material should never be
    welded with a loweralloy material such as would be the case in using a filler material matching
    to the 9Cr CSEF steel (such as ER90SB9 or ER80SB8) to join a 9Cr CSEF steel to an
    austenitic stainless steel This point is summarized below
    …It has generally been the rule in welding that because of dilution problems a higher
    alloy should not be welded with a lower one The successful welding of an austenitic
    material to a ferritic one with a ferritic electrode seems to break this longstanding
    rule…Secondly in some applications such as austenitic support attachments to
    superheater tubes it removes the point of high stress due to expansion from the colder
    face of the tube to a higher temperature zone having less temperature cycling and
    consequently lower operating stress Also as previously discussed it improves the
    metallurgical transition between the high alloy and the low one This results in
    advantages in the expansion stress problem and eliminates the sharp transition from a
    hard to a soft zone at the fusion line… [10]
    To confirm that DMWs could be made using a filler material nominally matching to the 9Cr
    CSEF steel and therefore potentially be applicable to attachmenttype welds a series of thick
    section dissimilar weldments were made in 20mmthick (075inthick) pipe using AWS
    type B8 and B9 filler metals The results are provided as macro images of crosssections for
    each weldment in Figure 47 Although it is not an explicit recommendation of this document to
    utilize a ferritictype filler material for welding austenitic stainless steel or nickelbase
    attachments to 9Cr CSEF steels there is prior historical evidence and recent investigations that
    support more investigation into this approach
    9842144

    Example Applications and InService Operation Experience
    48

    Figure 47
    Examples of thicksection weldments in P91 to Alloy 800A and stainless steel TP347H
    made with filler metals nominally matching in composition to P91 and free of welding
    defects
    A second consideration in welding attachments to tubes especially for thinwall tubes (generally
    defined as ≤025 in [635 mm]) is the control of heat input Due to failures that have been
    observed in the industry it is highly recommended that the selected welding procedure be
    suitably controlled so as to limit the throughthickness extent of the HAZ in the 9Crtype steel
    tube In practice control of the heating input is often difficult as different predictive equations
    exist and measurement of travel speed for common manual processes is inexact without a
    dedicated logging device However simple controls can be placed to limit the heat input the size
    of the welding consumable the welding technique (that is specification of stringertype beads)
    and training of the welder to prevent excessively slow travel speeds during the welding
    sequence A comparison of the depth of penetration and the resulting HAZ is compared in
    Figure 48 for different diameter nickelbase electrodes and the SMAW process

    Figure 48
    Bead on plate measurements for nickelbase filler metal ENiCrFe2 deposited on a Grade
    91 steel plate (Reference operating parameters 25 mm [75A 20V 63 ipm] 32 mm [100A
    26V 63 ipm] 40 mm [125A 26V 53 ipm])
    9842144

    Example Applications and InService Operation Experience
    49
    Figure 49A provides an example of a poorly controlled welding procedure which resulted in
    tube rupture due to a local seam weld being created in the tube In some cases such as for
    reheater tubing where the wall thickness may be <0150 in (38 mm) the extension of the HAZ
    through the tube wall may be unavoidable and for these types of applications the use of
    mechanical attachments or thickerwall tube segments should be considered In the example
    provided in Figures 49B and 49C the tube dimensions were consistent with superheater tubing
    even in this example the poor control of the welding procedure was sufficient to result in a HAZ
    that clearly penetrates to the ID of the tube (Figure 49C)
    Figure 49
    Examples of poorly controlled welding procedures premature failure of a DMW in a
    stainless steel sliding spacer to T91 in a final superheater (A) (failure mode in this example
    is attributed to a local seam weld effect) location of hardness map through the T91
    thickness at the attachment weld (B) and color hardness plot showing the extension of
    the HAZ from the attachment weld through the T91 (C) [19] For Figure 49 the reference
    operating conditions are as follows highest tube operating temperature 565°C (1050°F)
    Note design operating temperatures vary from 524 to 579°C (975 to 1075°F) 197 bar (2855
    psi) and initial failures in ~36000 hours Oxide thickness measurements (~100 microns in
    total scale thickness) suggested that the investigated weld in Figure 49B (note no
    damage) was operating ~535°C (995°F) and consistent with the expected design
    conditions
    An example of variable damage progression within similar attachment welds in a Grade 91 steel
    hot reheat system is provided in Figure 410 In the example images provided it is clear that
    there is evidence of creep damage in the T91 HAZ (Figure 410AB) oxide notching and jacking
    (Figure 410CD) and fusion line dominated damage (Figure 410EF) Depending on the
    specific details of operating temperature local cycling quality of weld and other potential
    factors the damage progression can be variable The identified failures in the example provided
    in Figure 410 prompted replacement of the component after only ~75000 hours of operation
    due to unpredictable failures and an inability to keep the unit online
    9842144

    Example Applications and InService Operation Experience
    410

    Figure 410
    Example of multiple damage modes in attachment welds in a Grade 91 steel hot reheat
    system (A B – cavitation in the HAZ resulting from the T91 to attachment weld C D –
    oxide notching and jacking E F – fusion line damage) For Figure 410 the reference
    operating conditions are as follows ~70000 hours of operation to removal of samples
    Oxide dating suggested an operating temperature in the vicinity of these attachments of
    590–600°C (1095–1110°F)

    9842144

    Example Applications and InService Operation Experience
    411
    The last consideration in welding attachments to tubes is to minimize welding defects and ensure
    full penetration through the attachment Because the attachment will be in operation at a higher
    temperature than the tube it is connected to and heat will flow from the attachment to the tube an
    insufficient weld penetration (such as shown in Figure 411A) or gross welding defects (such as
    shown in Figure 411B) will locally increase the temperature at the attachment This will
    increase the risk of premature failure because of a combination of higher attachment temperature
    and increased thermal stress due to the greater temperature mismatch between the attachment
    and the tube To ensure that the tubetoattachment weld is operating at as low a temperature as
    possible the weld should be full penetration
    Figure 411
    Examples of unacceptable fabrication practices in T91 attachment welds (A – incomplete
    penetration of the sliding spacer to T91 weld B – gross welding defects between the
    sliding spacer to T91 weld In Figure 411 the reference operating conditions are as
    follows highest tube operating temperature 565°C (1050°F) Note design operating
    temperatures vary from 524°C to 579°C (975°F to 1075°F) 197 bar (2855 psi) and initial
    failures in ~36000 hours

    9842144

    Example Applications and InService Operation Experience
    412
    ASME B&PV Code Section I guidelines for welding attachments to pressureretaining items is
    provided in Figure 412 The guidance in most Codes only addresses the size of the weld from a
    load bearing capacity Where attachments such as slip spacers are welded to tubes there is no
    formal guidance because these are not load bearing components However it is clear that the
    particular details of attachments especially in tubing need to address several key aspects to
    achieve a number of operating characteristics as summarized below
     Where additional loading is a concern the fillet should be oversized
     The welding process should be well controlled to produce a full penetration weld minimize
    welding defects and minimize the depth of the HAZ in the tube or pipe This is summarized
    in Figure 413 These precautions are vital to ensure adequate heat transfer (such as cooling)
    of the attachment and minimization of local increase in temperature at the pressure boundary
     Stress concentrations at the weld toes should be minimized or eliminated through controlled
    machining or grinding

    Figure 412
    ASME B&PV Code Section I guidance for welding lugs hangers and brackets to shells
    drums and headers [35]

    Figure 413
    Summary schematic of important details and notes for welded attachments on 9Cr CSEF
    steel tubing
    9842144

    Example Applications and InService Operation Experience
    413
    Thermowells and Steam Sampling Nozzles
    Conventional design approaches for making stainless steel thermowell to 9Crtype piping
    system connections typically utilize a bosstype connector and a DMW seal or fillet weld The
    bosses are typically produced to the designs provided in ASME B&PV Code Section I PW16
    (see Figure 414) and they may or may not be threaded Example failures of this type of
    configuration in a thermowell and a steam sample nozzle are shown in Figures 414 and 415
    Note that in both cases the cracking occurred along the nickelbase weld metal to Grade 22 or
    Grade 91 fusion line
    The failure in Figure 415 occurred after ~85000 hours of operation Upon failure the utility
    performed a detailed review of the fleet including investigations of previous failures that may
    have gone unreported It was found that a previous failure in the utility fleet had occurred after
    ~40000 hours of operation A total of 24 similar probes were identified in the utility’s fleet in
    hot reheat and main steam systems and immediately inspected Six of these contained significant
    indications and a review of operating records revealed that another eight had failed since going
    into operation
    Historically DMWs placed in thermowells have also posed a significant problem in lowalloy
    steel systems Immediately following the catastrophic longseam failure at Monroe power station
    Unit 1 in 1986 all welded connections in the four units were inspected including thermowells
    As noted in the summary findings of the 24 total thermowells inspected 8 required repair and 12
    were replaced reinforcing the severity of the problem noted by the utility with respect to these
    welded connections [36]
    Because of the fundamental propensity and history of cracking the welding of stainless steel
    thermowells into 9Cr type piping systems should be prohibited Under no circumstance should
    this type of DMW be permitted in fabrication or construction The thermowell material should be
    matching to the 9Cr type material

    9842144

    Example Applications and InService Operation Experience
    414
    Figure 414
    Example of a DMW failure in a Grade 91 boss to stainless steel 316 thermowell

    Figure 415
    Failure of a steam sampling nozzle in a Grade 91 hot reheat HRSG system [37] (Reference
    operating conditions 570°C [1060°F] 483 barg [700 psig] failures in 40000 to 85000
    hours)
    Steam Flow Elements
    A catastrophic failure of a poorly designed stainless steel flow element to Grade 91 steel piping
    system in an HRSG application is shown in Figure 416 The fabrication of stainless steel flow
    elements in 9Cr steel piping systems can be complicated as shown in Figure 417 where a
    representative crosssection of a flow element in a P91 system is detailed In this example there
    is evidence of a nickelbase butter layer in the Grade 91 steel a butter layer applied to a portion
    of the 316 flow element and the girth weld made with a nickelbase filler metal It is also clear
    that a notch exists at the toe of the girth weld and at the Grade 91 fusion line

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    Example Applications and InService Operation Experience
    415
    Due to the number of repeat failures [38–40] which in some cases have been catastrophic the
    welding of stainless steel flow elements in piping systems should be prohibited Under no
    circumstance should this type of DMW be permitted in fabrication or construction The flow
    element should be fabricated from material matching to the 9Cr type steel It is also
    recommended that even if a matching 9Cr steel flow element is used the design selected—
    particularly for application in combined cycle plants—should avoid sharp corners and large
    changes in crosssection that can promote fatigue crack initiation

    Figure 416
    Catastrophic failure in a DMW between a Grade 91 main steam piping system to a
    stainless steel flow nozzle (Reference operating conditions operating temperature 507°C
    [945°F] 1055 bar [1530 psi] and failure after ~8 years) [40]

    Figure 417
    Macro view (A) and a crosssection (B) of a stainless steel 316 flow element welded in a
    P91 main steam HRSG piping system

    9842144

    Example Applications and InService Operation Experience
    416
    Drains
    Where drain lines in 9Cr piping system are fabricated from austenitic stainless steels the
    location of the DMW must be placed away from the main run piping component and in a region
    with sufficient flexibility Failures have occurred where the DMW was located at the drain line
    to thicksection component weld or where the DMW was placed in a location in the immediate
    vicinity of the thicksection component In many cases positive material identification (PMI)
    testing of material in the drain line system revealed rogue material that had not been specified
    in the original design or drawings In all cases the best practice option is to remove the stainless
    steel from the drain line and remove the presence of the DMW in this application

    9842144
    51
    5
    METALLURGICAL RISK
    9Cr CSEF Steels
    The introduction of unnecessary metallurgical risk factors at the fusion line (such as those that
    result in the initiation and growth of embrittling phases) should be avoided through proper
    consideration of the selected filler material for the DMW For DMWs operating in components
    that operate at elevated temperature in the creep regime for 9Cr steels (generally >550°C
    [1022°F]) the impact of metallurgical risk due to relative high impurity content in the parent
    metal may be premature failure in the 9Cr HAZ [41–43] As detailed in Figure 44 when
    9Cr steel DMWs are operating well within the creep regime and well designed (that is not
    excessive system loads or stressconcentrating features) there is an evolution of creep damage in
    the 9Cr HAZ and under some scenarios the lifelimiting factor will be the HAZ To ensure
    sufficient longevity of the DMW the composition of the 9Cr steel in the DMW should be
    procured to the proposed ASME Code Case for Grade 91 steel (ASME B&PV Code Case 2864)
    which imparts new maximum limits for tramp and impurity elements such as As B Cu S Sb
    Sn and W as well as more stringent maximum limits for Mn Ni S and Si [44] Ongoing
    research is elucidating the factors contributing to damage in the HAZ and more information can
    be found in [45]
    Austenitic Materials
    Austenitic stainless steels when welded with overmatching filler materials and subjected to
    highrestraint applications can be susceptible to stress relaxation cracking In particular
    materials that have deliberate elemental additions of precipitationstrengthening elements may
    develop shortterm damage as a consequence of PWHT (depending on the manufacturing
    sequence of the DMW) or after operating in service These issues are typically observed in the
    HAZ a few grains removed from the fusion line and where a sufficiently high temperature is
    achieved during welding to resolutionize carbonitride particles The nature and extent of risk
    factors have been reviewed and evidence generally suggests the following contributors to
    susceptibility [46]
     Large grain size in the parent material
     Presence of precipitationstrengthening elements such as Nb Ti and V
     Impurity content that can lower creep ductility
     Exposure (either service or PWHT) in a region of reduced creep ductility

    9842144

    Metallurgical Risk
    52
    It is particularly important to note that although some types of stainless steels such as 304H are
    generally regarded as being immune from stress relaxation cracking the actual content of
    precipitationstrengthening elements may be highly variable since the TP304H specification does
    not require purposeful control of these types of elements Stress relaxation cracking has been
    documented in stainless steel H grades such as 316H 321H and especially in 347H With
    respect to candidate nickelbase alloys for use in transition pieces stress relaxation cracking has
    been observed in alloy 800800H and alloy 617
    Longterm evolution of damage is typically controlled by the presence of inclusions particularly
    MnS content [47] as well as the development of embrittlement phases such as sigma [48] The
    evolution of the sigma phase in austenitic stainless steels can take 10000s to 100000s of hours
    in service Formation of damage due to longterm microstructural evolution in the austenitic
    material (either stainless or nickelbase) is expected to be a secondorder concern The
    controlling issues in DMWs are expected to be shortterm issues (<10000h of operation) due to
    stress relaxation cracking in the austenitic material or development of damage in the longerterm
    at the fusion line against the 9Cr steel or in the 9Cr steel HAZ

    9842144
    61
    6
    INSPECTION
    Given that temperature and temperature cycling are contributors to DMW failures it is
    recommended that the metal temperature history be monitored in the vicinity of DMWs
    Depending on the location orientation and application of the DMW it may be advantageous to
    install thermocouples around the circumference or it may be sufficient to infer temperature from
    a nearby instrument location For scenarios where DMWs are present in large numbers (such as
    in tubing applications attachments or tube butt welds) it is recommended to install sufficient
    monitoring thermocouples on outlet terminal tubes (such as above roofline prior to entry to
    header) to provide information on temperature imbalances and excursions that can be used in
    fitnessforservice evaluations
    Because of the uncertainty in failure time DMWs that pose an elevated risk to personnel injury
    or collateral damage should be routinely inspected through the life of the connection In some
    cases such as in thermowelltype connections the position of the DMW relative to walkways
    and critical equipment should be evaluated For the failure presented in Figure 414 the utility
    found that many thermowells were adjacent to walkways and hightraffic areas prompting the
    utility to relocate these connections and plug the existing locations
    Suggested inspection methods require the use of qualified personnel and qualified procedures
    that have been explicitly developed for inspection of DMWs At a minimum inspection should
    include both surface techniques such as surface penetrant testing and volumetric evaluation of
    the fusion line and 9Cr HAZ using linear phased array ultrasonic testing around the entire
    circumference of the DMW A number of special considerations for DMWs must be addressed
    when the decision to inspect a DMW is made
     Does the NDE contractor have a DMWspecific procedure Refracted longitudinal waves are
    recommended due to attenuation and beam redirection from differences in sound
    transmission in DMWs
     How is calibration performed and what is the reference standard for identification of
    indications
     What is the contractor’s experience with DMW inspections and for Grade 91 Since damage
    may be observed at the fusion line andor HAZ it is important to review vendor references
    and case studies
     What damage is expected to be detected
     Is the NDE contractor aware of oxide notching fusion line cracking HAZ damage and so
    forth Damage can progress by these three mechanisms and may be present in the same
    DMW such as shown in Figure 61 Furthermore since the damage is not likely to be
    uniform through the thickness of the component it is important to ensure that the potential
    extent of damage is fully appreciated
    9842144

    Inspection
    62
     Are scan plans available
     Is encoded data provided for future reference A fingerprint is vital for comparison with
    future inspections and independent assessment(s) when necessary
     What engineering support is provided for disposition The NDE practitioner andor vendor
    should not be allowed to autonomously disposition indications This should be done jointly
    with appropriate engineering review to understand and when necessary to question the
    validity of the provided results
    A summary of common inspection techniques and the validity of each technique with respect to
    inspection of DMWs is provided in Table 61
    Table 61
    Validity and comments for common inspection techniques that may be applied to DMWs
    Technique Location Valid Comments
    Positive Material Identification (PMI) Surface  Confirm base metal and weld filler
    alloy types
    Penetrant Testing (PT) Surface  Surface cracking in weld fusion line
    or base metal
    Magnetic Particle (MT) Surface  Not valid due to nonmagnetic weld
    deposit
    Eddy Current (EC) Surface Needs specific setup for DMW
    Replica (REP) Surface 
    Surface metallurgical condition
    significantly affected by oxide
    notching and damage is subsurface
    Hardness (HRD) Surface  Limited value for indication of
    damage
    Ultrasonic (LPA) Volumetric  Detect cracking on fusion line
    Time of Flight Diffraction (TOFD) Volumetric 
    Challenging due to sound
    transmission through stainless steel
    and weld deposit
    Radiography (RT) Volumetric  Not recommended for detection of
    fusion line service damage
    Sample (SAMPLE) Volumetric  Verification of indications and detailed
    analysis of damage mechanisms

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    63

    Figure 61
    Example of damage through the wall thickness of a low alloy steel to stainless steel DMW
    (nickelbase filler metal) In Figure 61 note that depending on the type and quality of
    performed inspection the results may indicate that this weld is at end of life a stated life
    fraction or contain no damage aside from the oxide notch at the surface
    98421449842144
    71
    7
    LONGTERM PERFORMANCE MICROSTRUCTURAL
    EVOLUTION AND LIFE MANAGEMENT
    LongTerm Performance
    Previous attempts to develop lifing tools such as PODIS for low alloy steel DMWs have been
    largely unsuccessful as they have proven difficult to reduce uncertainty in operating parameters
    stresses and other factors to acceptable levels With respect to 9Cr steel components where
    damage is expected to be creepdominated a number of studies including EPRI evaluations
    have shown that whether failures are at the fusion line in the HAZ or mixed mode failures in
    DMWs fall within the scatterband of HAZdominated failures in similar welds (similar in this
    context means a 9Cr steel welded to itself using a nominally matching filler material such as
    ER90SB9 for welding Grade 91 steel) As shown in Figure 71 the EPRI and literature database
    of fusion line failures in DMW laboratory crossweld creep tests overlaps with the crossweld
    database for Grade 91 HAZ failures in similar welds In these cases use of the EPRI Life
    Calculator tool may be appropriate to provide an indication of remaining life

    Figure 71
    Comparison of reported fusion line failures [14 49–55] in Grade 91 DMWs to Grade 91 HAZ
    failure database and meanminimum relationships for the database (Note the fusion line
    failures in this comparison were produced as a consequence of wellcontrolled creep tests
    in a laboratory)

    19500 20000 20500 21000 21500
    50
    100
    150
    200
    Grade 91 HAZ Failure Database
    Grade 91 HAZ Failure Mean20
    Laha et al
    Yamazaki et al
    Fukahori et al
    EPRI 3002003363
    EPRI 1018991
    EPRI 3002007219
    Stress (MPa)
    LMP (273 + T C)(20 + log (tR hours))
    Grade 91 HAZ Failure Mean
    Grade 91 DMW Fusion Line Failures
    Source
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    LongTerm Performance Microstructural Evolution and Life Management
    72
    It is important to highlight that premature accelerated service failures such as those reviewed
    here and in [16] do not show a correlation with the calculated loading from pressure alone (for
    example whether the hoop or axial stress is assumed to drive cracking) These failures are
    concentrated at the fusion line and are observed in components operating ≤565°C (1050°F) For
    properly designed DMWs there is increasing evidence that axial stress drives the evolution and
    propagation of creep damage at the fusion line and this is supported by multiaxial test
    evaluations such as in [51] However it is clear that the review of inservice failures indicates a
    significant contribution from a secondary source that cannot be creepdominated or enhances the
    evolution of local creep damage at the fusion line Thus the assessment and lifing of DMWs in
    components operating at lower temperature and in a regime that is marginally affected by creep
    is particularly challenging In this regard operation of DMWs in 9Cr steels and in
    temperatures ≤550°C (1022°F) should be carefully assessed particularly for thicksection or
    highly restrained component geometries
    A significant fraction of DMWs in future 9Cr steel components will be placed in conditions
    that are expected to be creepdominated To evaluate the performance of a range of fabrication
    and design variables a number of welds are currently being tested to understand the influence of
    filler metal PWHT and joint design on performance Figure 72 illustrates several such weld
    configurations These welds were fabricated free of welding defects creep test results at standard
    crossweld creep conditions are reported in Table 71 The results presented in Table 71 show
    that most of the DMWs achieve sufficient performance when compared with the expected
    minimum and mean lives for HAZ failures in Grade 91 steel similar welds and the results of a
    control test (DMW5A) fabricated using a conventional manufacturing route However an
    important consideration in performance is not only when a DMW may fail but also how the
    DMW will fail This is addressed in Table 72 which compares the amount of observed tertiary
    creep in the straintime data for the failed crossweld creep tests at 625°C (1157°F) and 80 MPa
    (116 ksi) The failure behavior for singlevee welds with an aligned HAZ is generally poor
    unless a ferritic filler material is used Where a nickelbase filler material is specified there is
    clear merit in the application of a steptype configuration to increase the life fraction spent in
    tertiary creep shown by the comparison of DMW1B and DMW1C to DMW1D In the
    fabrication of components it is vital to understand the factors that impact performance not only
    from a lifetime perspective but also in how the component may fail In the case of DMWs
    failure must be regarded as an eventuality within the lifetime of the plant due to the variable
    nature of inservice performance More details can be found regarding the assessment of DMW
    performance in [18 55]

    Figure 72
    Fabrication of screening DMWs for evaluation in creepdominated test conditions
    9842144

    LongTerm Performance Microstructural Evolution and Life Management
    73
    Table 71
    Results of uniaxial feature crossweld creep testing for DMWs manufactured in P91 steel to the approaches provided in
    Figure 72
    ID Filler Metal and Welding Procedure 625°C (1157°F)
    80 MPa (116 ksi)
    625°C (1157°F)
    60 MPa (87 ksi)
    600°C (1112°F)
    80 MPa (116 ksi)
    Gr 91 Crossweld Min Life [29] 939 3700 5690
    Gr 91 Crossweld Mean Life [29] 2720 10725 15080
    DMW1B Nickelbase + No PWHT 6730 [Gr 91 HAZ] >17850 >15000
    DMW1C Nickelbase + 745°C2h 4142 [Gr 91 HAZ] >13500 >12500
    DMW1D Nickelbase + Step + No PWHT 3917 [Gr 91 HAZ] >13600 >12500
    DMW2A B9 + 745°C2h 1625 [347H FL] 3903 2749
    DMW3A B8 + No PWHT 811 [347H FL] 2850
    Not Tested DMW3B1 B8 + Step + No PWHT 1515 [347H FL] 4729
    DMW4A Stainless + No PWHT 4201 [Gr 91 HAZ] >13600
    DMW5A Nickelbase butter + 745°C2h + Nickelbase 2409 [Gr 91 HAZ] >13300 >12500
    Color Key
    Below Grade 91 HAZ Min Between Grade 91 HAZ Min and Mean At or above Grade 91 HAZ Mean


    9842144

    LongTerm Performance Microstructural Evolution and Life Management
    74
    Table 72
    Comparison of results at 625°C (1157°F) and 80 MPa (116 ksi) for time to failure (such as when the sample fails) and the life
    fraction spent in tertiary creep (such as how the sample fails)
    Note DMW5A in this comparison is used as the reference fabrication case
    Specimen Welding Procedure Time to Failure
    (hours)
    Life Performance Relative to
    DMW5A
    Percentage of Life Spent
    in Tertiary Creep
    DMW1B Nickelbase + No PWHT 6730 28X 2
    DMW1C Nickelbase + 745°C2h 4142 17X 7
    DMW1D Nickelbase + Step + No PWHT 3917 16X 21
    DMW2A B9 + 745°C2h 1624 07X 15
    DMW3A B8 + No PWHT 811 03X 26
    DMW3B1 B8 + Step + No PWHT 1515 06X 25
    DMW4A Stainless + No PWHT 4201 17X 6
    DMW5A Nickelbase butter + 745°C2h + Nickelbase 2409 (10X) 14
    Color Key
    Below Grade 91 HAZ Min Between Grade 91 HAZ Min and Mean At or above Grade 91 HAZ Mean
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    LongTerm Performance Microstructural Evolution and Life Management
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    Microstructural Evolution
    The evolution of the fusion line microstructure in 9Cr DMWs results in the development of a
    narrow band of potentially creepweak material (that is ferrite) Strain localization may occur in
    this location and result in premature or accelerated damage rates This has been investigated
    microstructurally and analytically through computational structural mechanics in [24 56]
    The evolution of the asfabricated zones in a DMW can contribute to strain localization and
    accelerate the rate of damage development particularly for fusion line dominated failures
    In the asfabricated state there are no less than nine distinct regions in a DMW as depicted in
    Figure 75 and Table 73 The complexity of these zones particularly across the fusion line is
    shown in Figure 74 for a representative DMW under SEM imaging (Figure 74A) composition
    variation as shown for the distribution of nickel determined using energy dispersive spectroscopy
    (SEMEDS) (Figure 74B) and electron backscatter diffraction (SEMEBSD) for the phase
    boundaries (Figure 74C) and distribution of grain boundaries with a misorientation of 2–180°
    (Figure 74D)

    Figure 73
    Identified regions in a DMW adapted from the classic description by Nippes [57] and as
    modified and reported in [58]

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    LongTerm Performance Microstructural Evolution and Life Management
    76
    Table 73
    Descriptions of the regions in Figure 73 and specific to a DMW in a 9Cr CSEF steel
    where the weld metal is nickelbase
    Zone Designation Description Ref
    Unaffected Parent Metal In the context of this description the unaffected
    parent metal is a 9Cr CSEF steel
    Overtempered HAZ OTHAZ
    The grain structure remains similar to the original
    matrix in the parent material and coarsening of
    secondary precipitates
    [59]
    Partially Transformed
    HAZ PTHAZ
    Original matrix of the parent metal is only partially
    reaustenitized and incomplete dissolution of the
    preexisting secondary precipitates
    [59]
    Completely
    Transformed HAZ CTHAZ
    Original matrix of the parent metal is fully re
    austenitized and a complete dissolution of the pre
    existing secondary precipitates
    [59]
    Partially Melted Zone PMZ
    Exists in all fusion welds made in alloys since a
    transition from 100 liquid to 100 solid must
    occur across the fusion boundary
    [58]
    Fusion Line Defined as the boundary between the PMZ and
    UMZ regions across a DMW
    Unmixed Zone UMZ Consists of melted and resolidified base metal that
    does not mix with the filler metal [58]
    Partially Mixed or
    Transition Zone TZ
    In heterogeneous welds where the filler metal is of
    different composition from the base metal this
    would represent a composition transition from the
    [fusion zone] to the UMZ…in welds between
    stainless steels and lowalloy steels martensitic
    structure may form in the transition region that does
    not occur elsewhere in the weld
    [58]
    Fully Fusion Zone
    In the context of this description the fully fusion
    zone shows a minimally diluted composition from
    the parent metal and is consistent with a nickel
    base filler metal

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    Figure 74
    Asfabricated condition of the fusion line between Grade 91 steel and a nickelbase filler
    metal (A – SEM image B – elemental distribution [shown for nickel] as determined using
    SEMEDS C – phase fraction of bodycentered cubic [BCC red] and facecentered cubic
    [FCC green] as determined using D – grain boundary fraction of grains with a
    misorientation of 2–180° and for the nickelbase FCC matrix and Grade 91 steel BCC
    matrix and as determined using SEMEBSD)
    The evolution of damage in the specific constituents at the fusion line are detailed in Figures 76
    through 78 On the basis of the evaluation of damage in exservice DMWs and wellcontrolled
    uniaxial creep tests where damage was observed at the fusion line it has been shown that all
    nickelbase filler materials are inherently susceptible to microstructural instability This
    instability results as a consequence of the welding thermal cycle and the compositional
    heterogeneity that results from the welding process The eventual transformation of the UMZ to
    ferrite is driven by a combination of time temperature and accumulation of strain under
    operating conditions PWHT has not been shown to contribute to this evolution for example the
    transformation of the UMZ to ferrite is observed for welds that have undergone PWHT or not
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    78

    Figure 75
    Microstructural regions and the phase(s) present composition details and association
    with inservice damage
    Examples of damage in the UMZ are provided in Figures 76 and 77 In Figure 77 EDS
    mapping of the nickel concentration shows that the creep damage (cavitation) is concentrated in
    a region with composition consistent with that of the Grade 91 steel Since the UMZ as defined
    in Table 73 should possess a composition that is nearly identical to that of the parent material
    the evidence suggests that damage is being accumulated in the UMZ The presence of martensite
    between the ferrite band and the fully fusion zone is consistent with a region where the
    composition is a mixture between the parent metal and fully fusion zone such as the described
    TZ in Table 73 It is interesting to note that in power generation creepdominated scenarios
    where the working fluid is steam damage is not seen to accumulate in the TZ This may not be
    the case for petrochemical or processing DMWs where the working fluid may be elevated in
    hydrogen content In these situations damage may be preferentially accumulated in the TZ
    martensitic band as a consequence of hydrogenassisted stress corrosion cracking

    Figure 76
    Location of damage in the ferrite band adjacent to the fusion line (Note nickelbase weld
    metal is on the left and Grade 91 steel on the right side of each image)
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    79

    Figure 77
    Location of damage in the ferrite band (left image) adjacent to the fusion line and
    composition mapping for nickel (right image) as determined using SEMEDS
    The elemental distribution of chromium in the region where damage was identified is provided in
    Figure 78 The evidence presented shows that the ferrite band is locally depleted in Cr which
    would lead to an increase oxidation rate (for example potential for accelerated oxide notching in
    this location)

    Figure 78
    Location of damage in the ferrite band (left image) adjacent to the fusion line and
    composition mapping for chromium (right image) as determined using SEMEDS
    The evolution of the UMZ to ferrite is shown schematically in Figure 79 Figure 79 explains
    two potential mechanisms for the evolution of the UMZ to ferrite—either as a consequence of
    PWHT or inservice operation As previously noted the UMZ has been observed in DMWs
    where PWHT was not applied and so it is not likely that PWHT contributes to the
    transformation of the UMZ PWHT may affect the stability of the fusion line with respect to the
    evolution of deleterious phases particularly where filler materials matching to alloy 625 are
    utilized Because DMWs in power generation materials require joining using arcwelding
    processes resulting in melting the necessary conditions for the formation and eventual
    transformation of the UMZ are a direct consequence of the compositional heterogeneity resulting
    9842144

    LongTerm Performance Microstructural Evolution and Life Management
    710
    from the welding process This means that all DMWs fabricated from nickelbase filler materials
    are inherently susceptible to this microstructural evolution Service exposure of DMWs leads to
    very local carbon diffusion (for example most of the carbon is expected to be tied up as stable
    carbonitrides having the form M23C6 or MX) and a driving force for carbide dissolution The
    maximum width of the observed ferrite band after inservice operation has consistently been <10
    µm (040 mils)
    Selection of nickelbase filler materials may delay the onset of the UMZ transformation to
    ferrite such as suggested in Figure 710 This delay is predicated by the use of a nickelbase
    filler material that is more matching in composition to the parent material such as EPRI P87 and
    through avoiding highly alloyed nickelbase filler materials It should be noted that the evolution
    of the fusion line constituents may not directly result in premature failure because the
    susceptibility to premature failure is a consequence of not only the potential metallurgical risk
    factor(s) present at the fusion line but an operational or loading condition that results in strain
    localization
    9842144

    LongTerm Performance Microstructural Evolution and Life Management
    711

    Figure 79
    Schematic of potential explanations for the evolution of ferrite in the unmixed zone [24]
    (Note that the supporting evidence shows that local carbon migration during service leads
    to carbide dissolution and the evolution of ferrite in the unmixed zone [such as the right
    hand side of the figure])
    9842144

    LongTerm Performance Microstructural Evolution and Life Management
    712

    Figure 710
    Results of computational analysis which show the sensitivity of the nickelbase weld to
    Grade 91 fusion line region to carbon migration (The sensitivity is represented as the
    variation in the minimum and maximum carbon values [weight percent] as a result of the
    analysis More details are available in [24])
    Life Management
    Due to the complexities and variability in the performance of DMWs there is a need to manage
    the life of DMWs using an integrated approach that addresses key details in the EPRI life
    management strategy outlined below
     Understanding of materials in the componentsystemfleet and damage mechanisms relevant
    to components of interest
     Appreciation of historical issues exemplar failures and statistical analysis of databases
     Development of better (ideally best) practice purchase specifications above the minimum
    Code of record (ASME ASTM or other)
     Guidelines for quality assurance during component manufacture and system fabrication
     When to look where to look how to look and damage tolerance disposition of damage and
    toolsmethods
     Alternative componentspecific methods for repair or replacement guidelines that exceed
    minimum Code designfabrication rules
     Effective technology transfer of information to the plants engineering departments and other
    relevant stakeholders in the process such as design codes architectengineers
    manufacturers materials producers and service providers

    16 13 14 15 7 8 5 4 1 3 2 6 11 9 12 10
    00
    01
    02
    03
    04
    05
    06
    07
    08
    [9Cr1Mo]

    Carbon at Fusion Line (Weight Percent)
    (Max Value Min Value)
    Simulation Number
    EPRI P87
    ENiCrFe3
    ENiCrFe2
    ENiCrMo3
    ENiCrCoMo1
    [P22]
    9842144

    LongTerm Performance Microstructural Evolution and Life Management
    713
    Approaches are needed for understanding whenwherehow to look for damage using riskbased
    inspection (RBI) methods aimed at riskranking the multitude of plants units or components
    systems within the unit or plant This is particularly important for DMWs placed in tubing
    applications where hundreds or thousands of DMWs exist across the unit and possibly in both
    thinsection (reheat) or thicksection (superheat) configurations When susceptible locations are
    identified it is vital that a wellengineered staged approach be embraced that uses a Level I II
    or III assessment to reduce risk and decrease the overall uncertainty in the current state of the
    component [60 61] A primary benefit of the risk rankingbased approach to life management is
    the staging of the evaluation by prioritizing componentsunitsplants and actions Such an
    approach can be planned costeffective and ensure safe operation of the plant
    DMWs in thicksection piping especially those that have been identified to pose a high
    consequence of failure should be subjected to the same Level I II or III assessment For these
    DMWs it is imperative to minimize the amount of uncertainty in the evaluation through detailed
    operating data and stress analysis Where unacceptable uncertainty exists—and therefore
    insufficient monitoring and pedigree of the DMW—the variation in predicted performance can
    easily exceed a factor of 10 The operating characteristics particular temperature (lower
    operating temperature appears to be worse) and the number and type of thermal cycles are
    extremely important in any assessment of DMWs This reality emphasizes the need for local
    monitoring particularly a sufficient number and distribution of thermocouples to better
    minimize the uncertainty in the assumptions

    98421449842144
    81
    8
    REPAIR
    Approved alternative weld repair methods that deviate from the rules for new construction in
    ASME Section I and ASME B311 are addressed for 9Cr steels in the National Board
    Inspection Code Part 3 Repairs and Alterations Welding Supplement 8 and Welding Method 6
    [62 63] Testing is ongoing to evaluate a repair scenario where ODconnected damage is
    identified at the fusion line or HAZ in a DMW between Grade 91 steel and a nickelbase weld
    metal Figure 81 Evaluation of alternative weld repair procedures using nickelbase filler
    materials without PWHT remain ongoing as part of the project scope detailed in [64] Sufficient
    data are being developed to allow for incorporation of the repair of DMWs between 9Cr steels
    and austenitic stainless steels using partial or fullpenetration repairs and select nickelbase filler
    materials

    Figure 81
    Simulated weld repair made using EPRI P87 filler metal and in a P91 to stainless steel 316
    DMW (the original weld metal is ENiCrFe3) (This simulated repair is being tested in
    uniaxial feature crossweld creep testing in the aswelded condition)

    98421449842144
    91
    9
    CONCLUSION
    It is often the case that failures—DMW or otherwise—result from a convergence of contributing
    characteristics in the design fabrication and operation of these components This point is
    emphasized by a number of recent failures in a finishing superheat wingwall arrangement where
    DMWs were fabricated between T91 and stainless steel 347H as shown in Figure 91 In this
    example a large number of unpredictable failures resulted in several forced outages major
    disruptions to the local supply chain providing fuel to the unit reduced management confidence
    in the unit’s operation and challenging weld repairs With respect to the engineering
    considerations reviewed in this report there were several identified factors that contributed to
    these failures
     Design DMWs were placed at rigid highly constrained locations This is evidenced in
    Figure 91A where the DMWs were placed immediately after tubemembrane construction
    The close arrangement of the constructed wingwall panels often resulted in catastrophic
    failures that would damage dozens of adjacent tubes The arrangement of wingwall panels in
    CFBtype units creates additional difficulties in the repair due to the lack of sufficient
    spacing
     Fabrication The use of automated GTAW in the fabrication shop and lack of a transition
    piece necessitated a large counterbore at the DMW see Figure 91B
     Operation A large spatial temperature variation was identified through a review of
    operating data and placement of additional thermocouples This variation was neither
    predicted nor anticipated by the OEM prior to the design and fabrication of the unit Failures
    were observed in the tubes operating at temperatures as high as 625°C (1155°F)
    Furthermore failure of solid tie attachments in the 347H tubing upstream of the DMW due to
    stress relaxation cracking placed additional loading on the DMW
     Metallurgy Failures were not identified at the fusion line and were consistent with
    cavitation in the Grade 91 HAZ andor parent material see Figures 91C and 91D
    9842144

    Conclusion
    92

    Figure 91
    Example of recent failures in DMWs in CSEF steel T91 to austenitic stainless steel 347H in
    a wingwall panel of a natural circulation fluidized bed boiler [65]
    In Figure 91 the reference operating conditions are as follows 565°C (1050°F) 198 barg
    (2875 psig) failures after ~15000 to 30000 hours of operation Oxide scale measurements
    suggested that failures were occurring in the highertemperature locations of the panel where the
    tube metal temperature was calculated to be 570–625°C (1060–1155°F)
    It may not be practical to eliminate all potential contributing factors to failures in DMWs through
    specification of best design and fabrication practices However it should be recognized that
    achieving better practice in lieu of best practice can greatly increase the longevity and
    performance of DMWs Furthermore with the lack of available tools for life management and
    general challenges in inspection there should be a greater emphasis placed on the proper design
    and specification of these connections than for similar weldments As the newest generation of
    DMWs in 9Cr steel tubing begin to exceed 100000 hours of operation in the supercritical
    power plants installed through the late 1990s and with DMWs being set to enter service for
    stateoftheart HRSGs it is clear that all future challenges in these components have not yet
    been realized

    9842144
    101
    10
    SUMMARY OF RECOMMENDATIONS
    The recommendations for inclusion into design and fabrication specifications contained
    throughout this document are listed below
     Restrictions on application DMWs should never be placed at these following types of
    welded connections
    – Thermowells steam sample probes or other similar types of components
    – Large attachment welds to piping
    – Stub to header or other small bore to thicksection components
    – Header end caps
    – Steam flow nozzles
     Weld joint geometry
    – For tubetotube welds such as in a reheater superheater or HRSG applications a wide
    cap should be used to extend the toe of the weld in the 9Cr steel beyond the fusion line
    and HAZ in the 9Cr steel
    – For thicksection welds nominally in piping components where the wall thickness is
    >127 mm (050 in) a step weld configuration combined with a wide cap is
    recommended if there exists sufficient access to complete the weld
     Weld process
    – DMWs should be fabricated with the following filler materials
    o GTAW or solid wire processes EPRI P87 ERNiCr3
    o SMAW process ENiCrFe2 ENiCrFe3 (until EPRI P87 fluxpool interaction is
    optimized it is not recommended for use)
    – A maximum interpass temperature of 315°C (600°F) particularly for automated welding
    processes should be imposed
    – For manual GTAW process the maximum diameter of the filler material should be
    limited to a maximum of 32mm (0125in) diameter and more preferably 24mm
    (0093in) diameter
     PWHT
    – DMWs should be given a PWHT in the lower end of the allowable range in the
    referenced Code of Construction As an example for ASME B&PV Section I the
    specified range would be 720°C (1325°F) ± 15°C (25°F)

    9842144

    Summary of Recommendations
    102
    – The maximum allowable PWHT temperature should never exceed 770°C (1420°F)
    – The PWHT time should be the minimum allowable time typically specified on the basis
    of the component thickness
     Transition pieces Where a thickness transition exists between the ferritic and austenitic
    stainless steel a transition piece should be utilized This transition piece should be higher
    alloy either matching in composition to the austenitic stainless steel or a nickelbase alloy
     Operating conditions
    – DMWs should never be placed in the vicinity of a soot blower downstream of an
    attemperator or other source of severe local thermal cycling which may include
    quenchingtype events
    – DMWs should never be placed at or in the immediate vicinity of a highrestraint location
    such as a roof or attachment weld in tubing
    – DMWs in thicksection or highly restrained applications should be placed in operating
    temperatures >550°C (1020°F) and more ideally ≥565°C (1050°F) For potential
    operating temperatures <565°C (1050°F) and regardless of the potential cycling
    conditions the use of a 9Cr steel DMW should be reviewed and substitution with a
    low alloy steel (such as 225Cr) should be considered
     Special considerations for tube attachments Quality assurance programs should address
    particular changes at tube attachments and include controls that
    – Limit the depth of penetration of the HAZ through the tube wall thickness
    – Achieve full penetration of the tubetoattachment weld

    9842144
    111
    11
    REFERENCES
    1 ASME Boiler & Pressure Vessel Code ASME B&PV I 2017 New York ASME
    2 ASME Code for Pressure Piping B311 American Society of Mechanical Engineers 2016
    New York ASME
    3 R D Thomas Jr and K W Ostrom Dilution of Austenitic Welds by Mild Steels and Low
    Alloys Welding Journal April 1941 pp 186s to 189s
    4 W Spraragen and D Rosenthal Welding of Dissimilar Metals Welding Journal February
    1945 65s to 85s
    5 H Thielsch StainlessSteel Weld Deposits on Mild and Alloy Steels Welding Journal
    January 1952 pp 37s to 64s
    6 J J B Rutherford Welding Stainless Steel to Carbon or LowAlloy Steel Welding
    Journal January 1959 pp 19s to 26s
    7 G E Lien Experience with Stainless Steel in Utility Power Plants Advances in the
    Technology of Stainless Steel and Related Alloys ASTM Special Technical Publication 369
    1965 pp 136 to 141
    8 G M Slaughter and T R Housley The Welding of Ferritic Steels to Austenitic Stainless
    Steels Welding Journal October 1964 454s to 460s
    9 L Schaeffler Selection of Austenitic Electrodes for Welding Dissimilar Metals Welding
    Journal October 1947 pp 601s to 620s
    10 O R Carpenter N C Jessen J L Oberg and R D Wylie Some Considerations in the
    Joining of Dissimilar Metals for HighTemperature HighPressure Service Proceedings of
    the ASTM Volume 50 1950 pp 809 to 857
    11 H Weisburg Cyclic Heating Tests of Main Steam Piping Joints between Ferritic and
    Austenitic Steels – Sewaren Generating Station Transactions ASME 71 (8) 1949 pp 643
    to 664
    12 DissimilarWeld Failure Analysis and Development Program Volume 1 Executive
    Summary EPRI Palo Alto CA 1985 CS4252V1
    13 D I Roberts R H Ryder B E Thurgood and R Viswanathan Improved Dissimilar
    Welds for Superheater and Reheater Tubes International Conference on Advances in
    Material Technology for Fossil Power Plants Sept 13 1987 Chicago IL
    14 Development of a New Nickel Filler for Dissimilar Metal Welds and Repair EPRI Palo
    Alto CA 2009 1018991
    15 Repair Methods for Dissimilar Metal Welds Development Weldability and Properties of
    EPRI P87 Solid Wire Filler Metal EPRI Palo Alto CA 2011 1019786
    9842144

    References
    112
    16 Cracking in ThickSection Dissimilar Metal Welds – Case Studies EPRI Palo Alto CA
    2015 3002006759
    17 DissimilarWeld Failure Analysis and Development Program Volume 8 Design and
    Procedure Guide for Improved Welds EPRI Palo Alto CA 1989 CS4252V8
    18 Factors Affecting Performance of Dissimilar Metal Welds Creep Performance of Screening
    Dissimilar Metal Welds Between Grade 91 Steel and Stainless Steel 347H EPRI Palo Alto
    CA 2016 3002007216
    19 Service Experience with Creep Strength Enhanced Ferritic Steels in Power Plants in the
    AsiaPacific Region EPRI Palo Alto CA 2015 3002005089
    20 ASME Boiler & Pressure Vessel Code ASME B&PV IIA 2017 New York ASME
    21 ASME Boiler & Pressure Vessel Code Code Cases Boilers and Pressure Vessels 2017 New
    York ASME Section I Code Case 23282 Austenitic Stainless Steel Tubes SA213SA
    213M UNS S30432 18Cr9Ni3CuCbN Approved September 18 2010
    22 ASME Boiler & Pressure Vessel Code Code Cases Boilers and Pressure Vessels 2017 New
    York ASME Section I Code Case 2864 9Cr1MoV Material Approved September 21
    2016
    23 ASME Boiler & Pressure Vessel Code Code Cases Boilers and Pressure Vessels 2017 New
    York ASME Section I Code Case 2839 9Cr3W3CoNdB Material Approved
    24 Factors Affecting Performance of Dissimilar Metal Welds Mechanical Analysis of
    Dissimilar Metal Welds Microstructural Characterization and Modeling of Dissimilar Metal
    Weld Failures Involving Grade 91 Steel EPRI Palo Alto CA 2016 3002007222
    25 ASME Boiler & Pressure Vessel Code ASME B&PV IIC 2017 New York ASME
    26 ASME Boiler & Pressure Vessel Code Code Cases Boilers and Pressure Vessels 2017 New
    York ASME Section IX Code Case 2733 FNumber Grouping for NiFeCr Classification
    UNS N08087 Welding Filler Metal Approved April 1 2014
    27 ASME Boiler & Pressure Vessel Code Code Cases Boilers and Pressure Vessels 2017 New
    York ASME Section IX Code Case 2734 FNumber Grouping for NiFeCr Classification
    UNS N08087 Welding Electrode Approved April 1 2014
    28 ASME Boiler & Pressure Vessel Code ASME B&PV I 2015 New York ASME Figure
    PG421 Welding End Transitions Maximum Envelope pp 38
    29 Life Management of Creep Strength Enhanced Grade 91 Steel Damage Calculator EPRI
    Palo Alto CA 2013 3002000082
    30 WellEngineered Weld Repair of Grade 91 Steel Welding Procedure Qualification to ASME
    Boiler and Pressure Vessel Code Section IX Part II EPRI Palo Alto CA 2016
    3002007233
    31 W F Newell Jr Welding and PWHT of P91 Steels Proceedings of the IPEIA
    Conference Banff Alberta 2011
    32 Guidelines and Specifications for HighReliability Fossil Power Plants 2nd Edition Best
    Practice Guideline for Manufacturing and Construction of Grade 91 Steel Components
    EPRI Palo Alto CA 2015 3002006390
    9842144

    References
    113
    33 J Henry and I J Perrin Dissimilar Metal Welds Involving Grade 91 Why Don’t They
    Last presented at the EPRI International Life Assessment Conference—Fossil Power 2012
    Hilton Head NC EPRI
    34 Babcock and Wilcox Memo authored by Mike Gold 1991 Heading Stainless Steel SH &
    RH Attachment Welds
    35 ASME Boiler & Pressure Vessel Code ASME B&PV I 2015 New York ASME Figure
    PW162 Some Acceptable Forms of Welds for Lugs Hangers and Brackets on Shells
    Drums and Headers p 102
    36 Detroit Edison Today Vol 14 No 4 Detroit MI February 1986
    37 A Olszewski Steam Sampling Probe DMW Failures on P91 Piping Presented at the EPRI
    Boiler Reliability and Interest Group Meeting December 2015
    38 Service Experience with Grade 91 Components EPRI Palo Alto CA 2009 1018151
    39 A Fabricius and P S Jackson Premature Grade 91 Failures – Worldwide Plant Operational
    Experiences Engineering Failure Analysis 66 (2016) pp 398 to 406
    40 C Matherne D DeRouen and W Baxter Catastrophic Failure of a Dissimilar Metal Weld
    in a High Pressure Steam Venturi Paper presented at the 2010 AIChE Spring National
    Meeting 2010 San Antonio TX AIChE
    41 J A Siefert J D Parker and R C Thomson Linking Performance of Parent Grade 91
    Steel to the Crossweld Creep Performance using Feature Type Tests Proceedings from the
    Eighth International Conference on Advances in Materials Technology for Fossil Power
    Plants ASM International 2016 pp 531 to 544
    42 WellEngineered Weld Repair of Grade 91 Steel Performance of Repair Welds
    Manufactured in an ExService Grade 91 Steel Header EPRI Palo Alto CA 2016
    3002004479
    43 J A Siefert J D Parker R C Thomson Factors Contributing to Heat Affected Zone
    Damage in Grade 91 Steel Feature Type Crossweld Tests Proceedings to the 4th
    International ECCC Conference on Creep and Fracture September 2017
    44 The Influence of Steel Making and Processing Variables on the Microstructure and
    Properties of CreepStrengthEnhanced Ferritic (CSEF) Steel Grade 91 EPRI Palo Alto
    CA 2014 3002004370
    45 Life Management of 9Cr Steels Evaluation of Metallurgical Risk Factors in Grade 91
    Steel Parent Metal EPRI Palo Alto CA 2017 3002009678
    46 A Dhooge Survey on Reheat Cracking in Austenitic Stainless Steels and Ni base Alloys
    Welding in the World 41 1998 pp 206 to 219
    47 N H Heo et al Dependence of Elevated Temperature Intergranular Cracking on Grain Size
    and Bulk Sulfur Content in TP347H Austenitic Stainless Steels ISIJ International 56 (6)
    2016 pp 1091 to 1096
    48 T Sourmail Precipitation in Creep Resistant Austenitic Stainless Steels Materials Science
    and Technology 17 (1) 2001 pp 1 to 14
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    49 K Laha et al A Comparison of Creep Rupture Strength of FerriticAustenitic Dissimilar
    Weld Joints of Different Grades of CrMo Ferritic Steels Metallurgical and Materials
    Transactions A 2011 Vol 43 No pp 1174–1186
    50 M Yamazaki et al Creep Rupture Properties and Fracture Type of 9Cr1MoVNb18cr
    8ni Steel Dissimilar Joints Welding in the World 2011 Vol 55 No 12 pp 67–77
    51 T Fukahori et al Experimental Study on the Creep Failure Modes of Dissimilar Metal
    Welds Presented at the ASMEEPRI Dissimilar Metal Weld Workshop at ASME Pressure
    Vessel and Piping Conference July 2017
    52 WellEngineered Weld Repairs of Grade 91 Steel Results for NickelBase Type Filler
    Materials EPRI Palo Alto CA 2014 3002003837
    53 WellEngineered Weld Repair of Grade 91 Steel Results for ThroughThickness Repair
    Welds EPRI Palo Alto CA 2014 3002004476
    54 WellEngineered Weld Repair of Grade 91 Steel Results of T91 Weld Repair Using the
    Nickel Base EPRI P87 Filler Metal EPRI Palo Alto CA 2014 3002003363
    55 Factors Affecting Performance of Dissimilar Metal Welds Creep Performance of Screening
    Dissimilar Metal Welds between Grade 91 Steel and Stainless Steel 347H 2017 Update
    EPRI Palo Alto CA 2016 3002007219
    56 Factors Affecting Performance of Dissimilar Metal Welds Mechanical Analysis of
    Dissimilar Metal Welds Part II Detailed Assessment to Support Best EPRI Palo Alto CA
    2017 3002007220
    57 E F Nippes The Weld Heat Affected Zone Welding Journal 38 (1) 1959 pp 1s to 17s
    58 J C Lippold Welding Metallugry and Weldability Wiley and Sons Inc New Jersey 2015
    pp 11–12
    59 X Xu G West J A Siefert J D Parker and R C Thomson The Influence of Thermal
    Cycles on the Microstructure of Grade 92 Steel Approved for publication in Metallurgical
    and Materials Transactions A on July 29 2017 Manuscript ETP161680AR
    60 Boiler Condition Assessment Guideline EPRI Palo Alto CA 2010 1019628
    61 Metallurgical Guidebook for Fossil Power Plant Boilers EPRI Palo Alto CA 2008
    1014183
    62 National Board Inspection Code Part 3 Repairs and Alterations United States of America
    2017 Section 2536 Welding Method 6 pp 46 to 47
    63 National Board Inspection Code Part 3 Repairs and Alterations United States of America
    2017 Supplement 8 – Weld and Post Repair Inspection of Creep Strength Enhanced Ferritic
    Steel Pressure Equipment pp 221 to 227
    64 Application of WellEngineered Weld Repairs for Grade 91 and Other CreepStrength
    Enhanced Ferritic (CSEF) Steels EPRI Palo Alto CA 2017 3002004332
    65 F Timmons and B Shelton Replacement of DMW Tube Sections in Finishing Superheater
    at Dominion’s Virginia City Hybrid Energy Center Presented at the ASMEEPRI
    Dissimilar Metal Weld Workshop at ASME Pressure Vessel and Piping Conference July
    2017
    9842144

    References
    115

    66 D I Roberts R H Ryder B E Thurgood and R Viswanathan Improved Dissimilar
    Welds for Superheater and Reheater Tubes International Conference on Advances in
    Material Technology for Fossil Power Plants Sept 13 1987 Chicago IL

    98421449842144
    1

    9842144Electric Power Research Institute
    3420 Hillview Avenue Palo Alto California 943041338 • PO Box 10412 Palo Alto California 943030813 USA
    8003133774 • 6508552121 • askepri@epricom • wwwepricom
    The Electric Power Research Institute Inc (EPRI wwwepricom)
    conducts research and development relating to the generation delivery
    and use of electricity for the benefit of the public An independent
    nonprofit organization EPRI brings together its scientists and engineers
    as well as experts from academia and industry to help address
    challenges in electricity including reliability efficiency affordability
    health safety and the environment EPRI members represent 90 of the
    electric utility revenue in the United States with international participation
    in 35 countries EPRI’s principal offices and laboratories are located in
    Palo Alto Calif Charlotte NC Knoxville Tenn and Lenox Mass
    TogetherShaping the Future of Electricity
    © 2017 Electric Power Research Institute (EPRI) Inc All rights reserved Electric Power
    Research Institute EPRI and TOGETHERSHAPING THE FUTURE OF ELECTRICITY are
    registered service marks of the Electric Power Research Institute Inc
    Program
    Technology Innovation
    3002007221
    9842144

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