A thesis submitted to the. Graduate School. of the University of Cincinnati. in partial fulfillment of the. requirements for the degree of

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2 Effect of Nitrogen Concentration in Shielding Gas on Microstructure and Mechanical Properties of ATI 2003 Lean Duplex Stainless Steel Autogenous Plasma Arc Welding A thesis submitted to the Graduate School of the University of Cincinnati in partial fulfillment of the requirements for the degree of Master of Science in the Materials Science and Engineering Graduate Program of the College of Engineering and Applied Science 2011 by Benjamin A. Sprengard B.Sc., The Ohio State University 2008 Committee Chair: Dr. Vijay K. Vasudevan

3 Abstract The effect of nitrogen concentration in the shielding gas on microstructure and mechanical properties of autogenously plasma arc welded lean duplex Allegheny Ludlum ATI 2003 material was studied. Six single-pass plasma arc butt welds were produced for the evaluation. One weld was performed using 100% argon shielding and backing gas, while the other five utilized varying additions of nitrogen in the shielding and backing gas, from 1 to 5%, with argon as the primary gas. 100% Argon was used as the plasma orifice gas for all welds. The coupons were labeled 0 through 5, which correlated with the concentration of nitrogen. The heat input was consistent for each weld and was representative of a typical value for production plasma arc welding of lean duplex stainless steel. Mechanical testing and evaluation was completed in accordance with typical customer requirements and industry standards. Each weldment was tested mechanically and analyzed microstructurally to investigate any correlations with the nitrogen additions in the weld shielding gas. Mechanical tests consisted of transverse and longitudinal tensile testing per ASME 2010 Section II, Part A, SA-370, Charpy impact testing per ASME 2010 Section II, Part A, -40 C, and micro-hardness testing per ASTM E384 with a test force level of 500 grams-force (gf). The microstructural analysis included ferrite testing, utilizing both the Fischer Ferritescope and PAXit software, and optical microscopy with focus on austenite formation and morphology, precipitate formation, and grain size comparison. The mechanical tests from this study revealed that only specific coupons met the requirements provided in ASME Boiler and Pressure Vessel Code. The transverse tensile tests revealed that all the coupons met the minimum requirements for the tensile strength of 620 MPa (90 ksi). Conversely, the elongation for each of the coupons except coupon 5, with 5% nitrogen in the shielding gas, fell short of the 25% elongation minimum specified for the base material. ii

4 Charpy impact tests disqualified coupons 0 through 2, which were unacceptable due to lateral expansions less than the minimum of 0.38 mm (.015 inches) per ASME Section VIII. Vickers micro-hardness testing was found to be optimum and below the ATI 2003 base metal requirement of 293 BHN per ASME Section II, Part A, SA-240. During ferrite testing, both using the Fischer Ferritescope and the PAXit software, it was determined that coupons 2 through 5 had optimum ferrite values of between 40% and 60%, for each the fusion zone, HAZ, and base metal. Conversely, Coupons 0 and 1 had values that fell outside this range. Neither secondary austenite precipitation nor formations of detrimental second phases were detected in any of the weld microstructures. In conclusion, 5% nitrogen added to the shielding gas had a beneficial effect on autogenously plasma arc welded ATI 2003 Lean Duplex Stainless Steel and resulted in optimum mechanical properties and microstructure of the weldment. Alternatively, autogenous plasma arc welding with 100% argon as the backing, shielding, and orifice gas resulted in unacceptable mechanical properties and an unbalanced ferrite-austenite ratio. iii

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6 Acknowledgements First, I would like to express deep gratitude toward Enerfab for all their support and for providing the opportunity to pursue this advanced degree. Without the help of Enerfab and their support for higher education, this achievement may have never been fulfilled. I especially would like to thank Kevin Kurtz, Senior Welding Expert, for his invaluable advice and Jim St. John, Enerfab s Director of Quality Assurance, for his continual support of my education over the past three years. I would like to thank Professor Vijay Vasudevan, my academic advisor, for his support and encouragement throughout my graduate studies at the University of Cincinnati. I would also like to express my appreciation toward Professor Jainagesh Sekhar and Professor Rodney Roseman for taking the time to review this study and provide professional advice. I am extremely grateful for the companies that supported me throughout this study including, ATI Allegheny Ludlum, for the materials support and helpful advice, and F & M Mafco, for the weld shielding gas support, and TEAM Industrial for allowing me to perform the research in their laboratory. TEAM Industrial and their great Materials Testing Laboratory (TMTL) staff were extremely welcoming and cordial. I especially would like to thank Tom Mosier, Metallurgical Engineer, and Anthony Clark, Level III Technician, for their invaluable advice and guidance. v

7 TABLE OF CONTENTS LIST OF FIGURES..viii LIST OF TABLES..ix 1.0 INTRODUCTION A Review of Lean Duplex Stainless Steel ATI 2003 (S32003) Lean Duplex Stainless Steel with Molybdenum Role of Nitrogen Objective EXPERIMENTATION ATI 2003 Test Material Weld Process Shielding Gas Mechanical Testing Tensile Testing Charpy Impact Testing Hardness Testing Metallurgical Evaluation Ferrite Testing Microstructural Evaluation RESULTS AND DISCUSSION Mechanical Properties Tensile Testing Charpy Impact Testing Hardness Testing Metallurgical Analysis Ferrite Analysis Microstructural Analysis...32 vi

8 4.0 SUMMARY AND CONCLUSIONS REFERENCES 44 vii

9 LIST OF FIGURES Figure 1 DSS Pseudo Binary Phase Diagram [3]... 2 Figure 2 ATI 2003 Base Metal, UNS S32003, Heat #827414, 200x Magnification... 9 Figure 3 Plasma Arc Welding Diagram [21] Figure 4 Specimen Test Coupon Layout Figure 5 Transverse Tensile Strength Figure 6 Transverse % Elongation Figure 7 Longitudinal Tensile Strength Figure 8 Longitudinal % Elongation Figure 9 Charpy Impact Test - Average Absorbed Energy Figure 10 Charpy Impact Test - Average Lateral Expansion Figure 11 Average Micro-Hardness Values Figure 12 Optical Micrograph of Specimen 0, Fusion Zone, 200x Magnification Figure 13 Optical Micrograph of Specimen 0, Fusion Zone, 79% Ferrite, 200x Magnification.. 28 Figure 14 Optical Micrograph of Typical Base Metal, 200x Magnification Figure 15 Optical Micrograph of Typical Base Metal, 47% Ferrite, 200x Magnification Figure 16 Ferrite Analysis using PAXit Software Figure 17 Ferrite Analysis using Fischer Ferritescope Figure 18 Optical Micrograph of Specimen 0, Fusion Zone, 50x Magnification Figure 19 Optical Micrograph of Specimen 5, Fusion Zone, 50x Magnification Figure 20 Optical Micrograph of Specimen 0, Fusion Zone, 100x Magnification Figure 21 Optical Micrograph of Specimen 5, Fusion Zone, 100x Magnification Figure 22 Optical Micrograph of Specimen 0, HAZ/BM, 50x Magnification viii

10 Figure 23 Optical Micrograph of Specimen 5, BM/HAZ, 50x Magnification Figure 24 Optical Micrograph of Specimen 0, Weld Cap, 50x Magnification Figure 25 Optical Micrograph of Specimen 1, Weld Cap, 50x Magnification Figure 26 Optical Micrograph of Specimen 2, Weld Cap, 50x Magnification Figure 27 Optical Micrograph of Specimen 3, Weld Cap, 50x Magnification Figure 28 Optical Micrograph of Specimen 4, Weld Cap, 50x Magnification Figure 29 Optical Micrograph of Specimen 5, Weld Cap, 50x Magnification Figure 30 Optical Micrograph of Specimen 0, HAZ, 50x Magnification Figure 31 Optical Micrograph of Specimen 1, HAZ, 50x Magnification Figure 32 Optical Micrograph of Specimen 2, HAZ, 50x Magnification Figure 33 Optical Micrograph of Specimen 3, HAZ, 50x Magnification Figure 34 Optical Micrograph of Specimen 4, HAZ, 50x Magnification Figure 35 Optical Micrograph of Specimen 5, HAZ, 50x Magnification Figure 36 Heat-Affected Zone Thermal Cycle [3] LIST OF TABLES Table 1 Chemical Composition of ATI 2003 [17][25]... 4 Table 2 Chemical Composition and Mechanical Properties ATI 2003, Heat # , [21]... 8 Table 3 Welding Parameters Table 4 Shielding Gas Composition and Certification [23] Table 5 Transverse Tensile Test Data Table 6 Longitudinal Tensile Strength Data Table 7 Charpy Impact Test Data ix

11 Table 8 Average Charpy Impact Test Data Table 9 Average Micro-Hardness Values Table 10 Ferrite Analysis using PAXit Software Table 11 Ferrite Analysis using Fischer Ferritescope Table 12 Shielding Gas Composition - Weldment Material Property Relations x

12 1.0 INTRODUCTION 1.1 A Review of Lean Duplex Stainless Steel Duplex stainless steels (DSS) are based on the Fe-Cr-Ni-N alloy system. Through compositional formulation and thermo-mechanical processing, they are designed to provide a two phase microstructure of nominally equal proportions of austenite and ferrite. The unique properties of duplex stainless steels are heavily weighted on this 50/50 balance of austenite, face centered cubic (FCC), and ferrite, body centered cubic (BCC), crystal structure. Their high corrosion resistance and increased strength, as compared with austenitic stainless steels, make DSS the favored material [3][5]. Due to their superior strength, they also have higher hardness making them attractive where abrasion and corrosion are of concern. Duplex stainless steels solidify as ferrite and depend on a partial solid-state transformation to austenite to result in nominally 50% ferritic and 50% austenitic microstructure. This solid-state transformation can effectively be controlled by optimizing the composition, processing temperature, and cooling rate. Figure 1 displays the elevated temperature region of a pseudo binary phase diagram for duplex stainless steel compositions [3, 20]. The Cr eq /Ni eq ratio for composition is on the x-axis, while the temperature is located on the y-axis. The shaded region represents the range of commercial duplex stainless steel alloys. This diagram reveals that austenite can only nucleate and grow below the ferrite solvus, thus making DSS rely completely on the solid-state transformation to austenite. 1

13 Nitrogen Alloying in Shielding Gas ATI = Cr eq /Ni eq Nitrogen Loss During Welding Figure 1 DSS Pseudo Binary Phase Diagram [3] During welding it is the solid-state transformation that is dependent upon cooling rate and composition which ultimately determines the ferrite-austenite balance and austenite distribution in the weld metal [3]. The solidification and transformation sequence of the microstructure for duplex stainless steel is as follows: L L + F F F + A 2

14 Austenite promoting elements, such as nickel, nitrogen, and copper are added to DSS to promote the transformation to austenite. When temperature decreases below the ferrite solvus, austenite begins to nucleate and grow along the ferrite grain boundaries. Austenite typically covers the grain boundaries completely and Widmanstätan plates may also protrude intragranularly into the ferrite grains [3]. Lean Duplex Stainless Steels (LDSS) are a subset of the duplex family of stainless steels. They have had significant growth and development over the past ten years [2]. In general, LDSS have a lower alloy content which makes them a lower cost alternative to standard duplex stainless steels and austenitic stainless steels. In addition to the benefit of lower cost, they also have high strength and good corrosion properties. LDSS have been utilized as a substitution for Types 304L and 316L for increased strength, stress corrosion cracking (SCC) and general corrosion resistance, and price stability [2]. They ve also replaced coated carbon steel in structural applications and Super Duplex Stainless Steels in subsea umbilicals with the use of an external Zn coating [2]. 1.2 ATI 2003 (S32003) Lean Duplex Stainless Steel with Molybdenum ATI 2003 (UNS S32003) is a lean duplex stainless steel (LDSS) which is molybdenumenhanced and was developed by ATI Allegheny Ludlum in It was created to bridge the performance gap between Type 316L (UNS S31603) austenitic and 2205 (UNS S32205) duplex stainless steels [9,16,17]. Although ATI 2003 is less costly than 2205 due to its reduced alloy content, it still retains the higher strength of the duplex grades and has increased strength and corrosion resistance as compared with 316L austenitic stainless steel. Strength, phase balance, 3

15 weldability, and impact toughness are the key characteristics of 2205 material that the developers of ATI 2003 wanted to maintain [2]. The chemical composition of ATI 2003 is listed in Table 1[17] along with the required mechanical properties per ASME Section II, Part A, SA-240, Table 2 Mechanical Test Requirements [25]. Due to the reduced levels of Cr and Mo, as compared with traditional duplex stainless steels, this alloy is more resistant to the formation of detrimental phases such as sigma, which can significantly lower the impact strength especially at low temperatures [17]. When compared with austenitic stainless steels, such as Type 316L, ATI 2003 has better pitting, SCC, and general corrosion resistance. In addition, ATI 2003 has significantly higher strength [2]. Table 1 Chemical Composition of ATI 2003 [17][25] Chemical Composition of ATI 2003 Element Weight Percent Carbon max Manganese 2.00 max Phosphorus max Sulfur max Silicon 1.00 max Chromium Nickel Molybdenum Nitrogen Iron Balance Mechancial Properties Tensile Strength, min 620 MPa (90 ksi) Elongation, min 25% Hardness, max 293 Brinell (~305 HV) 4

16 The weldablility of ATI 2003 is similar to 2205 material and can be welded by most of the methods used for all stainless steels [16]. Autogenous welding will increase the ferrite present in the fusion zone (FZ) and heat-affected zone (HAZ), as compared to the base metal. The increase occurs because these zones are rapidly cooled from temperatures near the ferrite solvus, inhibiting austenite formation. Nitrogen additions in the shielding gas are recommended to preserve phase balance and strength in the as-welded condition, although the exact amount of nitrogen has not been fully developed for this ATI Previous work proposed by Rogne, Fostervoll, and Rorvik has determined that 5% nitrogen added to plasma arc weld (PAW) shielding gas raised the contents of austenite by 15 to 20% for type 2205 (UNS31803) material, a standard DSS alloy [7]. Consequently, the rise in austenite formation effectively controlled the ferrite-austenite ratio and resulted in a 53-60% austenite microstructure [7]. Commercially available overmatched filler metals are suggested, such as AWS ER2209. These filler metals are boosted with nickel in order to promote austenite formation during the rapid solidification associated with welding [3,17]. When no filler metal is used, as in the case with autogenous welding, heat treatment is recommended to optimize corrosion resistance and formability by reestablishing the austenite-ferrite phase balance. For welding performance qualifications, per ASME Code Case , ATI 2003 (S32003) shall be considered a P-No. 10H, group 1 material [17]. 1.3 Role of Nitrogen Modern duplex stainless steels are intentionally alloyed with nitrogen to improve strength and pitting corrosion resistance [3]. The nitrogen helps maintain an optimum austenite to ferrite ratio, especially during welding. Extensive use of nitrogen over-alloying filler metals has been 5

17 made to make up for the loss of nitrogen that occurs during the welding process. When performing autogenous welding, nitrogen or nickel enhanced filler metals are not present, thus higher ferrite volumes are found in the weldment. This loss of nitrogen must be replaced by some type of alloying method [6]. Nitrogen additions in the shielding gas for standard duplex stainless steel has be The austenite-ferrite phase balance in duplex stainless steel weldments is greatly dependent upon diffusion. Because nitrogen is a small interstitial atom with more rapid diffusion rates than the large substitutional atoms in DSS, it is the key element for lowering the effective quench temperature to obtain the optimum phase balance in the weld metal and HAZ [3]. The concept of the effective quench temperature has been proposed by Vitek and David [17]. It describes a temperature during cooling, at which diffusion can no longer keep pace with the equilibrium compositions of the two phases. At this time and temperature, the composition and phase balance are frozen into the alloy. At temperatures above this point and with decreasing temperature below the ferrite solvus, the volume fractions of austenite and ferrite are continually changing and diffusion is slowing until it reaches the effective quench temperature [3]. The effective quench temperature for Cr, Ni, and Mo is approximately the same as the ferrite solvus temperature where nitrogen, due to its high mobility and diffusivity at significantly lower temperatures than the solvus, allows for more austenite formation at lower temperatures. Therefore, nitrogen content is of key importance for phase balance under weld cooling conditions. In addition, nitrogen also helps to decrease the Cr eq /Ni eq ratio, and thus increases the ferrite solvus temperature. Consequently, this increases the solid-state phase transformation start temperature resulting in a higher fraction of austenite in the fusion and heat affected zones. 6

18 1.4 Objective A common issue in the duplex welding industry is the question of exactly what quantity of nitrogen must be added to the weld shielding gas, especially during autogenous welding, to result in optimum properties for duplex stainless steel weldments. There is the issue of having an excessive amount of nitrogen in which the resultant weld can have excessive nitride formation, austenite, or even porosity. Conversely there could be insufficient nitrogen resulting in a high amount of ferrite. ATI 2003 is a relatively new alloy, developed in 2003, and has much economic potential. It has good weldability, although without the use of filler metal its optimal phase balance could be altered. This experiment will help define a way of retaining the optimum phase balance during autogenous plasma arc welding through the use of nitrogen alloying in the weld shielding gas. This study will not only reveal the resultant phase balance, but it will also describe the microstructure and material properties through metallography and mechanical testing. The ultimate objective is to make a recommendation for the percent of nitrogen in the plasma arc weld shielding gas for autogenously welded ATI 2003 lean duplex stainless steel. EXPERIMENTATION 2.1 ATI 2003 Test Material The test material used for this study was provided by Allegheny Ludlum. The material was ATI 2003 (UNS S32003) and the specific heat and lot numbers were and , respectively. It was solution annealed at above 1010 C (1850 F) and water quenched [21]. The 7

19 plate was in compliance with ASME Section II-A, SA-240 and had a thickness of 6.35mm (0.25 inches). The particular heat composition and mechanical properties are listed in Table 2 [21]. All the properties met the requirements of SA-240 for UNS S The chromium and nickel equivalents are and 7.56, respectively, with a Cr eq /Ni eq ratio of The calculations are shown below with applicable equations from the Welding Research Council (WRC) [3]. This Cr eq /Ni eq value is plotted schematically on Figure 1. Cr eq = Cr + Mo Nb = [3] Ni eq = Ni + 35 C + 20 N Cu = 7.56 [3] Cr eq /Ni eq = 3.12 Table 2 Chemical Composition and Mechanical Properties ATI 2003, Heat # , [21] Chemical Composition and Mechanical Properties of ATI 2003 (Heat # ) Base Metal Element Weight Percent Carbon 0.02 Manganese 1.7 Phosphorus Sulfur Silicon 0.28 Chromium Nickel 3.58 Molybdenum 1.7 Nitrogen Iron Balance Mechancial Properties (Heat # ) Yield Strength 493 MPa (71.5 ksi) Tensile Strength MPa (102.0 ksi) Hardness 228 Brinell (234 HV) 8

20 Figure 2 displays a photomicrograph of the specific ATI 2003 LDSS base material utilized for the testing. This micrograph reveals the duplex microstructure of the material and the relative 50/50 balance of austenite and ferrite at 200x magnification. It was found, through the use of PAXit imaging software that the base material consisted of an average of 48% ferrite. An image of the base material and the corresponding image including the PAXit image software are found in Figure 14 and Figure 15, respectively. These photomicrographs were captured at 200x magnification. Figure 2 ATI 2003 Base Metal, UNS S32003, Heat #827414, 200x Magnification 2.2 Weld Process The plasma arc welding (PAW) process used in this experiment utilized the transferred arc mode. In this mode the arc transfers from the tungsten electrode to the workpiece. A 9

21 diagram of this process is displayed in Figure 3. The joint configuration consisted of a square butt with no bevel or gap and was welded in the 1G position (flat) from one side. Because the ATI 2003 material was 6.35mm (0.25 inches) thick, only one pass was necessary to complete the weldment. The travel speed, amperage, and voltage were cm/min (6.9 ipm), amps, and 25.1 volts, respectively. This equates to a heat input of 19.7 kj/cm (50.1 kj/inch). No preheat or post weld heat treatment (PWHT) was performed. This heat input value is typical for welding duplex stainless steel in the industry. A complete list of the welding parameters is displayed in Table 3. Figure 3 Plasma Arc Welding Diagram [21] 10

22 Table 3 Welding Parameters Plasma Arc Welding Parameters 6.35mm Gas Coupon # Orifice Shielding Backing 0 Ar-100% Ar-100% Ar-100% 1 Ar-100% Ar-99%, N-1% Ar-99%, N-1% 2 Ar-100% Ar-98%, N-2% Ar-98%, N-2% 3 Ar-100% Ar-97%, N-3% Ar-97%, N-3% 4 Ar-100% Ar-96%, N-4% Ar-96%, N-4% 5 Ar-100% Ar-95%, N-5% Ar-95%, N-5% Electrical Characteristics Current Polarity Amparage Voltage Heat Input Direct Electrode Pos Amps 25.1 Volts 19.7 kj/cm (50.1 kj/in) Welding Technique Travel Speed Tungsten Tungsten Size PAW Orifice Size Oscillation cm/min 2% Thoriated mm mm None 2.3 Shielding Gas Six different compositions of shielding gas were used in this research. One weld was performed using 100% argon shielding and backing gas, while the other five utilized varying additions of nitrogen in the shielding and backing gas, from 1 to 5%, with argon as the primary gas. The shielding gas compositions and certifications for each weld coupon are listed in Table 4. This table displays the elemental concentration, blend tolerance, certified concentration, and certification accuracy for each mixture. Argon was used for the orifice plasma gas for the welding of each coupon. These argon-nitrogen gas mixtures were analyzed and certified to typical science and industry standards [23]. 11

23 Table 4 Shielding Gas Composition and Certification [23] Weld Coupon Number Component Concentration Shielding Gas Composition and Certification Blend Tolerance (±) Certified Concentration Certification Accuracy 0 Argon % Nitrogen 1.00% 10% 1.00% ± 2% Argon BAL --- BAL --- Nitrogen 2.00% 10% 2.00% ± 2% Argon BAL --- BAL --- Nitrogen 3.00% 10% 3.00% ± 2% Argon BAL --- BAL --- Nitrogen 4.00% 10% 4.00% ± 2% Argon BAL --- BAL --- Nitrogen 5.00% 10% 5.00% ± 2% Argon BAL --- BAL Mechanical Testing The following sections describe the testing performed on each of six (6) ATI 2003 autogenously plasma arc welded plates with varying amounts of nitrogen in the shielding gas. Figure 4 displays the layout of the test specimens for each coupon. Specimens A thru C and D thru F were used for heat affected zone and weld metal Charpy impact testing, respectively. Specimens H, and I were used for microstructure analysis, ferrite evaluation, and hardness testing. Transverse tensile tests were performed on specimens J and K, while specimen O was used for longitudinal tensile testing. 12

24 A B C D E F G H I J L K M N O Figure 4 Specimen Test Coupon Layout Tensile Testing Tensile tests were performed on each coupon per ASME 2010 Section II, Part A, SA-370 and Section IX, QW-150. Two transverse tests, perpendicular to the weld, and one all weld metal longitudinal test were completed on each coupon for analysis. The transverse tests were completed to evaluate the location of the fracture and the tensile strength of the specimen. The weld metal tensile specimens were used to give a more exact value for the weld metal strength and to be used if the transverse tensile specimens ultimately fractured in the base metal. Percent elongation was also calculated to roughly compare the ductility of each coupon. 13

25 2.4.2 Charpy Impact Testing Charpy impact testing was performed on each coupon to determine a correlation between the nitrogen additions in the weld shielding gas and the fracture toughness of each weldment. This test is commonly used in industry for the qualification of welding procedure specifications. Because temperature greatly affects the toughness of body centered cubic crystal structures, as found in ferrite, the temperature was reduced for testing to exaggerate the variation in absorbed energy in the duplex specimens. Charpy impact testing was performed on each coupon per ASME 2010 Section II, Part A, -40 C. Due to the thickness, 6.35mm (0.25 inches), of the ATI 2003 plate material, sub-size specimens were utilized for the testing. The sub-size specimens were 10 by 5 millimeters in compliance with the code. Six impact test specimens were taken from each weld coupon. Three were notched with the center in the weld metal and the remaining three were notched with the center in the HAZ Hardness Testing Because hardness is a general indicator of a material s strength, wear resistance, and ductility, hardness testing was utilized to help provide more evidence to differentiate which ATI 2003 weld coupon would exhibit the most optimum properties with respect to the shielding gas. Micro-indentation hardness tests were utilized rather than macro-hardness to allow for evaluation of more specific locations on the weldments cross-section, such as base metal, heataffected zone, and weld metal. Hardness testing was executed using the Vickers indenter per ASTM E384 with a test force level of 500 grams-force (gf). Hardness traverses were plotted on each of the six weld 14

26 coupons. The traverses were performed on the weldments cross-section and continually passed through the base metal, heat affected zone, and weld metal. 2.5 Metallurgical Evaluation The following subsections describe the metallurgical evaluations performed on each of the welded ATI 2003 lean duplex stainless steel coupons. The base metal, heat affected zone, and weld metal were all included in the examination Ferrite Testing PAXit Image software was utilized to evaluate the phase balance of the ATI 2003 lean duplex stainless steel weldments. This was performed by quantifying the fractional area of ferrite in the microstructure. The analyzed locations were captured using light microscopy at 200x magnification using mounted samples electrolytically etched with a potassium hydroxide (KOH) solution. This etchant distinguished the austenitic structure from the primary ferrite and allowed the PAXit software to be applied and ultimately determine the percentage of ferritic and austenitic phases. Another analytical tool, the Fischer Ferritescope, was used to detect the amount of ferrite in the duplex material. This was conducted by testing five different areas and taking the average to ultimately derive a final value. The areas tested were on the weld cap, as this is what is most practical for field evaluation of weld metal ferrite in the industry Microstructural Evaluation The analysis of the lean duplex stainless steel was performed using light microscopy. The purpose of the analysis was to determine a relationship between the nitrogen addition in the 15

27 weld shielding gas and the weldment microstructure. Some of the characteristics of high importance were the detection of chromium nitride precipitation due to high amounts of nitrogen, detrimental intermetallic phases, such as sigma and Chi phase, amount of austenite formation and morphology, and grain size. RESULTS AND DISCUSSION The following sections describe and summarize the results determined from the experimentation. They will also help to formulate correlations between the additions of nitrogen in the shielding gas and the outcome of the mechanical testing and microstructure. 3.1 Mechanical Properties The results from the mechanical testing revealed various trends and characteristics of the ATI 2003 weldment. Test results and discussions are found in the subsections below for tensile, Charpy impact, and micro-hardness testing Tensile Testing The tensile tests performed were transverse and longitudinal to the welds. The transverse test values were quite similar, except for a slight increase in the strength for coupon five, which contained five percent nitrogen in the weld shielding gas. These values are listed in Table 5 and displayed in Figure 5. This increase indicates that the five percent addition of nitrogen acted as an effective solid solution strengthening agent for the weldment. The increase was attributed to the nitrogen diffusing into the interstitial sites and making the material harder to deform. The specimens with no nitrogen additions, and thus 100% argon for shielding, fractured in the base 16

28 metal. This occurrence was a result of the lack of ductility in the fusion zone, as compared with the balanced austenitic-ferritic duplex base metal. In addition to the loss of ductility caused by the excessive amount of ferrite in the weld metal, the presence of Cr-rich nitrides, Cr 2 N, may have also helped to decrease the weldment s ductility and toughness. Nitrides form in the fusion zone and HAZ, as a consequence of the weld s inherent rapid cooling conditions. This results in high ferrite content and nitride precipitation because of insufficient time for the nitrogen to partition to the austenite. This theory of lack of ductility due to nitrides is undetermined because no nitride formation was evident in the light micrographs. Transverse percent elongation reveals a trend of increasing elongation with percent nitrogen. This is a direct result of the increase in the austenitic phase found in the weldment. This trend is displayed in Figure 6. Upon further analysis the elongation value per ASME Section II, Part A, SA-240, Table 2 Mechanical Test Requirements lists a minimum elongation value 25%. Coupon-5 is the only acceptable elongation value that complies with the ASME code requirements for ATI 2003 base metal. 17

29 Table 5 Transverse Tensile Test Data Transverse Tensile Test Data Test # %N Width (mm) Thickness (mm) Area (mm) Ult. Load (N) UTS (MPa) Length i (mm) Length f (mm) % Elongation Location 0-J Base Metal 0-K Base Metal 1-J Weld Metal 1-K Weld Metal 2-J Weld Metal 2-K Weld Metal 3-J Weld Metal 3-K Weld Metal 4-J Weld Metal 4-K Weld Metal 5-J Weld Metal 5-K Weld Metal 18

30 Tensile Strength (MPa) Base Metal Fracture Transverse Tensile Strength Data 0-J 0-K 1-J 1-K 2-J 2-K 3-J 3-K 4-J 4-K 5-J 5-K Specimen # Figure 5 Transverse Tensile Strength Elongation (%) Transverse % Elongation 0-J 0-K 1-J 1-K 2-J 2-K 3-J 3-K 4-J 4-K 5-J 5-K Specimen # Figure 6 Transverse % Elongation 19

31 The longitudinal tensile behaviour is quite peculiar. The coupon which had 100% argon as shielding gas had the lowest tensile strength and the coupon with one percent nitrogen increased in strength by over 5000 psi. As the nitrogen additions increased, the ultimate tensile strength decreased until it reached the five percent coupon, which had the highest value of them all. These values are listed in Table 6 and displayed in Figure 7. This scattered data may be a result of other variables, such as weld width. Because the fusion zone and HAZ widths were not the same for each coupon, one tensile bar may have included more HAZ or fusion zone than another, resulting in possible faulty values. These scattered results are also evident in Figure 8 for the longitudinal elongation to fracture. Table 6 Longitudinal Tensile Strength Data Longitudinal Tensile Test Data Test # %N Width (mm) Thickness (mm) Area (mm) Ult. Load (N) UTS (MPa) Length i (mm) Length f (mm) % Elongation Location Weld Metal Weld Metal Weld Metal Weld Metal Weld Metal Weld Metal 20

32 700.0 Longitudinal Tensile Strength Data Tensile Strength (MPa) Specimen # Figure 7 Longitudinal Tensile Strength 35.0 Longitudinal % Elongation Elongation (%) Specimen # Figure 8 Longitudinal % Elongation 21

33 3.1.2 Charpy Impact Testing The Charpy impact test values for the weld HAZ were much higher than that of the fusion zone. For autogenous welding, such as in this study, nucleation occurs by atoms arranging from the liquid metal upon the substrate grains without changing their existing crystallographic orientation [1]. This means that the nuclei in the melt (fusion zone) will have the same orientation and lattice structure as the base metal grains located at the solid-liquid interface [26]. This scenario results in epitaxial nucleation and grain growth of the ferrite, which ultimately leads to course, columnar shaped ferrite grain morphologies in the fusion zone lacking toughness. An example of the columnar shaped morphology, captured from test coupon-5 is displayed in Figure 19. The HAZ toughness was higher as compared to the fusion zone because of the finer equiaxed grains. Figure 23 displays the equiaxed grains of the HAZ found in coupon-5. The overall Charpy impact absorbed energy and lateral expansion behaviour for the coupons was relatively similar, except for coupon-5. The five percent nitrogen in the weld shielding gas had a strong effect on the fusion and heat-affected zone toughness. This amount of nitrogen lowered the effective quench temperature, as discussed in section 1.3, and allowed enough time for the ferrite to austenite solid state phase transformation to take place resulting in an optimum duplex phase balance and higher toughness. The impact test values are listed in Table 7 and 8. Graphic illustrations of the average absorbed energy and lateral expansion are displayed in Figure 9 and Figure 10, respectively. Per ASME, Section VIII, UHA-51(a)(2), a minimum of 0.38 mm (0.015 inches) of lateral expansion opposite the notch is required for each of the three specimens tested for the FZ and HAZ. Per this requirement, coupons 0 thru 2 were unacceptable, while the remaining three were acceptable. 22

34 Table 7 Charpy Impact Test Data Charpy Impact Test Data Specimen # Location Size (mm) Temp ( C) Absorbed Energy (Joules) Lat. Exp. (mm) % Shear 0-D WM 10 x E WM 10 x F WM 10 x B HAZ 10 x C HAZ 10 x G HAZ 10 x D WM 10 x E WM 10 x F WM 10 x A HAZ 10 x B HAZ 10 x C HAZ 10 x D WM 10 x E WM 10 x F WM 10 x B HAZ 10 x C HAZ 10 x G HAZ 10 x D WM 10 x E WM 10 x F WM 10 x B HAZ 10 x C HAZ 10 x G HAZ 10 x D WM 10 x E WM 10 x F WM 10 x B HAZ 10 x C HAZ 10 x G HAZ 10 x D WM 10 x E WM 10 x F WM 10 x A HAZ 10 x B HAZ 10 x C HAZ 10 x

35 Table 8 Average Charpy Impact Test Data Coupon # Location Average Charpy Impact Test Data % Nitrogen Absorbed Energy (Joules) Mils Lat. Exp. (mm) WM HAZ WM HAZ WM HAZ WM HAZ WM HAZ WM HAZ Average Absorbed Energy - Impact Testing Absorbed Energy (Joules) Heat Affected Zone Weld Metal % Nitrogen in Shielding Gas Figure 9 Charpy Impact Test - Average Absorbed Energy 24

36 Average Lateral Expansion - Impact Testing Lateral Expansion (mm) Heat Affected Zone Weld Metal % Nitrogen in Shielding Gas Figure 10 Charpy Impact Test - Average Lateral Expansion Hardness Testing Micro-hardness test results were all acceptable per ASME Section II, Part A, SA-240, Table 2 Mechanical Test Requirements. The results for each coupon were relatively similar, with a range of the average values for the fusion zone, heat-affected zone, and base metal of 6, 6, and 4 units, respectively. This small variance in values could be attributed to the nitrogen s strengthening effects counteracting the toughness of the austenitic phase and, consequently resulting in optimum hardness. The average Vickers mirohardness values are listed in Table 9 and the results are graphically displayed in Figure

37 Table 9 Average Micro-Hardness Values Average Micro-Hardness Values Test Number HAZ Weld Metal Base Metal Vicker's Microhardness HV Fusion Zone HAZ Base Metal % Nitrogen in Shielding Gas Figure 11 Average Micro-Hardness Values 26

38 3.2 Metallurgical Analysis The following subsections contain analysis of metallurgical effects on the ATI 2003 microstructure as a result of the varying nitrogen content in the plasma arc weld shielding gas. Ferrite analysis tools and light microscopy were utilized Ferrite Analysis As described in section 2.5.1, two different methods were used for the detection of primary ferrite in the microstructure. PAXit image software was utilized first to evaluate the coupons. Figure 12 and 13 demonstrate how the software was applied to a typical fusion zone micrograph, whereas Figure 14 and 15 display the software generated image for a typical base metal. The calculated ferrite values are listed in Table 10 and graphically displayed in Figure 16. By briefly reviewing the data, the characteristic behavior of decreasing ferrite content is easily revealed. This decrease in ferrite is attributed to the increase in nitrogen additions in the shielding gas which, in turn, were able to provide sufficient alloying of the fusion zone. The nitrogen alloying helped to decrease the effective quench temperature as discussed in section 1.3 and increase the ferrite solvus. The increase in ferrite solvus was a result of a compositional shift which decreased the Cr eq /Ni eq ratio. In combination, these two phenomena obtained the optimum austenitic-ferritic phase balance in the coarse grained heat affected zones (CGHAZ) and fusion zones for the coupons which contained 2 to 5% nitrogen in the shielding gas. 27

39 Figure 12 Optical Micrograph of Specimen 0, Fusion Zone, 200x Magnification Figure 13 Optical Micrograph of Specimen 0, Fusion Zone, 79% Ferrite, 200x Magnification 28

40 Figure 14 Optical Micrograph of Typical Base Metal, 200x Magnification Figure 15 Optical Micrograph of Typical Base Metal, 47% Ferrite, 200x Magnification 29

41 Table 10 Ferrite Analysis using PAXit Software % Ferrite - PAXit Software Test Number HAZ Fusion Zone Base Metal 0 76% 79% 49% 1 72% 60% 47% 2 55% 58% 47% 3 51% 59% 48% 4 50% 46% 50% 5 40% 47% 47% 90% % Ferrite - PAXit Software % Ferrite 80% 70% 60% 50% 40% 30% 20% 10% HAZ Weld Metal Base Metal 0% % Nitrogen in Shielding Gas Figure 16 Ferrite Analysis using PAXit Software 30

42 The second method, which utilized magnetic determination, required the use of a Fischer Ferritescope analyzing instrument. The results from this method are listed in Table 11, while Figure 17 displays the data graphically. Through the use of this technique, the results from the image software analysis were confirmed. Table 11 Ferrite Analysis using Fischer Ferritescope % Ferrite - Fischer Ferritescope Test Number Location 1 Location 2 Location 3 Location 4 Location 5 Average % 89.5% 71.4% 65.3% 76.8% 77.4% % 61.3% 60.8% 62.9% 57.1% 60.8% % 62.9% 66.8% 52.6% 56.1% 59.7% % 55.3% 58.3% 53.5% 52.2% 56.7% % 52.3% 48.7% 58.7% 55.2% 55.4% % 54.7% 51.8% 46.5% 41.8% 49.6% 90.0% % Ferrite - Fischer Ferritescope 80.0% 70.0% 60.0% % Ferrite 50.0% 40.0% 30.0% 20.0% 10.0% 0.0% % Nitrogen in Shielding Gas Figure 17 Ferrite Analysis using Fischer Ferritescope 31

43 Although this technique is very effective in the field, it can be difficult to accurately identify the location of the acquired data, especially between the base metal, HAZ and the fusion zone. For example, although the Fischer Ferritescope produced acceptable results for coupon 1, the PAXit software, produced high ferrite results in the HAZ. This concern must be considered when developing welding procedure specifications. In conclusion, a 5% increase in concentration of nitrogen in the PAW shielding gas resulted in a 25 to 35% increase in austenite content in the ATI 2003 weldment. In comparison with the 15 to 20% austenite increase effect on type 2205 DSS welding discussed in section 1.2, there is a larger degree of efficiency of austenite transformation for ATI 2003 material Microstructural Analysis The resultant microstructures of each of the six weldments had varying amounts of austenite formation. There was no secondary austenite precipitation detected in the microstructure, which is expected given that this study focused on single pass welding and no reheating of the original weld bead occurred. In addition, detrimental second phases, such as sigma and chi phase, were absent from the microstructure. This is most likely attributed to the reduced levels of Cr and Mo, as discussed in section 1.2. Chromium nitride formation was also undetectable for all compositions of shielding gas. All of the above microstructural analysis was performed with optical light microscopy. The principal distinction between each coupon was the increase in austenite formation. This increase is consistent with the pseudo-binary phase diagram in Figure 1 because the increased nitrogen shifted the composition to the left. When welding was performed on the duplex material, nitrogen diffused out of the metal and rapid solidification took place. Cooling 32

44 rate and composition are the two variables which are ultimately responsible for the final austenite to ferrite ratio. Because the cooling rate was unchanged throughout this study, the increase in austenite formation was attributed to the increase in nitrogen content in the microstructure by means of weld shielding gas alloying. This compositional shift from the original Cr eq /Ni eq to a higher value after welding and then ultimately to a lower value after nitrogen alloying is displayed schematically in Figure 1. This final shift increased the ferrite solvus, and consequently increased the phase transformation start temperature resulting in a higher fraction of austenite in the fusion zone. The higher ferrite solvus temperature not only increased the time for austenite formation, but it also allowed the diffusion controlled nucleation and growth to occur more rapidly because of the higher temperature and ease for diffusion of Cr, Ni, and N. The nitrogen also decreased the effective quench temperature allowing for more austenite formation at lower temperatures. In coupon 0, austenite formation occurred mainly along the ferrite grain boundaries, with some Widmanstäton sideplates and few small platelets within the ferrite grains as compared with coupon 5. Coupon 5 also revealed allotrimorphs at the prior ferrite grain boundaries, but to a greater degree and with extensive Widmanstäton sideplates and coarse intragranular austenitic precipitates. This comparison for the fusion zones is revealed in Figure 18, 19, 20, and 21. As discussed in section 3.1.2, epitaxial nucleation and grain growth of the ferrite resulting in course columnar shaped grain morphologies are evident in the figures stated above. These morphologies were characteristic of all six compositions of shielding gas. 33

45 Figure 18 Optical Micrograph of Specimen 0, Fusion Zone, 50x Magnification Single Grain Figure 19 Optical Micrograph of Specimen 5, Fusion Zone, 50x Magnification 34

46 Figure 20 Optical Micrograph of Specimen 0, Fusion Zone, 100x Magnification Figure 21 Optical Micrograph of Specimen 5, Fusion Zone, 100x Magnification 35

47 Figure 22 and 23 are micrographs containing the HAZ and a small fraction of the ATI 2003 lean duplex base material for coupons 0 and 5, correspondingly. Figure 36 depicts three separate time regions experienced by the CGHAZ adjacent to the fusion line [3]. In Region I, the austenitic phase in the lean duplex base material begins to transform to ferrite upon heating to the ferrite solvus temperature. Region II is representative of the time above the ferrite solvus in which ferrite grain coarsening occurs due to a lack of austenite in the mictrostructure. It is this time interval that is responsible for loss of toughness and ductility in the CGHAZ. In time region III, below the ferrite solvus, austenite begins to nucleate and grow. The amount of austenite formation is dependent upon the cooling rate and time in this region. When comparing these images, it is evident that the grain size in coupon 0 is coarser than in coupon 5, resulting in lower toughness as displayed in Figure 9. This difference in grain size is attributed to the nitrogen addition effectively increasing the ferrite solvus temperature near the weld fusion line. This increase in the ferrite solvus helps limit ferrite grain coarsening by restricting the time region above the solvus. During this time region, grain coarsening occurs because there is no austenite to inhibit growth, resulting in a dramatic effect on ferrite grains. Consequently, the coupons with less nitrogen composition in the shielding gas, and a lower ferrite solvus temperature, produced larger grains in the CGHAZ with lower toughness. In addition to grain size, the austenite formation correlates with the trend in the fusion zone. The steady increase in fusion zone and HAZ austenite is displayed sequentially in Figures 24 through 29 and Figures 30 through 35, respectively. 36

48 Base Metal Single Grain HAZ Figure 22 Optical Micrograph of Specimen 0, HAZ/BM, 50x Magnification Base Metal Single Grain HAZ Figure 23 Optical Micrograph of Specimen 5, BM/HAZ, 50x Magnification 37

49 Figure 24 Optical Micrograph of Specimen 0, Weld Cap, 50x Magnification Figure 27 Optical Micrograph of Specimen 3, Weld Cap, 50x Magnification Figure 25 Optical Micrograph of Specimen 1, Weld Cap, 50x Magnification Figure 28 Optical Micrograph of Specimen 4, Weld Cap, 50x Magnification Figure 26 Optical Micrograph of Specimen 2, Weld Cap, 50x Magnification Figure 29 Optical Micrograph of Specimen 5, Weld Cap, 50x Magnification 38

50 Figure 30 Optical Micrograph of Specimen 0, HAZ, 50x Magnification Figure 33 Optical Micrograph of Specimen 3, HAZ, 50x Magnification Figure 31 Optical Micrograph of Specimen 1, HAZ, 50x Magnification Figure 34 Optical Micrograph of Specimen 4, HAZ, 50x Magnification Figure 32 Optical Micrograph of Specimen 2, HAZ, 50x Magnification Figure 35 Optical Micrograph of Specimen 5, HAZ, 50x Magnification 39

51 Figure 36 Heat-Affected Zone Thermal Cycle [3] SUMMARY AND CONCLUSIONS ATI 2003 can be welded with acceptable and optimum results autogenously with PAW provided the correct weld parameters and shielding gas composition are utilized. It should be noted that this study is only applicable for ATI 2003 material and that other lean duplex stainless steel materials may react differently to nitrogen additions in weld shielding gas. In addition, this evaluation utilized acceptance criteria per the requirements of ASME Boiler and Pressure Vessel Code which may differ from other industrial codes and specifications. 40