Effect of Welding Parameters on Tungsten Carbide - Metal Matrix Composites Produced by GMAW

By: Leonard Choii, Tonya Wolfeii, Matthew Yarmuchiii, Adrian Gerlichiv

Abstract

Wear of materials in the oil sands industry includes severe low stress sliding abrasion and various types of slurry erosion. Tungsten carbide metal matrix composite materials are one of the various wear-resistant materials used in mining applications to combat these problems. The application of these coatings is used to extend the service life of production-critical components, reduce maintenance costs and avoid production outages. Tungsten carbide composite overlays are most commonly applied to the component surface by plasma transferred arc welding (PTAW) with the use of powder consumables. Although PTAW is the most common and efficient method for shop-production, this process cannot be used in all situations. Due to the required substrate geometry and immobility of the equipment or component, it is impractical to use PTAW for field welding applications.

The present work examines the performance of tungsten carbide wire consumables during gas metal arc welding (GMAW). Although the GMAW process provides the flexibility and mobility required for field repair welding applications, the wire consumables are currently not widely used in industry due to inferior wear performance compared to PTAW overlays. The influence of welding parameters on the resulting microstructure and wear performance are evaluated. Specifically, the properties will be investigated, including macro and microstructures of the carbides, percent dilution, percent carbide area, mean free path and scanning electron microscopy with Auger Electron Spectroscopy. Strategies for optimizing welding parameters to improve the wear properties of wire-based overlays will be discussed.

i MSc Graduate Student, University of Alberta and Research Engineer, Alberta Innovates - Technology Futures (formerly Alberta Research Council). ii Research Engineer, Alberta Innovates - Technology Futures (formerly Alberta Research Council) iii Team Leader, Welding Engineering, Alberta Innovates - Technology Futures (formerly Alberta Research Council) iv Assistant Professor, Chemical and Materials Engineering, University of Alberta

Effect of Welding Parameters on Tungsten Carbide - Metal Matrix Composites Produced by GMAW

By: Leonard Choi, Tonya Wolfe, Matthew Yarmuch, Adrian Gerlich

Background

Wear in the Oil Sands Industry The oil sands deposits in Northern Alberta are the second largest in the world,

behind only Saudi Arabia, and contain over 175 billion barrels of oil (DOE, 2010).

Current production from the oil sands is 55% of western Canada’s total crude oil

production, and is expected to grow from approximately 1.3 million barrels per day (bpd)

in 2009 to 2.2 million bpd in 2015 and 3-3.5 million bpd by 2020-2025 (DOE, 2010,

CAAP, 2010). Oil sands are processed either by in-situ operations (e.g., thermal or other

assisted recovery methods) or open pit mining operations (e.g. bucket and truck

methods). A general overview of the open pit mining process and hydrotransport

systems are illustrated in Figure 1. Oil sands mining operations present very unique

challenges; for example, the equipment and machinery must be tolerant enough to

withstand the extreme conditions present when handling two tonnes of oil sands ore for

every barrel of refined oil produced (Harper et al., 2002). The sands in oil sands are

mainly composed of quartz sand, silt, clay, water and bitumen (Flores et al, 2009).

Quartz particles comprise 80% to 95% of the total solids in oil sands in varying sizes and

forms. These hard particles are the main causes of wear attack on oil sands equipment,

primarily by the low stress abrasion, high stress abrasion, gouging abrasion, and impact

wear mechanisms, such as in Steps 2 through 5 in Figure 1. The introduction of water,

during hydrotransport operations, has lead to significant maintenance and reliability

issues due to the synergic erosion-corrosion damage mechanism (Step 6 in Figure 1).

For example, this is seen by the increased wear rates of separating screens, inclined

settling plates, slurry pumps, and tailing pipelines. (Flores et al, 2009; Llewellyn, 1997).

Wear attack of equipment is responsible for higher production costs resulting in

hundreds of millions of dollars being spent yearly on maintenance activities, not to

mention the loss of production (Llewellyn,1995). For example, the annual budget for

repair and maintenance of equipment at a major oil sands producer was in excess of

$450 million in 2003 (Anderson et al, 2004). A major portion of this budget is due to wear

damage to machinery and equipment. The lost time in production due to maintenance

and unplanned shut-downs also amounts to hundreds of millions of dollars. Production

goals and investment returns would be difficult to achieve without protective overlays

and coatings for key production equipment.

2 3 61 4 5

Hydrotransport Piping to Extraction Plant

Figure 1 - General Overview of the Oil Sands Open Pit Mining Process

Wear-resistant Materials in Oil Sands Mining Operations The equipment being used throughout the oil sands extraction process are

currently made of high-grade steels, cast irons with chromium carbide (CrC) overlays,

hardened carbon steels and chromium white irons (Flores et al, 2009). While these

materials have the characteristics to resist wear, weld overlay materials have been

developed to extend the service life of production-critical components in oil sands

operation. For the most demanding applications, the weld overlay materials of choice are

metal matrix composites (MMCs) which are typically deposited by plasma transferred arc

welding (PTAW) (Anderson et al, 2004).

A metal matrix composite (MMC) is composed of hard (reinforcing) particles

fused together in a (ductile) metal matrix alloy. In the case of a nickel-based tungsten

carbide MMC, which is a common material used in oil sands applications, the tungsten

carbide particles act as the reinforcing phase bonded in the ductile NiCrBSi or NiBSi

matrix alloy. These composites combine the wear resistance of the tungsten carbide

(2200 – 3100 HV) and the ductility of nickel to produce a tough material (Anderson et al,

2003). PTAW is the most common weld overlay method for depositing MMC overlays.

PTAW involves an arc being established between a non-consumable tungsten electrode

and the work piece, typically with argon gas shielding and powder feeding through the

welding torch. This process can produce relatively thick weld overlays, typically 4-6 mm

in a single pass, and with the appropriate selection of welding parameters it is possible

to achieve low penetration and low dilution levels (Yarmuch et al., 2008). Due to the

nature of the equipment, PTAW is usually automated which produces consistent MMC

overlays and achieves higher productivity than manual welding (Deuis et al, 1998).

Laboratory tests and field applications have shown that these Ni/WC MMC overlays,

applied by PTAW, can significantly improve the performance of oil sands mining

equipment and extend their service life by up to 500% depending on the specific

application (Flores et al, 2009, Harper et al, 2002).

Although PTAW can produce high-quality MMC overlays reliably through

automation, this process is limited to shop production environments. The relatively

expensive equipment and infrastructure necessary for PTAW renders it impractical for

field welding applications. Overlay deposition and proper powder feeding to the torch is

dependant on gravity; hence, some complex substrate geometries are not suitable for

PTAW without significant fixturing and manipulation of the work piece. Gas metal arc

welding (GMAW) is a very versatile process that may overcome these obstacles, as it

may be more applicable for complicated part geometries and is readily adaptable to field

welding (repair) applications.

This paper will focus on NiCrBSi/Wc-W2C MMC overlays applied by GMAW. The

consumable is in the form of a cored wire (similar to a flux-cored or metal-cored wire)

where the sheath is composed of nickel and the core material is the tungsten carbide

(hard phase) and additional matrix alloying elements. Angular tungsten carbide particles

can come in eutectic powders, which are made up of WC and W2C phases, or

macrocrystalline, which are homogeneously WC phase. Eutectic carbide particles are

used in this case. The present study focuses on the influence of GMAW welding

parameters on the weld overlay microstructures produced, with particular emphasis on

microstructural changes due to the thermal (heat input) cycle and metal transfer mode

during welding. Currently most cored wire manufacturer literature suggests that the user

should employ the “lowest practical” current and voltage in order to avoid degradation of

the primary tungsten carbides. It is postulated that greater degradation may occur during

GMAW, wherein the carbides are directly exposed to the welding arc, compared to

PTAW in which the powder is typically injected at the sides/rear of the arc. A thorough

understanding of the role of the GMAW process parameters, metal transfer mode, and

heat input on the resultant MMC microstructure is critical, as these are primary factors

influencing the performance (and service life) and the MMC in wear and corrosion oil

sands applications.

Experimental Approach The objective of this study was to investigate the effects of GMAW process

parameters on the quality of the MMC overlay macro- and microstructure. This was done

by changing the welding parameters to achieve different metal transfer modes: short-

circuiting, globular and spray. An additional three weld overlays were made using a

higher and more productive wire feed speed (WFS) setting, again achieving the different

metal transfer modes. The following describes the welding of the test specimens and the

evaluations protocols utilized to examine the overlay properties.

Welding of Testing Specimens For this study a 1.60 mm (1/16”) diameter Arctec Tungcore FCS cored wire was

investigated (supplied by Arctec Alloys Ltd) utilizing 98% argon - 2% oxygen gas, DC

electrode positive polarity, and 19 mm (3/4”) electrode stickout. For reference, the

manufactured recommended welding parameters are 18-20 V, 160-180 A, and 30-80

cm/min (12-32 in/min) travel speed. According to the product literature, the weld deposit

consists of fused tungsten carbide (comprising a eutectic of WC and W2C phases) in a

Ni-Cr-B-Si matrix. The reported hardness of the matrix is 560-620 HV. The carbides

have a reported hardness of 2340 HV with an approximate mesh size of +100-120 mesh

(120-150 m) (Arctec Alloys, 2010). Welds were produced using a Lincoln Electric

Power Wave 455M/STT® with constant voltage operation (Program 5), and the

instantaneous True Energy® value (the “active” real energy reported by the power

source) was recorded in kilojoules. Heat input calculations were based on the

instantaneous energy values, as specified in QW-409.1 of ASME Section IX (Melfi,

2010). The welding torch was attached to a Bug-O® modular drive system to maintain a

constant travel speed and electrode stick-out setting; thus ensuring consistent and

repeatable welds. The output welding parameters (voltage and current) were also

recorded using a FLUKE 345 Power Quality Clamp Meter to provide a measure of the

welding operation independent of the power source.

The test parameters for this study are shown in Table 1. To achieve the “low”

productivity condition, a WFS setting of 305 cm/min (120 ipm) was utilized; the voltage

was then set to achieve the different transfer modes. Lastly, the travel speeds were

adjusted to match the (nominal) heat input achieved using the manufacturer’s average

recommended parameters (note: Specimen B is within the manufacturer settings, as

described above). For the “high” productivity conditions, a WFS setting of 635 cm/min

(250 ipm) was employed. The conditions for Weld Sample D were established first, then

voltages and travel speeds for the Samples E & F were based on achieving the metal

transfer modes and maintaining a similar (nominal) heat input level as Sample D.

Table 1 - Welding Parameters of MMC Weld Overlays

Productivity Condition

WeldSample

WFS[cm/min (ipm)]

Voltage[V]

Travel Speed [cm/min]

A 305 (120) 13.0 27B 305 (120) 19.0 30LowC 305 (120) 28.0 47 D 635 (250) 15.5 30E 635 (250) 18.5 32HighF 635 (250) 31.0 49

For this study, all overlays were double pass overlays were produced. Many

demanding oil sands applications require multi-pass overlays as the thicker protective

layer extends component life. However, the heat input must be minimized to prevent

significant (detrimental) changes to the base metal HAZ and overlay macro and

microstructure.

Examination of Weld Overlay Macro and Microstructures Following welding, the weld overlay top surface (as welded) was examined

visually and using low-magnification stereo-microscopy. Metallographic cross-sections of

the welds were prepared using standard laboratory techniques in order to reveal the

macroscopic structural features of the MMC. The weld dimensions, percent dilution,

percent carbide area and mean free path ( ) were determined as outlined below:

Percent Dilution (%dil.) – is the ratio between the amount of base metal melted

and mixed with the molten consumable during welding. The %dil. can be

calculated using the cross-sectional area with the formula and schematic in

Figure 2. In the case of an MMC overlay, the %dil. should be as low as possible

(typically <10%) to retain the intended microstructure and properties of the

consumable. Some level of dilution is necessary (typically 3-5% minimum) to

ensure the overlay is metallurgically fused to the base material.

Percent Carbide Area (%carbide) – the level of primary carbide in the overlay

directly influences the MMC wear resistance, with high %carbide translating to

improved performance. The maximum %carbide can be significantly influenced

by the rate of degradation (dissolution) of the primary tungsten carbide during

welding. The %carbide level can be determined by measuring the area fraction

in a cross-sectional micrograph; this quantification was done using Clemex®

Microscopy Analysis Solutions image analysis software as shown in Figure 3.

Mean Free Path ( ) - the spacing of the carbides in the metal matrix drastically

influences the wear resistance of the MMC. By minimizing the , relative to the

particle size of the abrading media, a major increase in the wear resistance of

MMCs can be obtained (Hutchings, 1992). The value of can be quantified by

measuring the average distance between the carbides using image analysis

software as shown in Figure 4. The values were determined through the

thickness and the average and standard deviation determined.

A2A1A2 Dilution %

Figure 2 - Cross-Section of Overlay Bead Schematic

Figure 3 – Image of an MMC microstructure (left), and an Example of Utilizing the Image

Analysis to Determine %Carbide area (right)

Figure 4 – Image of an MMC Microstructure (left), and an Example of Image Analysis

Software Measurement Grid to Determine the Mean Free Path (right)

In addition to the analysis described above, the microstructures in Samples A &

B were also examined by scanning electron microscopy (SEM) with Auger Electron

Spectroscopy (AES) capabilities. AES analysis involves the bombardment of a sample

with a beam of energetic electrons, as with other SEM techniques; however, AES can

provide exceptional lateral resolution and can give reliable analysis of surface chemistry

(within 5 nm) with spatial resolution of <50 nm. The sensitivity is reasonably high and

can detect less than 1% of the atomic composition in a sample. The analysis was

completed to assess the ability of SEM-AES to delineate the specific components of the

microstructure and quantify the level of primary carbide dissolution.

RESULTS AND DISCUSION

Observations and Monitoring of the Welding Conditions The average voltage and current measured by the FLUKE 345 multimeter are

shown for each weld in Table 2, along with the corresponding transfer mode (based on

visual observation). The transfer modes listed in Table 2 are based on the classifications

applied to solid wire GMAW welding (Iordachescu and Quintino, 2008). However, it

should be noted that during cored-wire welding, a mixed-transfer behavior is often

achieved and the listings are the dominant mode observed. Specimen B was welded

with the manufacturer’s recommended parameters (see above), and the measured

current and voltage are within the suggested ranges. The heat input values, based on

the instantaneous power source readings and utilizing the new ASME IX formula, is

presented in Table 3. The results confirm that the progression from short-circuiting to

globular and to spray transfer results a net increase in average voltage, current,

instantaneous energy and the resulting heat input, for both WFS settings, when utilizing

cored MMC consumables. The influence of the heat input values will be discussed

below, with consideration of the macro and microstructure properties.

Table 2 - Average Voltage and Current Measured Using the Multimeter

Sample Voltage (V) Current (A) Transfer Mode 15.6 206A 16.0 222 Short-circuiting

20.2 179B 20.6 173 Globular

28.5 189C 28.5 188 Spray

15.9 380D 17.2 351 Short-circuiting

21.2 285E 21.2 293 Globular

31.9 298F 31.9 287 Spray

Table 3 - True Energy and Heat Input of Double Pass Welds

Wire Feed Speed 305 cm/min [120 ipm]

635 cm/min [250 ipm]

305 cm/min [120 ipm]

635 cm/min [250 ipm]

Transfer Mode True Energy ® (kJ) Heat Input (J/mm)* Pass 1 40.0 52.4 430 570Short-Circuiting Pass 2 43.3 47.5 466 516 Pass 1 54.7 76.6 570 774Globular Pass 2 51.1 77.5 532 783 Pass 1 68.9 100.9 718 1121Spray Pass 2 61.5 91.4 641 1016

*[mm]Length

][JEnergy Input Heat (ASME Boiler and Pressure Vessel Code, 2010)

Weld Bead Macro Features The bead surface appearance of the “low” productivity samples A, B and C are

shown in Figure 5, and the corresponding bead surface for the “high” productivity

samples D, E and F are provided in Figure 7. The top surfaces for the welds produced

with short-circuiting (Samples A and D) and globular (Samples B and E) transfer are

visually very similar. Sample F produced with spray mode is more uniform and aesthetic

in macro-appearance, which was attributed to the higher heat input and extended

solidification time at temperature which likely facilitates a more fluid weld pool.

The height and width of the weld overlays, measured using Clemex® calibrated

image analysis software, are shown in Table 4. As expected, higher wire feed speeds

offer higher deposition rates, which resulted in thicker and wider overlays. These

conditions promote increased productivity with fewer passes required to cover the

substrate. However, as shown by comparing the weld overlay cross-section of the “low”

and “high” productivity conditions (Figure 6 and Figure 8, respectively), as the heat input

increased there was a trade-off in the quality of the MMC microstructure. Quantification

of the key features of the microstructures is discussed below.

Table 4 - Dimensions of Double Pass Weld Overlays

Wire Feed Speed 120 ipm 250 ipm 120 ipm 250 ipm

Transfer Mode Height (mm) Width (mm) Short-Circuiting 4.64 6.27 9.61 9.66

Globular 4.26 4.72 10.18 13.39Spray 2.73 3.40 11.64 16.37

Weld Bead Microstructure As shown in Figure 6 and Figure 8, based on the parameters employed and the

resultant heat input achieved, there are significant differences in the %dil., %carbide,

distance, the carbide morphology, and the extent of their dissolution. The quantification

of these values is shown in Table 5.

Table 5 - Results of % Dilution, % Carbide Area and Mean Free Path

Mean Free Path ( m)Sample % Dilution % Carbide Area Avg Std Dev

A <1 45 84 73B 8 34 93 91C 5 40 60 60D <1 49 56 49E 16 38 89 84F 25 15 215 236

Considering the %dil. and %carbide values of Table 5, the data suggest that

welding an overlay with short-circuiting mode (Sample A and D) will yield the lowest

dilution and the highest fraction of primary carbides in the metal matrix. According to the

manufacturer, the original ratio between carbides and metal matrix in the wire

consumable was ~50 vol%, and the as-deposited MMC %carbide approached this level.

However, a dilution level of <1% is likely too low for some oil sands applications,

especially those involving impact loading or mechanical bending, and such an overlay

would likely be prone to spallation failure. Hence, although the short-circuiting mode

promoted low carbide losses, the low level of base metal fusion would potentially be

problematic during actual service. Further testing would be necessary to quantify the

actual wear performance of these samples, and investigate if mechanical loading could

cause an alternative failure mechanism. Conversely, on the other extreme, Sample F

achieved excessively high %dil. and low %carbide, which indicates that the heat input

was far too high to retain the desired metallurgical properties. This supports the

hypothesis that when welding MMC overlays, selection of parameters within the

appropriate “operating window” (i.e., not too far above or below the optimized values) will

cause a detrimental effect on the microstructure and resulting wear properties.

In the case of the mean free path ( ) values shown in Table 5, Samples C and D

achieved the lowest and most consistent primary carbide space (i.e., considering

average and standard deviation, respectively). However, the microstructures were

substantially different, as Sample C (Figure 6c) typically exhibited small, closely-packed

carbides, while sample D (Figure 8a) was comprised of mainly large carbides. This

implies that the measurements do not take into account variations in the primary

carbide size due to dissolution, but the %carbide values does indentify this variation.

Hence, alone does not provide a complete indication of the quality of the MMC overlay.

In future studies, it may be beneficial to measure the average as-deposited carbide size

(width). Additional future work could include wear testing of Samples C and D, to

determine the performance difference between smaller and large carbides, respectively.

The heat input levels of the “high” productivity overlays, particularly Sample F

achieved high values, with sparsely and unevenly space carbides, as confirmed in the

microstructure shown in Figure 8c. This is likely due to the very high temperatures and

long solidification time-at-temperature associated with the higher arc energy when using

spray transfer. This is consistent with previous work examining GMAW using

NiCrBSi/Wc-W2C wire consumables, where severe dissolution of the WC/W2C particles

was noted with increasing arc current (Badisch et al, 2008). Based on the development

of an un-optimized combination of dilution, carbide area and spacing, as a function the

elevated heat input, the MMC would be expected to have reduced wear resistance. This

is because the abrading particles can preferentially remove the matrix material before

fracturing and/or removing the remaining exposed tungsten carbide due to insufficient

matrix binding-material to ensure proper resistance to wear, impact, chipping, and micro-

fracturing. Additional wear testing and analysis would be necessary to verify this

concept.

Auger Electron Spectroscopy (AES)AES analysis was performed on the cross-sectioned surfaces of Sample A and B

near the tops of the overlays. The SEM-AES was used to identify the specific

components of the microstructure and quantify the level of primary carbide dissolution.

The results are shown in Table 6 and Table 7. Figure 9 compares the SEM micrographs

for the overlays obtained from Sample A and B.

Table 6 – Results of AES Analysis on Points Shown in Figure 10 (Sample B)

Element (atomic %) PointC W Ni B

1 36.0 64.0 - -2 92.7 7.3 - -3 6.1 18.7 75.1 -4 - - 71.5 28.5

Table 7 - Results of AES Analysis on Points Shown in Figure 11 (Sample A)

Element (atomic %) PointC W Ni B

1 11.1 6.2 82.6 -2 - - 71.1 28.93 25.4 74.6 - -4 40.1 59.9 - -5 41.8 58.2 - -

Chemical analysis of a representative area (similar to Figure 9) showed that nickel

content is close to 50% which corresponds with the 50% fraction expected in nickel

matrix for this wire consumable. Table 6 shows the analysis performed on a carbide

particle and the surrounding matrix from Sample B, and Figure 10 show the points the

where AES analysis was performed. The white particle (Point 1) is shown to be tungsten

carbide, which corresponds to a homogeneous primary W2C particle. The dark region on

the interior of the tungsten carbide particle (Point 2) was found to be almost pure carbon.

This is most likely residual graphite inclusions originating from the powder manufacturing

process as a result of incomplete mixing of the tungsten and carbon. The lighter phase

in the metal matrix (Point 3) corresponds with the Ni dendrites in the metal matrix, and

contains a significant fraction of tungsten and carbon in solution. The dark phase in the

matrix (Point 4) is shown to be composed of nickel and boron, most likely Ni3B. Hence,

these primary Ni dendrites are surrounded by a fine lamellae eutectic structure of Ni -

Ni3B at the inter-dendritic regions, see Figure 10. The Ni and Ni3B phases form a matrix

that is partially dendritic and lamellar microstructure, and has been observed in many

studies of NiCrBSi alloy based overlays (Flores et al, 2009). The addition of B into the

nickel alloy is a common melting point suppressant, since it is highly insoluble in nickel

and so small concentrations will promote the formation of a Ni/Ni3B eutectic with a

melting point of around 1100ºC (ASM Handbook). It has been found that the presence of

Ni3B also contributes significantly to the hardness of the metal matrix in these composite

overlays (Liyanage, 2010). In addition, there may be hardening of the matrix alloy from

super saturation of W and C, which may positively influence abrasive wear resistance

(assuming no or mild impact conditions), but may result in premature failure during

impact or combined impact-abrasive oil sands conditions (Harper et al, 2002). Between

some of the lamellae in this eutectic structure there were fine precipitates or secondary

carbide phases observed containing tungsten (as indicated by the very light regions in

Figure 10). The formation of secondary phases have been observed for PTA welded

MMC overlays (Anderson et al, 2003), but further work would be necessary to determine

the natures of these secondary carbide phases.

Examination of the primary carbides structures in Figure 9 identified that some

level of dissolution had occurred around the circumference of the primary carbide phase.

The level of dissolution appears to be greater in Sample B versus A (which correlates

with the increased heat input). This preferential dissolution has been observed during

PTA welding, and has often been referred to as the “halo effect” (Harper et al, 2002).

The high magnification image in Figure 11 shows a tungsten carbide particle with some

dissolution around the edge. Table 7 shows the composition of the numbered points.

The lighter matrix phase (Point 1) is the primary Ni dendrite containing W, and C and the

darker matrix phase (Point 2) is Ni3B. However, the tungsten carbide particle is not

homogeneous. AES analysis shows that the two phases are present, with Point 3

identified as W2C and Points 4 and 5 as WC (Nagender Naidu et al, 1991). In the “halo

region”, the W2C appears to be preferentially dissolved from the carbide compared to the

WC phase. This implies that the WC is thermally more stable than the W2C phase. This

correlates with evaluation of PTA welded MMCs, in which macrocrystalline WC-only

carbides are more resistant to dissolution than WC-W2C eutectic carbides (Harper et al.,

2002). Similar phases were observed in Sample B, in which a matrix containing primary

Ni dendrites with a Ni-Ni3B eutectic in the interdendritic regions, along with tungsten

carbide particles that had a similar structure of a WC-W2C eutectic core with a WC outer

shell with preferential W2C dissolution. Future work will include evaluation of the wear

properties as a function of the level of this dissolution mechanism.

SUMMARY AND CONCLUSIONS 1) The changing of parameters to achieve the three different transfer modes drastically

changed the macro and microstructure and properties of the MMC weld overlay. As

the energy and heat input increased, the amount of dilution, and carbide degradation

increased as evident by % carbide and results.

2) Severe carbide degradation and areas devoid of primary carbides occurred for the

spray transfer mode at the high current and voltage levels. These MMCs will likely do

poorly in service, and hence such “high productivity” parameters are not

recommended. Conversely, too low of heat input may lead to insufficient fusion to the

base material and potential for failure by other modes than wear. These results

emphasize the importance of utilizing operating “windows” to optimize and maintain

MMC quality.

3) AES analysis show the metal matrix is composed of Ni and Ni3B arranged in dendritic

and lamellar microstructure. AES analysis also identified that a eutectic is present,

identified as W + W2C, and WC.

4) High magnification SEM analysis showed that there was still small amounts of

dissolution of primary carbides around the edges but the bulk of the carbide particles

remain intact, also known as the “halo effect”. This suggests that the WC and W2C

eutectic structure of the carbides may be inferior to homogeneous solid WC carbides.

FIGURES

(a) Parameter A (b) Parameter B (c) Parameter C

Figure 5 - Bead appearance and weldment cross-sections of low productivity samples A, B & C (left to right)

(a) Parameter A (b) Parameter B (c) Parameter C

Figure 6 – Mapping of the Macro Cross-Sections of Samples Welded with Parameter A, B and C (left to right). The scale bar on all images is 0.5mm (500 m).

(a) Sample D (b) Sample E (c) Sample F

Figure 7 - Bead appearance and weldment cross-sections of high-productivity samples E, E & F (left to right)

(a) Sample D (b) Sample E (c) Sample F Figure 8 – Mapping of the Macro Cross-Sections of Samples Welded with Parameter D,

E, and F (left to right). The scale bar on all images is 0.5mm (500 m).

(a) (b)

Primary Carbides

Carbides showing some level of dissolution around the perimeter, or

the “halo” effect

Figure 9 – SEM micrographs of Sample A and B respectively

Ni dendrites

34

Primary Carbide

2 Ni+Ni3B eutectic - lamellae structure of Ni (light phase) and Ni3B(dark phase)

1

Fine precipitates or secondary carbide phases containing W

Figure 10 – SEM micrograph of a tungsten carbide particle and surrounding metal matrix from Sample B

“Halo Region” around a primary carbide where the WC phase remains intact and the W2C phase has undergone preferential dissolution. (note: spot 4 & 5 are WC and spot 3 is W2C)

Figure 11 - High Magnification Image (4000x) of Carbide Particle from Sample A

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