OPTIMIZING COPPER TO COPPER CONTACT PERFORMANCE IN MARINE BATTERY DISCONNECT SWITCHES By Eric J. Graham A Project Submitted to the Graduate Faculty of Rensselaer Polytechnic Institute in Partial Fulfillment of the Requirements for the Degree of MASTER OF MECHANICAL ENGINEERING

Approved:

Dr. Ernesto Gutierrez-Miravete Project Adviser

Rensselaer Polytechnic Institute Troy, New York November 2005 (For Graduation December 2005)

CONTENTS LIST OF TABLES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

v

LIST OF FIGURES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

vi

Project Schedule / Time Line . . . . . . . . . . . . . . . . . . . . . . . . . . . vii ABSTRACT . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . viii 1. MARINE ROTARY BATTERY SWITCH CONSTRUCTION AND CONSIDERATIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.1 Typical Switch Construction

. . . . . . . . . . . . . . . . . . . . . .

1.2 Battery Switch Ratings and Implications . . . . . . . . . . . . . 1.2.1 Mechanical Endurance . . . . . . . . . . . . . . . . . . . 1.2.2 Voltage Limits . . . . . . . . . . . . . . . . . . . . . . . 1.2.3 Continuous Current . . . . . . . . . . . . . . . . . . . . . 1.2.4 Intermittent Current . . . . . . . . . . . . . . . . . . . . 1.2.5 Cranking Current . . . . . . . . . . . . . . . . . . . . . . 1.2.6 Operating Environment . . . . . . . . . . . . . . . . . . 1.2.6.1 Temperature Limits . . . . . . . . . . . . . . . 1.2.6.2 Humidity / Gaseous Contaminants / Salt-Spray 1.2.6.3 Vibration . . . . . . . . . . . . . . . . . . . . . 1.2.6.4 Potentially Explosive Atmospheres . . . . . . . 1.2.7 Insulating Materials . . . . . . . . . . . . . . . . . . . . 1.2.8 Ergonomics . . . . . . . . . . . . . . . . . . . . . . . . . 1.2.9 Attached Cable Ranges . . . . . . . . . . . . . . . . . . . 1.2.10 Economic / Financial . . . . . . . . . . . . . . . . . . . .

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3 3 3 4 4 4 5 5 6 6 6 7 8 8 8

1.3 Design Considerations of Switch Contacts . . . . . . . . . . . . . . 1.3.1 Conductor Material . . . . . . . . . . . . . . . . . . . . . . . 1.3.2 Conductor Thickness . . . . . . . . . . . . . . . . . . . . . . 1.3.3 Sealing of Enclosure . . . . . . . . . . . . . . . . . . . . . . 1.3.4 Internal Enclosure Atmosphere . . . . . . . . . . . . . . . . 1.3.5 Internal Conductor Sizing . . . . . . . . . . . . . . . . . . . 1.3.6 Contact Surface Finishing . . . . . . . . . . . . . . . . . . . 1.3.6.1 Expectation of Surface Impurities (Cupric / Cuprous Oxide) . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.6.2 Use of Liquid / Solid Lubricants . . . . . . . . . .

. . . . . . .

9 9 9 10 11 11 12

ii

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1

. 13 . 14

2. ELECTRICAL CONTACT RESISTANCE FACTORS . . . . . . . . . . . 15 2.1 Summary of Boundary Conditions and Limitations . . . . . . . . . . 15 2.2 Contact Resistance Parameters and Calculations . . . . . . . . . . . . 16 2.2.1

Practical considerations of contact dimensions . . . . . . . . . 17

2.2.2

Super Temperature Calculations and Implications . . . . . . . 18

2.3 Plating Material and Considerations . . . . . . . . . . . . . . . . . . 19 2.3.1

Effect on Contact Resistance . . . . . . . . . . . . . . . . . . . 20

2.3.2

Effect on Oxidation, Fretting, and Corrosion Control . . . . . 22

2.3.3

Effect on Sliding Contacts . . . . . . . . . . . . . . . . . . . . 24

2.4 Film Resistance and Sources . . . . . . . . . . . . . . . . . . . . . . . 24 2.4.1

Natural Oxidation . . . . . . . . . . . . . . . . . . . . . . . . 25

2.4.2

Corrosive Films . . . . . . . . . . . . . . . . . . . . . . . . . . 26

2.5 Sliding / Wiping Contacts . . . . . . . . . . . . . . . . . . . . . . . . 26 2.5.1

Contact Wear, Friction, Shear, and Delamination . . . . . . . 27

2.5.2

Boundary Lubrication . . . . . . . . . . . . . . . . . . . . . . 28

2.6 Summary of Findings . . . . . . . . . . . . . . . . . . . . . . . . . . . 30 3. CU to CU CONTACT THERMAL RESISTANCE OPTIMIZATION . . . 32 3.1 Contact Geometry and a-spot Optimization . . . . . . . . . . . . . . 32 3.2 Lubricant Effects . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34 3.3 Surface Roughness . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35 3.4 Plating Effects

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

3.5 Rotary Contact Break-In . . . . . . . . . . . . . . . . . . . . . . . . . 35 4. THERMAL / ELECTRICAL MODEL OF ROTARY DISCONNECT SWITCH 37 4.1 Assumptions / Fixed Parameters . . . . . . . . . . . . . . . . . . . . 37 4.2 Variable Parameters . . . . . . . . . . . . . . . . . . . . . . . . . . . 37 4.3 Output Parameters . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38 4.4 Thermal Effects of Contact Resistance . . . . . . . . . . . . . . . . . 38 4.5 Steady-State Equations . . . . . . . . . . . . . . . . . . . . . . . . . . 38 4.5.1

Power Connection Cables . . . . . . . . . . . . . . . . . . . . 38

4.5.2

Internal Contact Pairs . . . . . . . . . 4.5.2.1 Electrical Contact Resistance ture Calculations . . . . . . . 4.5.2.2 Thermal Contact Resistance . iii

. . . . . . . . . . . . . 40 and Super Tempera. . . . . . . . . . . . . 40 . . . . . . . . . . . . . 40

4.5.3

Copper Rotor Conductor . . . . . . . . . . . . . . . . . . . . . 40

4.6 Finite Difference Model . . . . . . . . . . . . . . . . . . . . . . . . . . 42 4.7 Results of Calculations . . . . . . . . . . . . . . . . . . . . . . . . . . 42 5. EXPERIMENTAL VALIDATION OF ELECTRICAL/THERMAL MODEL 43 5.1 Experimental Set-Up and Procedure . . . . . . . . . . . . . . . . . . 43 5.1.1

Tests Conducted . . . . . . . . . . . . . . . . . . . . . . . . . 43

5.1.2

Measurement of Switch Resistance . . . . . . . . . . . . . . . 44

5.2 Measurement of Internal and External Conductor Temperatures . . . 45 5.3 Conducted Experiments and Results . . . . . . . . . . . . . . . . . . 46 5.4 Results and Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . 47 5.4.1

Spring Force Variations . . . . . . . . . . . . . . . . . . . . . . 47

5.4.2

Lubricant Variations . . . . . . . . . . . . . . . . . . . . . . . 48

5.4.3

Plating Variations . . . . . . . . . . . . . . . . . . . . . . . . . 48

6. TABLE OF SYMBOLS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49 LITERATURE CITED

. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

iv

LIST OF TABLES 5.1

Construction Parameters for Samples . . . . . . . . . . . . . . . . . . . 47

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LIST OF FIGURES 1.1

Cross-Section Exploded View of Rotary Switch . . . . . . . . . . . . . .

1

1.2

Isometric Exploded View of Rotary Switch . . . . . . . . . . . . . . . .

2

1.3

Cross-Section Assembled View of Rotary Switch . . . . . . . . . . . . .

2

1.4

Current Flow Constriction From Current Re-Direction . . . . . . . . . . 10

1.5

Transient Temperature Profile Accross Contact . . . . . . . . . . . . . . 12

2.1

Contact Point Location Affects . . . . . . . . . . . . . . . . . . . . . . . 18

2.2

Contact Resistance vs. Force . . . . . . . . . . . . . . . . . . . . . . . . 29

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Project Schedule / Time Line Plan Date Deliverable Act Date th Aug 19 Submit Project Outline and Preliminary List of Refer- Aug 19th ences Aug 26th Submit Project Time Table, Revised Outline, Expanded Aug 29th Background Section nd Sep 2 Finalized Outline, Completed Background Sep 5th th Sep 9 Preliminary Oxidation and Corrosion Section / ExperiSep 5th mental Analysis Proposal Sep 16th Preliminary Fretting Section / Finished Oxidation Sec- Sep 17th tion / Initiate Procurement of Experiment Materials rd Sept 23 Preliminary Wear and Shear Section / Finished Fretting Sep 25th Sect. th Sep 30 Preliminary Lubricant Section / Finished Wear Section Oct 9th th Oct 7 Complete All Existing State of Art Section Oct 15th Oct 14th Preliminary Thermal / Electrical Model of Rotary Switch Oct 31st Oct 21st Expand Thermal / Electrical Model Oct 28th Begin Experimental Studies / Finalize Therm / Elect Model th Nov 4 Continue Experimental Studies / Begin Experiment Results th Nov 11 Finish Experiments / Continue Documenting Results Nov 18th Preliminary Draft of Masters Paper th Nov 30 Final Smooth Draft Submitted

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ABSTRACT Marine battery switches are a unique animal within the broad group of electromechanical products known as electrical switches. Their uniqueness is gained both by the performance standards which they must be measured against as well as the environment within which they are required to perform. This paper will attempt to achieve three main goals. First, it will act to summarize the unique set of design challenges encountered when one sets-out to determine a contact structure for such a device. Second, the critical contact resistance parameters which must be considered will be investigated and scrutinized, with specific emphasis on two specific types of copper to copper contact mating schemes common in this field; transverse rotary, and plunger. Lastly, an investigation into the thermal resistance between copper to copper contacts will be undertaken with particular interest in the relationship between contact design parameters, thermal resistance, and overall switch performance.

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CHAPTER 1 MARINE ROTARY BATTERY SWITCH CONSTRUCTION AND CONSIDERATIONS 1.1

Typical Switch Construction Rotary type marine battery switches are typically constructed as shown in

Figures 1.1 and 1.2 below. Their relatively simple construction consists of an operating knob for manual operation of the switch, typically connected to an internal rotor shaft to translate rotational torque from the knob to the copper rotor. The internal spring responsible for developing normal contact force is compressed between the rotor shaft and the copper rotor, thus forcing the copper rotor to contact the copper terminals. It should be noted that insulative housing pieces have been removed from Figures 1.1 and 1.2, which would be located between the Knob and Rotor Shaft,

Figure 1.1: Cross-Section Exploded View of Rotary Switch 1

2

Figure 1.2: Isometric Exploded View of Rotary Switch

Figure 1.3: Cross-Section Assembled View of Rotary Switch

3

as well as between the Copper Rotor and Copper Studs (the copper studs actually protrude through the bottom insulator in order to provide an electrical path. As can be seen, rotating the Knob (and inherently the Copper Rotor) 90 degrees will place the contacts of the Copper Rotor either in direct contact with the Copper Studs or separated from the terminals. Figure 1.3 below shows the assembled rotary battery switch from Figures 1.1 and 1.2, clearly showing the electrical conducting path from the input terminal, into the copper stud, through one copper to copper contact into the copper rotor, back through a second copper to copper contact into a second copper stud, and finally out through the output terminal.

1.2 1.2.1

Battery Switch Ratings and Implications Mechanical Endurance Marine battery switches are considered safety disconnect devices, and are not

intended to switch live loads or currents. As such, the mechanical endurance requirements do not result in arcing of contacts. Rather, the mechanical wear of the switch contacts is the primary concern. Because of the typically high normal forces between contacts, rotary switches are particularly prone to galling and other plastic deformation failures of the conductor materials, while plunger style switches must endure the multitude of forceful impacts between the electrical contacts. 1.2.2

Voltage Limits Marine battery switches operate in low voltage DC environments less than

42 Vdc. Typical marine battery systems operate at either 12Vdc, 24 Vdc, or 36 Vdc. While this is much less than many AC or DC industrial environments, it will be demonstrated later that these voltages are more than sufficient to create contact arcing upon separation (if required to do so), as well as create contact area super-temperatures in excess of the softening or melting temperature of the base materials.

4

1.2.3

Continuous Current The continuous current rating of a marine battery switch is the highest cur-

rent which can be continually transferred through the switch for a period of 1 hour without the external temperature of the wire connecting terminal exceeding 100 ◦ C above the ambient. Unfortunately, other factors limit the ability for this metric to be fully utilized. Notable, the maximum ambient temperature in the marine environment has widely been accepted to be 50 ◦ C, and the typical marine wiring cable insulation covering is rated at 105 ◦ C. Since the terminal temperature referenced above is directly in contact with the cable insulation, the actual maximum terminal temperature is only 55 ◦ C. For the purpose of this research, a fixed continuous current of 250 A will be the required rating and will be fixed in all calculations. 1.2.4

Intermittent Current Marine battery switches are required to pass significant currents during a short

period of time. Engine starting events, AC inverter loads, and bow thruster utilization, can draw significant currents for up to 5 minutes. Therefore an intermittent rating for these switches has universally been developed. Under this rating, the wire connecting terminal mentioned above cannot exceed a certain temperature rise within 5 minutes. As described above, because of other mitigating factors, this temperature rise limit is 55 ◦ C. The importance of this rating as well as the cranking rating yet to be discussed lies in the desire to achieve maximum current throughput in a shortened period of time. The thermal dynamics of this transient event within the switch will be discussed later, however, it is important to note here the origin of the demand. For the purpose of this research, a fixed intermittent current of 350 A will be the required rating and will be fixed in all calculations. 1.2.5

Cranking Current Similar to the Intermittent Current rating, the Cranking Rating for a marine

battery switch is intended to measure the capacity of a switch to withstand a high

5

amount of current for a shortened period of time. In this case, the time duration is 2 minutes, with temperature limitations equal to the first two ratings already described. As will be shown later, a further limitation to the wire terminal temperatures of marine battery switches during short-term (Intermittent and Continuous) testing is the extreme gradient of temperature between the external wire terminals and the internal copper conductors. During such short transients, the thermal capacitance of the wiring system for the battery switch has little problem managing the higher than continuous currents. However, the thermal resistance of the internal switch contacts effectively insulate the internal copper rotor which is heated not only by the current flowing through the part, but by the constriction resistances of both sets of contacts, as will be explained later. For the purpose of this research, a fixed cranking current of 500 A will be the required rating and will be fixed in all calculations. 1.2.6

Operating Environment Marine battery switches must endure a wide breadth of extreme operating

conditions, including broad storage and operating temperatures, high humidity airborne salt levels, gaseous environments from diesel and gasoline engines as well as cleaning chemicals, high vibration loading, and potentially explosive atmospheres from stored gasoline or fuel leaks. The two primary guidance documents driving the design, testing, and performance specification of marine battery switches is Underwriter’s Laboratories (UL) standard 1107 (UL 1107, Marine Battery Switches), Society of Automotive Engineers (SAE) standard J1171 (SAE J1171, Ignition Protection of Automotive Electrical Products), and American Boat and Yacht Council (ABYC) directive E-11 for electrical equipment aboard vessels. 1.2.6.1

Temperature Limits

UL 1107 defines the required storage temperature of a switch to be between −40 ◦ C and 60 ◦ C. Within these limits, the product should be able to withstand handling and shipment, but not necessarily operate. UL defines the required operating temperature limits within the referenced standard to be between −10 ◦ C

6 and 50 ◦ C. Because these products produce considerable heat while operating, the lower temperature limit is of little consequence in this discussion. However, the upper limit, so chosen to match the expected maximum temperature within a vessels engine compartment, provides the lower basis temperature used to define the acceptable internal copper temperature of the battery switch. 1.2.6.2

Humidity / Gaseous Contaminants / Salt-Spray

The marine operating environment is not considered benign with respect to material longevity. An almost constant 90 percent plus humidity ensures that water molecules are present in abundance. Gasoline and diesel motors generate sulfur rich exhaust and fumes and often leak petroleum based oils in standing bilges, a multitude of cleansers exist in close proximity to electrical products, and airborne salt molecules are ever present to lend a hand at corrosion. While no minimum levels or standards are currently used to evaluate marine battery disconnect switches with regard to humidity, gaseous contaminants, and salt-spray; it should generally be acknowledged that the intended application environment is relatively hostile and that unless extreme precautions and measures are taken, surface contamination will occur. 1.2.6.3

Vibration

The presence of gasoline and diesel powered engines aboard almost every vessel which requires a battery disconnect switch is well known. The vibrations inherent in operating these machines are often quite severe. Little work has been done to quantify amplitudes and frequencies of these vibrations because each boat manufacturer implements their engine mounting is a slightly different manner, and engine manufacturers vary in their ability to remove vibration during operation. It should suffice to say that significant vibrations exist aboard recreational craft and should be strongly considered when designing battery disconnect switches, including 1.2.6.4

Potentially Explosive Atmospheres

Gasoline tanks and portable containers are often left in closed compartments in the company of electrical switches aboard recreational craft. Historically, when

7

fume levels became concentrated, electrical sparks from switches located in the same compartments as the gasoline containers would ignite the gasoline fumes resulting in an explosion and subsequent fire aboard the vessel. ABYC requires that battery disconnect switches, intended for use in a compartment which will or may contain a gasoline container, be tested for ignition protection. The ignition protection test requires the successful repeated operation of the switch within an atmosphere with a high concentration of gasoline fumes. The successful switch design capable of withstanding an ignition protection test must be well sealed. Because the test does not provide for the time necessary for switch ’breathing’ between internal switch and external atmospheric condition, a hermetic seal is not necessary to pass this test and therefore is not a common design element. However, because of the nature of the sealing required to pass ignition protection testing, the internal compartment of a typical battery disconnect switch will be free from particles and contaminants not able to pass through the o-ring seals typical of switch designs. 1.2.7

Insulating Materials While much could be discussed relative to the insulating material choices for

battery disconnect switches, the primary focus area for this discussion is the maximum temperature limit of the insulating materials in direct contact with the current carrying copper components. While thermoplastic materials such as Nylon have thermal limits of approximately 90 ◦ C, Thermoset plastics can support temperature in excess of 150 ◦ C. When considering the current limitation of any single switch design, the temperature limit of the insulating material; whether that be the power cable insulation, the switch backplate plastic housing, or a plastic insulator in contact with the copper rotor; will define the maximum possible current. This is because the typical contact super temperatures and resulting conductor temperatures will be well below the softening or melting points for copper, but they may well be above the temperature limits for the plastic insulation materials. For the purposes of this investigation, a glass-filled thermoplastic material with an absolute temperature limit of 150 ◦ C will be assumed for all battery switch

8 insulative plastics, and a power cable insulation temperature limit of 105 ◦ C will be assumed, corresponding to the typical available cable offering for mariners and boat builders. 1.2.8

Ergonomics While electrically activated plunger battery switches do not require user in-

teraction, rotary switches must be turned by hand in order to turn on or off. The amount of force required to perform this operation, while not limited by formal standards, must allow an average adult to easily turn on or off the switch with one hand in a reasonable amount of time. The above ergonomic limitations will most directly affect electrical contact design and performance through the limitation of internal spring force used to generate normal contact pressure. Because this force is by far the largest contributor to the rotational friction force resisting the turning of the knob, it will naturally limit the working range of force for the internal spring. 1.2.9

Attached Cable Ranges The temperature limitations, which define the official ratings of marine bat-

tery switches, are defined by many design factors and choices, including the size of internal bus bars, the force between copper contacts, etc. While these and other mechanisms affect the amount of internal heat generation, a major factor in determining the resulting terminal temperatures lies in the size of the cable used to connect to the device. Because electrical cables generate heat based on the current flowing through as well as the cross-section of copper inside the cable, larger electrical cable will allow more current to flow given a fixed conductor temperature, or similarly pass the same amount of current with a lower cable heat rise. Because of this system level affect, a standard conductor size of 4/0 AWG cable will be used during the battery switch evaluations within this research. 1.2.10

Economic / Financial

Marine battery switches are sold through distribution and retail outlets at very reasonable prices. While they perform a critical task of allowing isolation and

9

emergency disconnect between DC batteries and DC loads, the market expectation for product cost and value placed on a quality product are both relative low. This reality drives many battery switch design decisions, including conductor material and plating selection as well as determination of insulative plastic materials and the inherent thermal and electrical properties they possess.

1.3 1.3.1

Design Considerations of Switch Contacts Conductor Material The base conductor material of choice for marine battery switches is copper.

More specifically, electrolytically tough pitch copper (ETP) or C11000, which is made by refining the copper electrolytically. Because this process is prone to embrittlement, oxygen-free high conductivity copper (OFHC) is often used instead. OFHC is produced from refining ETP copper in the presence of an inert gas free of metallic oxidizers. Battery switches are sometimes also built with brass as the base material, or otherwise somewhere within the electrical conducting path. While brass provides good corrosion and other properties advantageous to the marine environment, the poor electrical conductivity it possesses compared to ETP copper makes this choice quite undesirable, and will not be considered herein. 1.3.2

Conductor Thickness The internal copper components within a typical battery isolation switch are

made from stamped copper bar or plate material, because manufacturing processes and low tooling costs allow for low part costs. However, limitations to material thickness result from the stamping processes utilized. Typical material thickness of internal copper parts are between 0.125 inches and 0.187 inches. Conductor thickness may play a measurable part when considering overall battery switch resistance. While not technically part of the contact-contact pont resistance, material thickness plays a part in the overall resistance calculations by adding a second constriction resistance in the area of the contact. After passing through the contact constriction, a rotary battery isolation switch current path

10

Figure 1.4: Current Flow Constriction From Current Re-Direction requires the current to ’turn’ 90 ◦ before proceeding. Depending on the thickness of the conductor material, the constriction of current flow lines in this area may measurably affect overall switch resistance and resulting performance. 1.3.3

Sealing of Enclosure Because of the harsh marine environment and atmosphere, sealing of a battery

switch enclosure is highly recommended. This provides excellent protection for the internal contacts against corrosion and tarnishing once sealed, and eliminates the ability for water and other moisture to attack the contacts. An additional benefit of sealing battery switch enclosures is that ignition protection is accomplished. Ignition protection, as defined for marine battery switches, is the ability of the switch to make and break live current while immersed in an explosive atmosphere, without causing that immersed atmosphere to ignite. The seal of ignition protection, as one could surmise, can be a powerful design and marketing advantage because of the inherent safety it implies. A negative side effect of a well-sealed enclosure is the limited thermal heat dissipation it allows. By effectively trapping the internal air, convective heat transfer from internal bus bars is severely hindered. Relative to other heat transfer mechanisms such as heat conducted out of the switch via the copper cables and then convected to the surrounding air through the insulation, the trapped air inside the switch effectively insulates the internal components from any

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heat escape other than through conduction to power cables. 1.3.4

Internal Enclosure Atmosphere While ignition protection from sealed enclosures is considered a high level

of environmental protection, the enclosure seal is far from considered hermetically sealed. As a result, breathing of air between the inside of such a switch and its external environment is inevitable, and will result in some level of moisture and air contaminants entering the switch. As an alternative, some electrically actuated solenoid plunger switches employ a hermetically sealed chamber, which after being fitted with inflow and exhaust fittings is evacuated of air an filled with an inert gas. After being sealed, these switches not only keep harsh external elements from entering the internal switch structure, the inert gas inside reduces arcing between contacts when sufficiently high voltages and/or current are connected or disconnected. The high costs of hermetically sealing an enclosure prohibit such actions from being taken in a low cost high volume product such as mechanically actuated rotary switches, and will therefore not be considered herein. 1.3.5

Internal Conductor Sizing The nature of battery disconnect switches and their respective ratings dic-

tate conductor sizing to a large extend. Because these switches are not required to operate under load, the major performance characteristics which define the value proposition between competing products are continuous and short-term current carrying capacity. As will be shown later, the size of the internal copper conductors, both in terms of thickness and in general overall mass, play an important part in defining the limitations to these two performance characteristics, most importantly the short-term current capacities. The reason why the size of the internal copper rotor in particular has a critical role in determining the current ratings lies in the physics of the transient shortterm high current events. During these events, currents in excess of 150 to 200 percent of the continuous rated current are passed through the switch. The power loss across the two contacts is dissipated through the two power cables; however, the power generated at both contact pairs is also conducted into the copper rotor.

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Figure 1.5: Transient Temperature Profile Accross Contact Depending on the heat capacity (derived from the mass) of the copper rotor, the rotor temperature reaction will vary from a) A slow increase closely coupled to the increase in power terminal temperature, to b) A rapid increase over a short period of time until the rotor temperature approaches the contact super temperature, after which the rotor temperature rises in concert with the power terminal temperature because the majority of the heat produced at the contact flows towards the power cables. Figure 1.5 shows the typical transient temperature profile of the contact region and the conductor material before and after the contact. The initial time measurement at the bottom is close to six times the temperature of the conductor, which has not yet increased in temperature due to it’s thermal capacity and the short time duration. By the last time point, the conductor has increased in temperature from both the current applied and the contact super-temperature, however, the ability for the conductor to remove heat energy away from the contact has lowered the relative difference between conductor and contact to slightly over two times. Figure 1.5 was take from the work of Filippakou [13] regarding coating material effects on copper joints. 1.3.6

Contact Surface Finishing With only very minor exceptions, the vast majority of rotary battery isolation

switches are constructed using copper to copper contact interfaces without plating materials applied. Not withstanding the previous statement, consideration and examination of tin, silver, and nickel plating as an option to improved switch performance will later be investigated. Because of cost obligations, more sophisticated

13

plating options will not be considered in this text. Plating internal contacts in sliding rotary switch has many potential drawbacks, including the introduction of insulative oxide layers from fretting, wearing the surface plating rapidly after just a few cycles of operation, and of course the potential increased product cost. 1.3.6.1

Expectation of Surface Impurities (Cupric / Cuprous Oxide)

As will be discussed later in detail, copper tends to oxidize immediately when exposed to an oxygen atmosphere. Many authors have discussed contact resistance in terms of cleaned samples free of surface contaminants, and measurements taken in vacuum chambers to avoid variation due to the effects of oxidation. Such laboratory evaluations are necessary in order to evaluate proper variations and eliminate errors, but have little absolute use when evaluating marine battery disconnect switches. The cost limitations of this product category immediately eliminate the concept of performing cleaning before assembly. Similarly, the growing cost reduction methods in manufacturing chain selection often results in copper contact plates manufactured across oceans from where they will be assembled, and then shipped for several weeks in a relatively humid and salt-rich environment. Finally, the ability to evacuated the internal compartment of a rotary battery switch is again diminished by the cost implications presented, thus ensuring a non-hermetically sealed chamber capable of ’breathing’ external atmospheric gases. The typical marine installation location of such switches will not only be high humidity, gases from engines, cleaning agents, and industrial adhesives related to the building of the vessel itself are all prone to enter the product Due to the aforementioned issues, it is unreasonable to expect an installed switch to be free from surface oxidation. The amount, thickness of oxidation layers, and implications therein will be discussed later. It is promising, however, that the sliding operation of the switch, and the potentially high enough fritting voltage involved, may render this issue less important to the discussion at hand.

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1.3.6.2

Use of Liquid / Solid Lubricants

The switches being investigated operate by sliding during the rotation of an operational knob. Some variants operate back-and-forth only, rotating often in a fixed 90 ◦ degree path, while others allow for a full 360 ◦ of motion. The issues related to both styles of switches are similar. Depending on contact geometry, rotor force and resulting contact pressure, relative material hardness, and other factors, contacts can and often do wear down during the typical service life. It is customary in the industry to apply a surface lubricant between mated rotary contacts to ensure minimal wear during the life of the product. The resulting performance of the product with varying lubricants is therefore an area of significant investigation. Reducing electrical contact resistance, improving heat conduction from the contact a-spots, remaining effective for thousands of operation cycles, and reducing dynamic friction during operation are all objectives of lubrication.

CHAPTER 2 ELECTRICAL CONTACT RESISTANCE FACTORS Significant research has been conducted for many decades regarding the subject of contact resistance theory and real-world validation. The most widely referenced work on the subject was written by Holm [22]. Holm’s dedication to the subject is widely known and appreciated, his efforts also resulted in the on-going Holm Conference on Electrical Contacts, now facilitated through The Institute of Electrical and Electronics Engineers, Inc. (the ”IEEE”). Many references within this paper have been cited from the Holm Conference proceedings. Because Holm’s book was most recently updated in 1967, Slade [40] made an effort in 1999 to provide a compilation of the current state-of-art knowledge on the subject of electrical contacts. This chapter will review and examine the current thinking from all of the above sources as well as ancillary documents.

2.1

Summary of Boundary Conditions and Limitations The present investigation is only concerned with marine rotary battery switches

as outlined in the previous chapter. Consequently, the following assumptions and limitations are explicitly followed: • Only ETP Copper substrates will be used. • Possible plating materials will be limited to Tin, Nickel, and Silver • The effect of current direction changes will not be considered • Typical operation voltage is 12 Vdc • Contact load will vary between 2 and 15 lbs. • Electrical performance of contacts during sliding is not critical (stationary only) • Contacts are surrounded by Air at sea level 15

16 • Contact geometry will consist of a convex spherical arc on a flat disc

2.2

Contact Resistance Parameters and Calculations Mating contacts in marine battery switches typically consist of protruding

circular contacts on rotating rotors in contact with stationary flat bus bars, both parts consisting of ETP copper bus bar stock. While macroscopically both contact surfaces appear smooth, microscopic roughness exists and surface asperities consisting of peaks and valleys dominate the landscape. When the contact pairs are brought together, the two parts make metallic contact at only the microscopic locations where the two parts touch. These locations, known as ’a-spots’ and typically simplified in analysis as circular, are the basis of the Constriction Resistance. The constriction resistance between monometallic contacts [46, 6], known as Maxwell’s formula for a single circular spot is defined as Rc =

ρ 2a

(2.1)

valid for a wide range of problems. Although the geometry of the two contact battery switch would suggest that each contact meets at only one point, real-world contacts have more than one a-spot contact point because of the contact loading, the softness of the copper substrate, and the microscopic peaks which exist. In order to estimate the total contact area from the multitude of microscopic a-spots, Timisit [46] outlines the well accepted plastic deformation model for contacts which results in a powerful relation between normal contact force F , material hardness H, and the contact area Ac as F = Ac H

(2.2)

Holm [22] has shown that hardness (H) is related to the yield stress (σy ) by H = 3σy

(2.3)

This relation clearly shows that softer materials and larger contact forces have

17

a linearly positive effect of effective contact area. This relation can be used to estimate electrical contact resistance. The relationship between contact area and radius, Ac = πa2, can be combined with the single a-spot resistance formula resulting in:




Rc =

ρ2πH 4F

(2.4)

Examining the equation, one sees that in general contact resistance in effected by the square root of contact force and material hardness, while it is affected linearly by material resistivity. The choice of copper for a contact material is therefore an obvious choice due to its high conductivity and relative softness. High force compression springs are regularly used for rotary high current switches as well. As will be discussed later, other factors may suggest that alternate interface materials would be more beneficial. 2.2.1

Practical considerations of contact dimensions Microscopically it could be argued that with conventional material process,

asperity distribution is relatively random and uniform [44]. Indeed, much discussion has covered whether surface asperities heights follow a purely Gaussian or some form of a truncated Gaussian distribution [34]. However, these discussions weight more deeply on the plastic deformation models used to define effective contact area. Macroscopically the physical nature and design of a contact can have significant effect on contact resistance. Timisit [46] shows the effective contact area for a variety of asperity distributions. Figure refa-spot-options from Timisit’s work [44] shows the significant changes that can result from different a-spot distributions. Indeed, Timisit found contact resistance reductions of 50 percent when comparing equal sized circular disc a-spots to those that were ring-shaped [44]. If this concept were to be extended to beyond the microscopic realm and into the contact design area, one could see that contacts designed to meet at distributed points spread-out as much as possible would support a reduced effective constriction resistance. Based of the ability to maintain close geometrical relations either through accurate design practices or by contact ”break-in”, once could expect a contact pair to macroscopically meet at several distributed a-spots and improve overall contact resistance

18

Figure 2.1: Clusters of a-spots with corresponding radius of equivalent single contact (shaded area) and Holm radius (outer radius) A striking example of the benefit of macroscopically spreading-out the contact spots is found by Braunovic [7] who studied the contact resistance of overlapping bolted wide bus-bar joints. It was demonstrated that by cutting lengthwise slots in the bus bars between bolt locations contact area could be increased by 1.5 to 1.7 times that of a non-slotted joint. Braunovic’s results also indicated that cutting 45 ◦ grooves in mating surfaces improved resistance, theoretically by increasing effective contact area [7]. This flys directly in the face of Holm’s assertion that contact areas is dependent only on contact force and material hardness. While Braunovic concluded that rougher surfaces have lower contact resistances, this was most likely a result of the spreading-out of a-spots rather than the roughness as a whole. Indeed, surface roughness comes into play only with very small contact loads as shown by Timisit [44]. He showed definitavely how contact resistance lessened significantly when loading was initiated. However, when un-loading was commenced, contact resistance did not increase appreciably. His assertion that asperities are deformed plastically when loading is increased matches findings from Milanez and Bryant [34, 10]. 2.2.2

Super Temperature Calculations and Implications Possessing a knowledge of the temperature of the contact interface is very

important for a number of reasons. Because it is nearly impossible to measure

19

these points, it is essential to know how high the contact temperatures reach in order to consider issues such as material softening and melting, ramifications on system temperatures during steady-state and transient time-frames, and effects on boundary lubricants. While simplified voltage-temperature relations exist for contact systems with relatively little rise in super temperature over the bulk material temperature (less than 3 ◦ C), the nature of the high amperage rotary switch designs lend themselves more to high temperature difference between bulk conductors and internal contact areas temperatures. The delta-T for representative switches can easily reach 20 ◦ C - 30 ◦ C, and therefore the temperature effect on material properties should be considered. Timisit [46] outlines the rigorously valid Voltage-Temperature relation 2 − T12) V 2 = 4L(TM

(2.5)

where L is the Lorenz constant (2.45 x 10−8 V 2 K −2 ) and T1 is the bulk contact temperature. Equation (2.5) is independent of material properties and therefore will apply to both mono-metallic contacts as well as bimetallic contacts. The relation (TM − T1) is known as the contact supertemperature. While it takes additional time to evaluate the supertemperature directly from the voltage drop V , the increased accuracy of the equation is well worth the effort. Equation (2.5) can be re-written to solve for the contact temperature TM [45], 

TM =

2.3

T12 + V 2 /4L

(2.6)

Plating Material and Considerations There are many potential plating materials that could be used on copper bus

bars. The three most prevalent are Tin, Silver, and Nickel [12]. This paper will contrast and compare the relative benefits and drawbacks of these three coating materials as well as their various application options, as they relate to the successful operation of copper contacts. The following table outlines the material properties [12, 20] for the three platings considered

20

2.3.1

Property

Units

Copper

Nickel Silver Tin

ρ, electric resistivity

µΩcm

1.72

6.84

1.59

26.6[20]

δ, thermal conudctivity

W/m ◦K

380

70

418

-

H, microhardness

MPa

500

8000

700

E, elastic modulus

GPa

115

207

71

σt, tensile strength

MPa

250

317

125

14.5[20]

αt , temperature coefficient

10−3 ◦ C

3.97

6.90

4.01

-

αe , coefficient of thermal expansion

10−3 ◦ C

16.8

13.3

19.6

-

Tm , melting temperature

◦

C

1084

1455

1482

-

Ts , softening temperature

◦

C

190

520

180

-

Tb , boiling temperature

◦

C

2582

2837

2193

-

t, coating thickness

µm

-

3-5

3-5

-

Effect on Contact Resistance Plating of contact surfaces can improved or reduce constriction resistance de-

pending on material properties. A thin soft plating such as tin may allow increased effective contact area. Timisit [46] develops the equation of total spreading resistance from a coated surface by approximating the film resistance as ρf d/πa2 with plating thickness d and resistivity ρf and adding this to the general resistance formula to provide



Rt =

ρ 4a



 

4 1+ π

ρf ρ

 

d a

(2.7)

The total constriction resistance can therefore be reduced if the hardness of the plating layer improves the a-spot diameter while having sufficient conductivity and minimal thickness as not to cause a significant negative effect on the overall spreading resistance in the base material. Equation (2.7) can also be used to show when the resistance of the film thickness overshadows the constriction resistance. If (ρf /ρ)(d/a) is much larger than unity [46], the contribution from film resistivity will begin to dominate. This relation holds true not only for conducting metallic layers, but for insulative films layers as well. The beneficial effect of plating materials on copper joints has been validated

21

by research including Farahat [12] who found that both Silver and Nickel plated contact demonstrated both superior initial and long-term resistance over bare copper. His findings determined that Silver plating performed better than Nickel. Nickel is also stable up to 400 ◦ C [12] providing sufficient thermal stability for the application under discussion. Conversely, Tin softens very quickly at a relatively low temperature of 100 ◦ C [27, 36] making it undesirable for expected contact temperatures approaching 150 ◦ C. Silver was also found to operate reliably at high temperatures [36] Because of the high currents involved, the corresponding measurable reduction in voltage drop for plated contacts would result in significant supertemperature reductions as outlined in Equation (2.5). This difference could easily affect whether the contact temperature is approaching softening or melting regions. Plating copper bus bars also develops sub-surface inter-metallic layers of alloys consisting of copper and plating material. Braunovic [8] studied the effect of intermetallic layers developed with tin plated copper. He found that the micro-hardness of the intermetallic layer was much larger (345kg/mm2 ) than that of the tin plating (6kg/mm2 ) or the copper bulk material (100kg/mm2 ). The intermetallic layer was also very brittle, had a significantly higher resistance than the copper, and worse yet the intermetallic layer tended to measurably grow at elevated temperatures. Because of the brittleness of the intermetallic layer, initial contact resistance of tin plated contacts is higher than un-plated copper. Braunovic’s assessments were substantiated by Hamman [19] who adds that the intermetallic layers of tin are uneven and that the electrical properties of the intermetallic layers are similar to those of tin. Nickel plating also forms intermetallic layers with copper, and results in a relatively hard surface [3, 11]. Because of its hardness, it is a prime candidate for a successful bimetallic contact pair. An interesting finding by Timisit [44] warns that Nickel plating is subject to what he terms as ”Thermal Runaway” above contact voltages of 0.16V due to the temperature coefficient of resistivity of Nickel. This is much lower than the calculated melting contact voltage of 0.65 V. Silver plating has also shown relatively high hardness when compared to the copper base material

22

[11], and possesses excellent conductivity. 2.3.2

Effect on Oxidation, Fretting, and Corrosion Control Nickel plating provides improved protection against natural oxidation versus

copper because the oxide structures are denser for Nickel [42]. Chudnovsky studied the relative effects of copper plating materials and found electroless Nickel plating to be a good alternative to tin or silver plating [11] in industrial atmospheres. However, the plating parameters should be closely controlled to optimize corrosion protection, by limiting Phosphorous content to 2 atomic percent and ensuring thickness is at least .0005 inches for even mildly corrosive atmospheres [11]. Chudnovsky concluded after evaluation of corrosive specimens that a properly applied electroless Nickel plating would perform better than traditional silver and tin platings. Tin plating has a long and storied history with copper based contacts. Chudnovsky found that electroplated tin plating provided good corrosion control and relatively low resistivity [11]. Abbott [1] explains that while tin plating shows rapid surface film development up to 300 Angstroms, the relatively brittle oxide films developed are easily broken by loads greater than 100 grams due to the softness of the plating. The result is significant rupture of the oxide layer and reduced resistance. While this benefit would otherwise make tin a good plating candidate, the combination of low material hardness and high rate of oxidation can result in deleterious effects on contact that are subject to minute relative motion. One may consider the potential of enclosing the tin plated contacts in a sealed enclosure to minimize oxidation, however studies in this environment showed similar results as those that were un-sealed [43]. Fretting is defined as incidental relative motion between two contact pairs (typically less than 0.025 mm [1] ) which results in contact failure over a period of time. Because the rotary battery switch employs movable contact rotors and operates in a high vibration marine environment, the motion necessary to cause fretting should be expected. Malucci [31] develops the model of tin oxide formation and demonstrates how microscopically the contact region changes from metallic contact to a semi-conductive insulative film of tin oxide mixed with tin and finally

23

to a thick insulative film of tin oxide. This happens within only 1000 fretting cycles. While the obvious answer to this problem is to eliminate the possibility of fretting motions [32], fretting should be a significant consideration to any marine or vehicle application; and in this case tin plating is by far a poor performer. Silver plating has long been used for plating copper and other noble metals, mainly due to it’s hard surface and excellent conductivity. However, recent work has shown Silver to be very susceptible to hydrogen sulfide corrosion [11]. Un-plated copper raises significant concerns relative to oxidation, corrosion, and fretting as well. Gagnon [14] analyzed bare copper to copper fretting with a 400 gram load, 100 µm slip, and 50 mA direct current. His results showed steep resistance increases at about 50,000 fretting cycles, with continued resistance climbing thereafter. Bryant [10] also outlines this microscopic deleterious mechanism and defines the maximum effective contact separation before all conductive mechanisms are ineffective to be 20 µm. Gagnon and Bryant together outline the three stages of fretting as: • Stage 1 - Surface asperities on both contact penetrate natural oxide films establishing good metallic contact. While fretting will break apart these metallic bonds resulting in metallic particles and additional oxidation of fresh metal flakes and a-spots, continued motion will facilitate the escape of this ’wear debris’ resulting in a ’burnish’ of the surface and a relative stable contact. • Stage 2 - Continued fretting causes a build-up of wear debris abound the contact providing no place for debris to escape. The resulting growth of semiconductive material consisting of both contact material and oxide particles continues during this stage. As contact resistance increases, contact supertemperature correspondingly grows resulting is softening of the contact material and enlarging of the a-spot and thus cooling down of the contact. This stage continues with relatively flat, stable contact resistance. • Stage 3 - Continued fretting finally causes contact resistance break-down as the semi-conductive film thickness grows to a point which cannot be overcome through tunneling or other means until contact resistance peaks sharply

24

causing significant contact heat and effectively causing contact failure. 2.3.3

Effect on Sliding Contacts The ability for plating materials to withstand sliding motions under is impor-

tant when considering rotary switch applications. Additionally, a low coefficient of friction between contact is desirable in order to achieve maximum normal contact force while ensuring operation torque is within limits. Silver as a plating material is relatively hard, however, in order to provide sufficiently low friction and resistance to wear, it must be accompanied by a lubricant [5]. Chudnovsky [11] cites several disadvantages for sliding tin contacts including a lack of hardness to prevent galling, while lamenting on the benefits of electroless nickel in sliding contact applications. Indeed, Chudnovsky notes that not only does electroless nickel plating present high hardness to resist galling but a natural lubricity to minimize friction. Bryant [10] argues that the optimal sliding contact plating would consist of a hard nickel over a soft gold plate covering the copper substrate. Tin plating has shown to be soft, and subject to softening and galling at low and high temperatures [20, 36]. Overall, it is not recommended for movable contacts [20, 11, 31].

2.4

Film Resistance and Sources It is well known that bare copper oxidizes immediately in the presence of an

air atmosphere [22, 40, 24, 1, 30, 32]. This is a very important factor for rotary battery switches because product cost limitations do not allow for product to have their contact surfaces cleaned immediately before assembly, or to remove the air within the switches and replace ti with an inert gas such as Nitrogen. Consequently, we will investigate what types of surface films to expect and what levels of build-up we must overcome.

25

2.4.1

Natural Oxidation Oxidation on bare copper begins immediately after bare surface ions are ex-

posed to oxygen diffusion as noted above. Within 1 hour typical film thickness reaches 2 to 10 Angstroms; while over the next 2 to 3 days film thickness will reach 80 to 100 Angstroms [24]. This data was also supported by Malucci [32] who found copper surface films quickly reaching 140 Angstroms at elevated temperatures. It was also found that ambient temperature and humidity significantly affected the rate of oxidation, with rates as much as 50 percent higher under the existance of 80 percent relative humidty or 85 ◦ C [24, 30]. While the oxidation of copper contacts appears to continue at an increasingly slower and slower pace, it appeared that electron tunneling and fritting plays a part in reducing the overall affects of thin oxidation films [24]. Liu [29] also found fritting able to overcome copper surface films as thick as 55 nm with 2.0 Volts of contact potential. An additional method of film breakdown could also be mechanical, as Malucci [30] found that copper oxide films were very brittle and easily displaced. Oxidation is not limited to durations before contacts are mated, and in fact recent work by Sun [42] has shown that oxidation around contact spots can in fact be accelerated by the heat induced constriction resistance. Sun found that Copper Oxide films diffuse onto the contact by creating rings of surface film around the mettalic a-spots. Oxide films then diffuse through the existing oxide ring towards the center of the a-spot until the entire a-spot is covered with surface film. Sun foun that contact temperature accelerated contact film formation and postulates that thermal vibrations provide the energy for the atomic diffusion. Bryant [10] found similar results and cautions that such contact area surface film growth can ’wedge’ mated a-pots apart. Interestingly Sun [42] also found that increased compressive forces between contacts slowed oxide film growth measurably. Liquid and grease lubricants have been thought to improve corrosion resistance of bare copper by seperating the clean mettalic contact area from the oxygen in the air [17]. However, other evidence has shown that caution should be used when selecting lubricants for bare copper due to the the ability of copper to degrade polymers through catalytic enhanced oxidation [33]. An investigation by McCarthy

26

showed that currently accepted ASTM standards for grease aging were not consistent with the typically thin layers of grease applied to contacts [33]. McCarthy found that thin layers of grease were much more likely to oxidize when heated in the presence of a copper electrical conductor, including greases such a Nye 362. Lubricants with copper corrosion inhibitors appear to fair much better in this application [33]. 2.4.2

Corrosive Films Gaseous elements other than oxygen exist in certain areas relevant to marine

activity, including Sulfur dioxide, Chlorine, Ammonia, and very high humidity water vapor [41]. While one could argue the relative effect of these elements for a marine rotary battery switch, this paper will not include a detailed investigation into the ramifications they may have on contact design or performance. Because these products are assembled in water-tight, sealed plastic enclosures it is believed that the deleterious effects of these agents is relatively small.

2.5

Sliding / Wiping Contacts High current rotary switches are other time referred to as having ’sliding con-

tacts’. This characterization, however, is not technically correct. A ’sliding contact’ is technically a mated conducting pair that must hold specific resistance limits in both a static and dynamic moving condition. They are often continuously moving for long periods of time, or are required to operate during movement at such a low resistance in order to minimize electrically generated noise [17]. ’Wiping Contacts’ are not limited by such stringent requirements. They fall into two separate categories. The first are contacts which operate essentially as open/close and incur some small form of ’wipe’ as incidental to the action of opening and/or closing. The second form, which is the focus of this paper, involves contacts which operate through a deliberate sliding motion. However, these contacts are not required to operate at optimal efficiency during the operational movements because these movements are intermittent and related to operation of the switch.

27

2.5.1

Contact Wear, Friction, Shear, and Delamination Sliding between mated contact members is a significant area of concern for

high current rotary switches. The desire to reduce contact resistance through high normal force is counter-productive to creating a switch which provides thousands of repeatable operating cycles. If a low friction force between moving contacts is established and maintained through the life of the switch, a higher normal force (and thus lower contact resistance) can be maintained while keeping torque limits during hand operation within specifications. The three major sources of friction within mechanical systems are asperity interaction, adhesion, and plowing [35]. Moran [35] defines these three sources as follows: • Asperity Interaction - This friction element is caused by the asperity peaks on mated surfaces colliding resulting in a resistance to motion. As forces is increased, asperities are sheared-off or deform and eventually become loose wear particles. • Adhesion - Adhesion is the result of localized mechanicals welding at the microscopic interaction points. These welded junctions must be sheared before the surfaces can slide, the force required to shear these points results in friction • Plowing - Plowing is considered the largest contributor to mechanical friction. It is a result of wear particles becoming trapped between mated surfaces. Excessive loads are then concentrated on these wear particles, embedding these particles in or or both of the mated surfaces. When the surfaces attempt to move relative to each other, the wear particle cuts a grove in the surface(s). The energy required to ’plow’ the particle through a surface is a component of friction Moran points out that plowing friction is dependant on depth of asperity penetration and that consequently materials of significantly dissimilar hardness results is penetration mainly within the softer materials. Similarly, materials of similar hardness will result in equal penetration depths in both materials, and consequently a larger total penetration depth will be seen resulting in a larger friction forces.

28

There are several contact and system design features which could reduce the deleterious effects of wear, including the formation of wear particle traps through surface grooves or undulating surfaces [35], selecting mating materials of significantly differing hardnesses [35, 10], as well as the use of lubricants as will be discussed later. However, one negative aspect of producing particle traps is the increased manufacturing expense that would result. Additional insight provided by Glossbrenner [17] indicates that a measurable fraction of the contact load is supported by contact films as well as metallic contact. This would make sense given the natural oxidation process that the bare or plated/lubricated contact undergoes. As a result, many of the asperities removed during sliding motion will result in insulative films trapped in contact grooves. With repetitive sliding motions, Glossbrenner also found good evidence that metallic conduction was likely to occur between ’matched’ a-spots from previous matings. This was proven by examination of the continuous resistance ’signature’ of a rotating contact pair and the repetitive nature of the contact resistance at each angular rotation point. Therefore, caution should be take to avoid significant surface film creation between mating cycles which could increase contact resistance. However, research has found that any wipe greater than .025 mm effectively removed surface films and reduced contact resistance significantly [30, 32, 1, 27]. Malucci [32] investigated an extremely relevant case to this paper of a rounded copper probe on a flat copper plate, and found contact wipe to produce up to two orders of magnitude lower resistance than without wipe. Figure 2.2 taken from Malucci [32] shows how contact wipe produced significantly better results regardless of contact force; and contact force greater than 100 grams resulted in no resistance reduction when wipe was used. This contrasted the improved contact resistance without wipe as contact force increased to 500 grams. 2.5.2

Boundary Lubrication Lubrication of sliding contacts can have both positive and negative effects de-

pending on application and environmental variables. If the localized environment within the switch is not free of loose debris, it can easily be attracted and retained

29

Figure 2.2: Contact Resistance vs. Force by a grease or liquid lubricant [35]. This could easily result in particle entrapment and increased friction, wear, and resistance. In fact, depending on lubricant properties and contact interface geometry, boundary lubricants are typically represent the largest quantity of materials which sliding contacts must remove in order to establish effective electrical connections [17]. Lubrication is widely used in sliding contacts in an effort to either reduce electrical contact resistance, limit sliding wear, or both. While Holm [22] believed that boundary lubricants posed no greater threat to increase constriction resistance than physisorbed films such as CuO, Klungtvedt [23] reminds us that grease lubricants pose an incremental risk due to the grease matrix which holds the lubricant. He identifies the critical negative aspects of silicate based lubricants being the hard glass-like silica particles which can easily become trapped between contacts and because of their hardness and resistance to breakage can cause significant contact damage through plowing [23]. Additionally, silicate particles were shown to cause problems in any potential arcing environments as they carbonize and fuse themselves to contacts in the presence of a drawn arc [23]. Furthermore, Klungtvedt postulates that any refractory material in the area of an arc would exhibit the same splattering and fusing effect. Conversely, Klungtvedt’s analysis found that lithium based lubricants were found to significantly improve contact life, which he postulated was a result of the soap-like consistency of the particles which are easy to crush and move out of the

30

immediate contact region [23]. When chosen properly, lubricants can serve to effectively shield contact areas from natural oxidation by forming a protective barrier around the contact region [17]. Contact geometry, base material properties, and contact force, all play a role in determining the proper grease viscosity. A properly greased contact will act anaerobically initially, however, after repeated sliding operations, the eventually removal of sufficient lubricant will result with an aerobic contact condition. This will provide an avenue for oxidation films to be re-formed over time. For this reason, contact lubricants are only believed to delay the onset of increased contact resistance, as has been proven by Leung and Glossbrenner [27, 17]. Consequently, a variety of research has found that regular wiping action during contact make and break was essential to maintain a minimum contact resistance [23, 43]. The list of possible lubricants for the application under discussion is very long. Recent work by Gagnon [15] with perfluoropolyether (PFPE) oil based lubricants indicated they were relatively ineffective at low loads (100 grams), but provided measurable benefit at higher normal forces (500 grams). Gagnon’s study found no benefit from lubricants with metallic additives, but he postulates that they may have benefit if able to improve the penetration of the oxide layers on the contact surfaces [15]. His findings for PFPE based lubricants reccomended Superlube or Uniflor 8622. Finally, an innovative solution which utilizes silver iodide as a solid lubricant was proposed by Arnell [5]. His investigations showed that silver, like many other soft, high conductivity plating, do not have significant wear resistance. By applying silver iodide on contact regions, Arnell was able to significantly reduce contact wear while not increasing contact resistance. Even with this finding, however, Arnell found that the use of a perfluorinated grease improved the performance of the solid lubricant in open air conditions.

2.6

Summary of Findings The proceeding discussion outlines the state of art in current thinking relative

to copper to copper contacts used in high amperage sliding switches. From this

31

investigation, the following conclusions are reccomended when considering design variable for future products of this nature. Experiments will be made to attempt to qualify the conclusions drawn. • Surface Geometry - Contact force should be sufficient to establish superior resistance and remove oxidation films without causing unnecessary wear. Macroscopic a-spot locations should be evenly distributed to maximize the effective Holm radius. • Plating - A hard electroless nickel plating one one contact member should be superior to two copper contacts. Tin plating is not reccomended because fretting vibrations are unavoidable and removing air from contact regin is cost prohibitive. Silver plating is cost-prohibitive for this application, and the poor wear properties would present increased lifetime concerns. • Lubrication - No lubrication may be possible if contact force is held sufficiently low (not considering potential thermal benefits). Boundary lubricants, if used, should be selected so viscocity matches contact force, temperature, and geometry conditions. • Sliding Contacts - Working-in of contacts after assembly of switch is reccomended to ensure contact pairings are allowed to remove initial surface films on copper parts resulting from time between fabrication and assembly.

CHAPTER 3 CU to CU CONTACT THERMAL RESISTANCE OPTIMIZATION A discussion of the thermal contact to contact interface is necessary in order to effectively evaluate and estimate contact performance. It is a common misconception that the electrical resistance of a contact is the same as the thermal resistance. While there are many similarities between the equations and dynamics of the two concepts, there is a significatn difference between the two. While the electrical constriction resistance is derrived mainly from the constriction of electrical lines of flow from an even distribution in the main cunductor to the small a-spots at the contact interface, thermal constriction resistance is base partially on th electrical constriction resistance but also includes the thermal conduction paths that are not electrically viable. These thermal paths include conduction through electically insulative surface films as well as the often used contact lubricant material This chapter will investigate the currently accepted knowledgte regarding contact thermal resistance and conductance. Areas whihc will be covered include the effects of contact geometry, lubricants, surface roughness, copper plating, and rotary contact break-in.

3.1

Contact Geometry and a-spot Optimization Like many aspects of contact design which affect contact thermal resistance,

contact geometry can not be examined by itself but rather by examining the combined effects of the contact geometry with other aspects such as lubrication or surface roughness. In order to achieve a minimum of gap thickness for the largest possible thermal contact spot area, the contact pairs should be a flat and smooth as possible while allowing for proper operation and motion. The radius of lead approach for the movable contact should be as large as possible and the nominal flat contact area should be sufficiently large enough to provide two flat surfaces capable of supporting improved thermal contact conduction through both mettalic and semi-insulative 32

33

thermal conduction. Recent studies by Leung and Yovanovich [28, 48] have shown that while voids between mettalic contact a-spots do not contribute to electrical conductivity, they do contribute to thermal conducticity, and that heat transfer accross mated contacts is a result of two parallel thermal paths. Leung explains that the two effective paths are the solid to solid conduction accross the ”a-spots” as well as conduction conducted through the fluid trapped at the microscopic surface voids [28]. Leung defines the thermal contact model as metaphorically two parallel conductors and presents the equation for the total thermal contact resistance as: Rtherm =

1 hc A



1 Ac Av 2ka kb + hc = ∗ ∗ kf Lg A ka + kb A where: hc

(3.1) 

(3.2)

= coefficient of conductance

Lg

= thickness of void space

ka

= thermal conductivity of material a

kb

= thermal conductivity of material b

kf

= thermal conductivity of fluid in void

A = apparent contact area Ac

= solid to solid contact area

Av = void area Yovanovich himself explains that while surface roughness and aesperity slope effect micro-thermal joint resistance, relative flatness between contact members dominates the macro-thermal joint resistance [48]. Yovanovich develops his own model for the gap resistance [48] with the resulting equation: Rgap =

d1 + d2 2πKf A2ln(A/Ac )

(3.3)

where: d1 , d2 = flatness of contact members which must be added to the spreading resistance in order to determine the thermal joint resistance by the equation:

34

1 Rtherm

=

1 1 + Rs Rgap

(3.4)

Because of rotary contact break-in expectations, contact geometry can also be used to develop significant reduction in thermal constriction resistance by designing evenly spread-out contact points design to ensure that primary contact points meet before incident contact points. This effectively reduces the electrical spreading resistance as was discussed in Chapter 2 and well documented by Timisit [?, 44]. By designing a small gap distance between the bulk contact material surfaces and the primary mettalic contact points, additional thermal resistance reduction can be achieve through lubricants or surface films between the bulk material surfaces.

3.2

Lubricant Effects Lubrication can play a significant role when evaluating the thermal resistance

of a contact. Lubricants applied to the contact mating surfaces and surrounding the adjacent conducting pairs will fill the microscopic peaks and valleys between mating contact mettalic material. In contrast to the relatively insulative air gas that would otherwise fill these valleys, lubricants can provide a significantly improved conduction path between contact pairs. There are many lubricants which advertise significant thermal heat transfer results through the use of mettalic filler particles suspended in the lubricant substrate. The majority of these products, however, have shown little improved advantage over lubricants developed to improve wear properties. One investigation into the mechanism which results in improved thermal properties with the use of mettalic infused lubricants suggest that the mettalic particles act to decrease the required a-fritting voltage accross a seperated pair of contact peaks. This would result in the development of a great number of additional electrical connections via electron tunneling and subsequently create permanent dendritic growth connections between the mated contact pairs.

35

3.3

Surface Roughness Surface roughness has been shown to measurably affect the thermal resistance

of mated contacts. When two contacts are mated, the surface roughness of each respective material surface dictates the minimum, maximum, and average gap distance between adjoining contact pairs. If the two contact surfaces have relatively small roughness, the average air gap between the contacts will be smaller. While this would marginally help contact pairs without lubrication, the addition of contact lubriaction to fill these smaller valley distances means that a thinner thermal conduction resistance will exist through the lubricant filler.

3.4

Plating Effects Surface plating of one or both contact pair members can affect thermal re-

sistance between the two contacts. Because of the difference in material hardness of various platings versus the copper base material under investigation, a harder plating material such as electroless nickel for example on only one contact would act to improve electrical and thermal resistance by forcing the bare copper contact to conform to the surface features of the nickle plated contact. This plowing effect will act to reduce average contact gap size and improve contact resistance in a similar manner to the effect gained by reducing surface roughness. Another way that surface plating can act to affect thermal resistance is through the oxidation affect that the plating material attracts. While a copper material oxidizes relatively quickly, the copper oxide film presents a natural filler material between mettalis contacts which provides some form of thermal conduction across contact pairs. Tin plating, however, oxidizes into a relatively insulative film material. Besides the well-documented deleterious affects tim-oxide has on the electrical contact resistance, tin oxidation films will not provide a beneficial thermal conduction path for contact valley gaps.

3.5

Rotary Contact Break-In As mentioned earlier, the operation of wiping contact pairs will act to provide

some benefit to reduce thermal contact resistance. Dissimilar materials will wear

36

together in such a way that the harder material will plow into the softer material. This action can effectively reduce the average contact to contact valley gap space by creating defined wear lines in the softer material and removing the softer matierla in a more orderly manner. Conversely, when two similar copper contacts are mated, they will act to ’plow’ evenly between each member. Friction effection such as mechanical adhesion and asperity deflections are more likely to cause transverse shearing of both materials in a random manner during operation of the contacts. With continual interaction and plowing of both materials, mettalic debris will also become lodged between contact memebrs and re-adhere to one of the contacts. This process will not result in a closely mated pair of contacts with small valley gap distances, rather it would result in roughened conect grooves and increased surface roughness and waviness.

CHAPTER 4 THERMAL / ELECTRICAL MODEL OF ROTARY DISCONNECT SWITCH 4.1

Assumptions / Fixed Parameters In order to allow a flexible thermal/electrical model to be developed which

takes into account the key parameters of interest, it was necessary to set fixed several problem parameters. The following list of system parameters were held fixed to the identified reference: - Conductor materials will all be pure electrolytic copper. - Contact geometry will consist of a semi-circular pin shape on a flat surface. The radius of the semi-circle will remain a variable for the overall solution to take into account. - Thermally Insulated Current Path Inside Switch - Sliding wear due to rotating rotor will not be considered. Only newly assembled switches will be considered. - Due to switch assembly ans terminal assembly practices and previous empiracal measurements, cable terminal connection resistance and resistance from switch terminal to internal bus bars will be considered negligle and ignored. - External power cables will convect and radiate heat to the surrounding area via natural convection and radiation heat transfer.

4.2

Variable Parameters The following variables will be input into the model and are required to solve

the system. - Contact force; which will be derrived directly from the internal spring force divided by the number of contact pairs which in this case is two. - Contact Geometry, including number of contact pairs and radius of pin contact

37

38

- Rotor cross-sectional profile and length between rotor contacts. - Connecting power cable; conductor size and diameter, insulation properties and thickness. - Plating material, if any; physical, electrical, amd thermal properties. - Contact lubricant, if any; properties including viscosity and thermal impedance.

4.3

Output Parameters The model will estimate several quantifiable and relevant output parameters,

including the electrical contact resistance, the contact super temperature, the temperature profile along each power feed cable, and the temperature profile along the internal copper rotor

4.4

Thermal Effects of Contact Resistance In order to accurately estimate the output parameters of the thermal/electrical

model, material properties and other model aspects which depend on temperature will be varied accordingly. For example, the temperature coefficient of restivity will be used to more accurately estimate the conductor resistance and the constriction resistance. Additionally, the more complicated variation of the super-temperature calculation will be used because of the expected large difference between bulk material temperature and contact temperature. One aspect that will not be considered is the potential effect of softening or melting of materials. It will first be assumed that material phases remain largely the same with similar mechanical properties. If after evaluation of contact resistance and voltage drop it is found that softening and/or melting voltages were approached, the results will highlight the potential erroneous results.

4.5

Steady-State Equations

4.5.1

Power Connection Cables For the probem being discussed, the expected dominant steady-state heat

transfer mechanism to remove the heat generated by the two contacts are the two

39

power feed cables. Heat energy created by the mated contact resistance and the continuous current will be conducted to the power terminals and then conducted to the power feed cables. The power feed cables will act as linear heat sinks dissipating the heat most effectively along the sections closest to the switch through convective and radiative transfer between the outisde cable surface and the ambient enviornment. Even though the cables themselves will be creating heat energy, the temperature difference between the current induced thermals and the a-spot temperature is sufficient to provide an excellent thermal path. This natural deduction was verified by Filippakou [13] who found that power delivery conductors and bus bars serve as contact coolers. His findings also concluded that increases in contact temperatures affected feeder conductor temperature as far as 50 cm away from the contact pair. Furthermore, Filippakou highlights that as convective cooling air currents around the feeder conductors approaches natural convection, the cross-section of feeder conductor play an increasing role in contact cooling. Finite section of cable Energy Balance, steady-state dT /dt = 0, uniform copper wire temperature at node i, Qi−1 + Qgen = Qi+1 + Qout

(4.1)

Qgen = I 2Ri

(4.2)



Qout

kins Ains = Ti − Tiwall dins









4 4 − Tamb Qout = hAwire Tiwall − Tamb + Awire Tiwall



Qi−1

kcu Acu = Ti−1 − Ti ∆x

Qi+1

kcu Acu = Ti − Ti+1 ∆x



(4.3) 

(4.4)



(4.5) 

(4.6)

40

where: Qi−1

= energy from node i-1 into node i

Qi+1

= energy from node i into node i+1

Qgen

= energy generated in node i

Qout

= energy loss along node i to ambient surroundings

I Ri

= electrical current passing through cable = finite resistance of node i

kins

= thermal conductivity of wire insulation

kcu

= thermal conductivity of copper at current node temperature

Ti

= conductor temperature at node i

Tiwall

= temperature of outside insulation at node i

Tamb

= ambient temperature

Ti−1

= conductor temperature at node i-1

Ti+1

= conductor temperature at node i+1

dins

= conductor insulation thickness

Ains

= average insulation area for heat loss

Awire

= insulation surface area for element i

Acu

= copper conductor cross-sectional area

∆x = distance from node to node  = emissivity * Stefan-Boltzmann Constant * shape factor

4.5.2

Internal Contact Pairs

4.5.2.1

Electrical Contact Resistance and Super Temperature Calculations

4.5.2.2 4.5.3

Thermal Contact Resistance Copper Rotor Conductor

The internal copper rotor will be heated by contacts on both ends and in general is insulated from dissipating heat because it resides inside of a plastic insulated enclosure. The main mode of heat rejection for the copper rotor is through the two contacts which are the primary cause of the heat in the first place. As a result, the heat flux at the contact interface is not zero rather at steady-state conditions the heat flux at the contact will be flowing away from the rotor and into the conducting

41

cables. However, it is a desire to ensure completeness and therefore the potential heat loss directly from the rotor itsef will be considered. The potential heat loss methods beyond conduction through the contacts are convective and radiative parallel paths between the rotor and the inside surface of the insulative housing, in series with the thermal conduction through the housing followed by a second set of parallel paths from the outside. The energy balance equations therefore are: Qi−1 + Qgen = Qi+1 + Qout

(4.7)

Qgen = I 2Ri

(4.8)







4 Qout = hAirotor Ti − Tins−in + Airotor Ti4 − Tins−in



Qout =

khousing Ahousing Tins−in − Tins−out dhousing











4 4 − Tamb Qout = hAins−out Tins−out − Tamb + Ains−out Tins−out



Qi−1 =

kcu Acu Ti−1 − Ti ∆x 

Qi+1 where: khousing Ti

kcu Acu = Ti − Ti+1 ∆x

(4.10) 

(4.11)



(4.12) 

= thermal conductivity of switch housing = conductor temperature at node i

Tins−in

= temperature of inside housing insulation

Tins−out

= temperature of outside housing insulation

dhousing

= housing thickness

Ains−out

= external housing area

Ahousing

= average housing area for heat loss

Airotor

(4.9)

= rotor surface area for element i

(4.13)

42

4.6

Finite Difference Model

4.7

Results of Calculations The thermal model was utilized to estimate the

CHAPTER 5 EXPERIMENTAL VALIDATION OF ELECTRICAL/THERMAL MODEL In order to assess the viability of the electro-thermal system model developed in the previous chapter, a set of empiracal tests were conducted on a respresentative marine rotary battery switch. A fixed set of variable parameters were adjusted and controlled in order to provide increased information about the affect of each parameter as well as to measure the ability of the finite difference model to take each parameter into affect with a correct weighting factor.

5.1

Experimental Set-Up and Procedure All tests were conducted with Blue Sea Systems model 9003E rotary battery

switches. These switches provide a simple On/Off operation by rotating in a 90 degree motion, with subsequent operational motions from Off to On returning over the same path as On to Off. All switch variants were constructed by the same personnel and assembled to the same manufacturing standards. 5.1.1

Tests Conducted Each switch was taken through the same test and operational sequence as

indicated below: • Initial Switch Resistance - The switch was assembled and operated exactly one (1) time from Off to On to Off. It is then placed ’On’ and assembled into the test fixture to determine the switch resistance and terminal temperatures after one (1) hour passing the rated current. • Phase I Rotary Contact Cycling - The switch was then operated mechanically by hand at a rate of 360 operations per hour (one every 10 seconds) for 250 cycles. Following this cycling, the switch was again assembled into the test

43

44

fixture to determine the switch resistance and terminal temperatures after one (1) hour passing the rated current. • Phase II Rotary Contact Cycling - The switch was then operated mechanically by hand at a rate of 360 operations per hour (one every 10 seconds) for 1750 cycles. Following this cycling, the switch was again assembled into the test fixture to determine the switch resistance and terminal temperatures after one (1) hour passing the rated current. • Dis-assembly and Visual Inspection - Lastly, the switch was dis-assembled and visually examined to determine the state of contact pairs and any deleterious affects on other switch components. 5.1.2

Measurement of Switch Resistance The battery switch test fixture shown in Figure xxx was built to test the

battery switch in the same manner consistently for all tests. It included a Xantrex Model xxxx power supply, two four foot power cables, two 1 foot voltage leads design to assemble to the switch terminals, and a Fluke Model xxxx multi-meter. Using flat and lock washers and controlling the tightening torque for each switch assembly to the fixture, resistance variation at the cable connection points was minimized. Voltage drop probes were connected between the flat washers and the cable terminals to ensure repeatablility. Before testing, the probes were calibrated and zeroed by attaching the probe ends to each other.

*** Show figure of Test

Equipment and Switch Assembled *** A continuous current of 300 Amperes was used to ensure sufficient accuracy of the voltage drop measurements and closely approach the operating limits of the devices. Each test was conducted for one hour to allow contact supertemperatures and all other conductors to thermally stabilize and ensure contact resistance values were representative of a continuous current situation. The voltage drop probes were measured at the end of each hour long test. Equation (5.1) below was used to convert the voltage differential at the contact points to resistance in Ohms from each point. R = V/I = Vmeas /300A

(5.1)

45

The following resistance calculations were made to estimate the electrical resistance of the bulk copper material between the two voltage drop measurement points. R-Bus-bar = (x2) *** R-Bus-bar to terminal = (x2) *** R-Terminal = (x2) *** R-Rotor = (x1) *** Total Bulk Material Resistance (Calculated) = ??? *** Finally, the bulk material resistance value was subtracted from the measured results to determine the actual contact to contact resistances. Because each switch has two contact pairs of contacts, the average contact resistance for each switch was calculated by dividing the previous result by a factor of two.

5.2

Measurement of Internal and External Conductor Temperatures Figure xxx shows the three locations used to consistently measure the conduc-

tor temperatures at the end of each one hour continuous operational test. Points A and B refer to the ends of each threaded terminal post. As outlined in Chapter 1, each power post is a single-piece forged copper component that carries the current from the power ring terminals through the switch housing and into the fixed bus bars. Because this component is a single piece of solid copper, the temperature variation of the power post is assumed to be negligible and therefore the temperature reading at the end of the threaded post will represent the entire part. The third temperature probe location is identified in Figure xxx as point C. For each battery switch, a hole was drilled at the same point; which provides external access to the middle of the movable copper rotor when the switch is in the ’On’ position. At the end of the one hour operational time duration, each temperature reading point was probed until readings stabilized. This allowed accurate switch conductor measurements while not affecting the integrity of the test results with permanently attached thermocouples. *** Show picture of bottom of battery switch, outlining the three temperature reading locations ***

46

5.3

Conducted Experiments and Results The design variations between each sample were chosen to provide sufficient

variation in construction parameters so that individual aspects of the numerical estimation model could be seperately evaluated for corectness. Additionally, a large enough quantity of different samples were chosen so that comparison of deviation could be performed against the same parameter for many samples. The design variations were also chosen to evaluate the qualitative results from Chapter 2.The three design variation areas were as follows: • Contact Force - The compression spring which internally generates force between the mated contacts defines the contact force. There are two sets of contacts inside the switch due to the rotatble internal rotor and two fixed bus bars. The variation of contact force will range from 3 lbf to 25 lbf. While the results should show reduced contact resistance with larger spring force, the wear affects of the internal conductors after 2,000 cycles is of particular interest • Lubrication - The contact lubricant for the samples will include Uniflour 8512, Nygel 759G, Nygel 761G, and no lubrication. The Uniflour product is designed to operate at very high temperatures (175 ◦ C) and also has a resistance to carbonization in the presence of an arc. While the high operating temperature is desired, the grease matrix particles are known to embed themselves in porous metals such as copper and increase resistance accordingly. The Nygle 759G is a proven lubricant, however it have a relatively low operating temperature of 125 ◦ C. The Nygel 761G lubricant is an experimental hybrid product which is touted to proved the excellent contact properties of the 759G but with a much higher temperature rating (150 ◦ C − 160 ◦ C. Finally, the option of using no lubrication will also be investigated. Particular attention will be paid to the rate of switch cycling, since un-lubricated contacts should wear faster when operated more frequently because of the friction induced heat. • Plating - Nickel and Silver plating were considered for this evaluation, as well as using bare unplated copper specimens. Combinations of plated bus bars

47

with un-plated rotors and plated rotors with plated bus bars were all tested with and without a grease lubricant to ensure lubrication would not affect the validity of the results variation. With the above design variations decided upon, Table 5.1 outlines each test sample variation of construction. The remaining construction paramenters for all samples were identical. Table 5.1: Construction Parameters for Samples Sample Spring Number Force Lubricant 1 3 lbf 8512 2 17 lbf 8512 3 25 lbf 8512 4 31 lbf 8512 5 17 lbf None 6 17 lbf 759G 7 17 lbf 761G 8 17 lbf 8512 9 17 lbf 8512 10 17 lbf None 11 17 lbf None 12 17 lbf 8512 13 17 lbf 8512 14 17 lbf None 15 17 lbf None

5.4

5.4.1

Rotor Bus Bar Plating Plating None None None None None None None None None None None None None None None Nickel Nickel Nickel None Nickel Nickel Nickel None Silver Silver Silver None Silver Silver Silver

Results and Conclusions

Spring Force Variations *** Graph of Contact Resistance vs. Number of cycles, four curves for each

spring force*** *** Internal contact area photgraphs if informative

48

5.4.2

Lubricant Variations *** Graph of Contact Resistance vs. Number of cycles, four curves for each

lubricant with bare copper contacts*** *** Internal contact area photgraphs if informative 5.4.3

Plating Variations *** Graph of Contact Resistance vs. Number of cycles, four curves for Nickel

plated contacts with and without lubrication*** *** Graph of Contact Resistance vs. Number of cycles, four curves for Silver plated contacts with and without lubrication*** *** Internal contact area photgraphs if informative

CHAPTER 6 TABLE OF SYMBOLS Symbol ρ ρf Rc a Ac F H σy L T1 TM V d

Definition Bulk Material Restivity Plating Resistivity Electrical Resistance From Current Constriction Radius of a Single a-Spot Total Contact Area Total Normal Contact Force Material Hardness Yield Stess Lorenz Constant Bulk Temperature of Contact Material Absolute Temperature of Contact a-Spot Voltage Drop Accross Contact Plating Thickness

49

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