Effects of temperature and span amplitude on fretting corrosion behavior of tin-plated electrical contacts

Effects of temperature and span amplitude on fretting corrosion behavior of tin-plated electrical contacts

Microelectronics Reliability 69 (2017) 80–87 Contents lists available at ScienceDirect Microelectronics Reliability journal homepage: www.elsevier.c...

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Microelectronics Reliability 69 (2017) 80–87

Contents lists available at ScienceDirect

Microelectronics Reliability journal homepage: www.elsevier.com/locate/microrel

Effects of temperature and span amplitude on fretting corrosion behavior of tin-plated electrical contacts Min-Jung Kim a, Ho-Kyung Kim b,⁎ a b

Graduate School of NID Fusion Technology, Seoul National University of Science and Technology, Republic of Korea Dept. of Mechanical and Automotive Engineering, Seoul National University of Science and Technology, Seoul, 139-743, Republic of Korea

a r t i c l e

i n f o

Article history: Received 4 October 2016 Received in revised form 23 December 2016 Accepted 23 December 2016 Available online 8 January 2017 Keywords: Fretting corrosion Electric contact Failure lifetime

a b s t r a c t Fretting tests on electric contacts were conducted to investigate the effects of temperature and contact span amplitude on fretting corrosion behavior at 298 K, 323 K, 348 K and 373 K under a constant force of 0.85 N. Riders and flats made of 0.3 mm-thick brass sheet were coated with 10 μm of tin. The electric resistance was measured during the fretting test period. The electric resistance was very low; afterward, however, the resistance increased rapidly and intermittently with the number of fretting cycles. Electric failure lifetime (Nf) was found to decrease with increase in the testing temperature. It was assumed the failure cycle is a cycle with an electric resistance of _ due to the degrading of the contact was inverse-linearly proportional to the number 0.01 Ω. The damage rate (D) of cycles. A formula for lifetime prediction was developed by applying the Arrhenius equation. Using this formula, the lifetime can be predicted within a factor of two in a range below 35 μm span amplitude. © 2016 Elsevier Ltd. All rights reserved.

1. Introduction With increasing application of various advanced electronic control systems in vehicles, the use of connectors to connect those electronic components has been steadily increasing in recent years. Field data has shown that a large portion of the electrical failures of electronic systems can be attributed to connector problems. Primarily, electrical failures are caused by high contact resistance or the intermittence of electrical signals. Even a slight contact failure can have severe impacts on the vehicle's safety and operation. Therefore, the reliability of electronic systems in vehicles during long-term operation has become a very important issue. Contact failures are caused by diverse factors, and by the interactions between those factors, including chemical corrosion, fretting corrosion, inflow of dust or other foreign matter to contact points, wear, etc. [1]. Fretting corrosion, which is one of the major types of connector contact failure, refers to a type of corrosion caused by the micro relative motion of the contact points of the connectors; this corrosion produces fretting wear. In the case of an engine speed sensor connector, for example, vehicle vibrations as well as temperature variations due to stopping and restarting of an engine can lead to micro movement and wear of a connector contact point. In addition, higher temperatures around the engine accelerate the creation of an oxidation film over the contact point, leading to higher electrical contact resistance. Eventually, the

⁎ Corresponding author. E-mail address: [email protected] (H.-K. Kim).

http://dx.doi.org/10.1016/j.microrel.2016.12.014 0026-2714/© 2016 Elsevier Ltd. All rights reserved.

signal value sent by the sensor can be distorted, causing the vehicle ECU to send an output signal with the wrong values. In such a case, the operation of an actuator can be affected, resulting in fatal errors. Major factors involved in fretting corrosion include current magnitude, contact force, relative humidity, frequency, span amplitude, temperature, environment gas, etc. [1,2]. Many studies have been performed to understand the factors contributing to fretting corrosion of connectors [2–5]. For example, Ito et al. [2] evaluated several control factors to determine their effects on fretting corrosion of tin plated contacts. They found that contact force and tin thickness have higher impact on contact resistance than do other control factors such as temperature, humidity, frequency and amplitude. Park and Lee [3] conducted fretting corrosion experiments using tin-plated brass specimen, and proposed two three-dimensional empirical equations by considering corrosion and wear accelerator. Liskiewicz et al. [4] investigated the impact of corrosion on the fretting damage of electrical contacts. They reported that a larger span amplitude leads to reduction of the unexposed surface area and, as a result, shorter lifetimes for electrical contact. Jedrzejczyk et al. [5] conducted research on the effects of temperature and normal force on the transition amplitude from the partial slip to the gross slip regime which yields a threshold between the theoretical infinite and the actual finite lifetime of electrical contacts. They proposed a fast method to quantify the transition amplitude of an electrical contact by adopting a conventional constant-span-amplitude test methodology. Hannel et al. [6] proposed fretting sliding transition amplitude as a criterion for electrical contact performance. They suggested that the sliding transition amplitude between the stabilized partial slip condition and the gross slip condition can be predicted analytically or

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by FEM analysis. However, there have been a limited number of studies on the effects of temperature on the lifetime of the electrical contacts. This study aims to analyze the effects of temperatures and span amplitudes on a tin-plated copper alloy contact resistance under a fixed contact force. Finally, in order to provide basic data necessary for design of connects, this study will formalize an equation for predicting the lifetime as a function of temperatures and span amplitudes. 2. Experimental procedures 2.1. Test apparatus For the fretting corrosion experiment, a system for controlling and measuring the displacement and contact force has been designed and manufactured. The design has been specifically constructed to allow the application of constant force onto a specimen. In Fig. 1, it can be seen that the experimental equipment includes: a loading device for generating specimen load; system for measuring load and displacement; load train system for applying load onto a specimen; grip and jig for fixing and supporting a specimen; and system base. In addition, since the specimen is subminiature in size, an X-Y-Z stage for adjusting the locations of the specimens is installed. The signals of the horizontal and vertical directions detected from the load-cell are amplified using a dynamic amplifier (Vishay Micro-Measurements, 2300 Signal Conditioning Amplifier). To measure the relative displacement of the two contact points, a non-contact CCD camera with 0.5 μm resolution is used. The displacement and load signals are recorded using a data acquisition device (National Instrument, DAQ Card-6024). As for the software for data acquisition and control, ver. 8.0 LabVIEW is used. 2.2. Specimen preparation and test condition The electrical-contact specimen used in the experiments is a tinplated brass sheet of 0.3 mm in thickness (C2600, Ni: 1.82%, Si: 0,75%, Zn: 0.01%, Sn: 0.37%, Cu: balance). To improve the adhesive strength, after thin-coating the sheet with copper, specimen was electrically coated with tin with a thickness of 10 μm. In order to form a single contact point, the upper specimen was produced as a hemispherical specimen with a curvature radius of 1.04 mm; the lower specimen was made as a flat specimen. The upper specimen was fixed to the upper moving table while the lower specimen was fixed to the lower holder. The specimen is placed in a small chamber; the experiments were conducted at 298 K, 323 K, 348 K, and 373 K. To eliminate displacement errors due to temperature changes caused by the ambient temperature,

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fretting experiments were performed by isolating the whole testing device within a chamber. The span amplitudes of the specimen were 25, 28.5, 30, 32, 34, 50, and 77 μm. Three specimens were used for each test. Control of the span amplitude was realized by applying the image analysis method of the camera on the basis of knife edge gaps for the measurement of gage length closest to the pair of contact points. A frequency of approximately 0.003 Hz was applied to minimize the impact of frictional heating. The experiments were conducted for up to 3000 cycles. Scanning electron microscope (SEM) analysis of the specimen was performed to analyze the degree of surface damage due to fretting wear. To measure changes in the electric contact resistance during fretting testing, the four-wire resistance measurement method was applied. Contact resistance readings were taken every cycle at setup measurement points along the contact area with a constant current of 0.1 A, as shown in Fig.2; the readings were recorded on a computer. A highresolution digital multi-meter (Agilent 34410A) capable of measuring up to 1 nV was used for voltage measurement. 3. Results and discussion 3.1. Change of electric contact resistance In this study, changes of the electric contact resistance have been observed by applying different span amplitudes with contact force of 0.85 N and temperature of 298 K, 323 K, 348 K and 373 K. From the results, it has been found that as the temperature increases, the contact resistance increases rapidly; the pattern of resistance change was similar for most experiment conditions. Fig. 3(a), (b), (c) and (d) show the change of electric resistance at each temperature for an amplitude of 34 μm. The contact resistance increases for a few initial cycles, but drops rapidly. The initial increase might be due to the oxide layer, which formed on the additional coating layer, but the resistance soon drops as partial contacts are made between tin and tin. Fig. 3(a) that the contact resistance is maintained in a stable manner approximately the 80th cycle, but the resistance continues to move up afterwards. As has been explained in other studies [7,8], this resistance increase is presumed to be caused by wear debris and oxidation products accumulated due to the wear of the oxide and tin layers, as the particles build up and then block the contact parts of the surface. Around the 225th cycle, an electric resistance peak of 0.01 Ω or higher is reached for the first time; then changes take place unevenly. This lack of consistency is due to a decrease in the conductive contact area caused by the build-up of wear debris and oxide particles accumulated on the matched surface. 3.2. Structural analysis on the specimen When normal load is applied between the bump and plate of a contact specimen, deformation around the contact point occurs. Threedimensional structural analysis was performed to determine the contact stress of the contact specimen. In order to obtain tensile properties for the structural analysis, a miniaturized tensile specimen was extracted from the 0.3 mm-annealed brass plate and the 1 mm-thick tin plate. The specimen had a gage width of 4 mm, gage length of 8 mm, shoulder

Fig. 1. Fretting test assembly: dead weight (1), top specimen grip (2), bottom specimen grip (3), normal force load-cell (4), 10× lens (5), CCD camera (6), XYZ-stage (7), Xstage (8), ball-screw (9), stepping motor (10), displacement gauge (11), tractive force load-cell (12), and counter balance (13).

Fig. 2. Schematic of measurement for contact resistance.

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Fig. 3. Change in contact resistance of the connector across the contact zone as a function of fretting cycles with a span amplitude of 34 μm at 298 K, 323 K, 348 K and 373 K.

radius of 30 mm, and total length of 55 mm. Fig. 4 shows the tensile stress-strain curves of brass and tin. To determine the size of the contact area, structural analysis was performed considering two different cases, one in which there was a tin-coated layer of 10 μm and another in which there was no such layer. FEM analyses were conducted adopting ABAQUS (version 6.6) as the solver and HyperMesh (version 7.0) as the pre- and postprocessor. The specimen was modeled using shell elements. The number of nodes and elements of the fretting specimen model were 7740 and 7418, respectively. The minimum element size of the contact point was 1.5 μm. Perfect elastic and non-linear kinematic hardening elastic-plastic material model, as shown in Fig. 4, were used in the structural analysis. As can be seen in Fig. 5 which shows the loading and constraint conditions, the load is applied to the upper plate, and the lower plate is fixed. To apply the contact condition between the upper and lower plates, the upper plate is set to be the master and the lower plate is set to be the slave. The friction coefficient is assumed to be 1.3 [9]. Fig. 6 shows the three-dimensional distributions of pressure when the structural analysis was conducted using the non-linear kinematic hardening elastic-plastic material model. From the result of the analysis, the radius of the contact area was found to be 75.7 μm; the maximum contact pressure was 86.5 MPa. Referring to the figure, it can be seen that the pressure increases rapidly at the point where contact between the upper and lower specimen ends. The abrupt increase of pressure is thought to be caused by the pile-up phenomenon in which contact formed because the specimen is coated with very soft tin.

Fig. 4. Tensile stress-strain curves of (a) brass and (b) tin.

resistance. Whether this is actually occurring needs to be confirmed. The total resistance Rt becomes [10]

Rt ¼ Rs þ R f ¼

ρs ρ f d f þ 4a πa2

3.3. Theoretical contact resistance Fig. 3(a) shows that the initial electric resistance of approximately 0.002 Ω suddenly drops after approximately five cycles at 298 K. This seems to be due to fracturing of the initial tin oxide layer. If the oxide layer existed at the initial phase of the fretting test, the electric resistance will increase to the value corresponding to that of tin oxide layer; at this point, metal-to-metal contact spots will be formed when the oxide layer is destroyed due to fretting, resulting in reduced contact

Fig. 5. 2-D mesh of the cross section of the fretting specimen.

ð1Þ

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Fig. 6. 3-D contact pressure distribution of tin coated brass specimen under 0.85 N, considering elastic-plastic deformation behavior.

where Rs and Rf are spreading resistance and film resistance, respectively. ρs and ρf are spreading and film electrical resistivity, respectively. If the tin oxide is considered to be a film and if the tin-plated layer contact radius is a, the oxide layer thickness df can be determined. Tin oxide films formed at temperature lower than 393 K were reported to be amorphous SnO [11]. The thickness of the tin oxide layer is determined by applying electrical resistivity to SnO and tin of ρf = 1Ωm and ρs = 1.16 × 10− 7Ωm [12]. As a result of the structural analysis of the load P = 0.85 N, the radius of the contact surface is found to be 67 μm, as can be seen in Table 1. Therefore, when assuming this contact radius, the thickness of the oxide layer can be calculated as follows: Rt ¼

1Ωm  d f 1:16  10−7 Ωm þ ¼ 0:002Ω 4  0:000067m π  ð0:000067mÞ2

This value is just about 7% of the actually observed electric resistance of 0.002 Ω. This fact appears to show that the effective contact area is much smaller than the apparent contact area of π × (67 μm)2. This means that the contact is made through multiple tiny contact points, not a large single contact point. Conclusively, the reason that the electric resistance increases initially is because a partial contact is formed between two layers of tin, as the tin oxide layer is damaged as soon as the contact force is applied. After that, with fretting cycles, multiple contact points are formed between the two layers of tin, resulting in

ð2Þ

Based on this formula, when we calculated the film contact resistance using the tin oxide value as the electrical resistivity of film, the _ oxide layer thickness df was calculated as:d ¼ 2:2  10−11 m ¼ 0:22A. f

In this study, we stabilized the specimens in a chamber for about 1 h before starting the fretting tests. The tin oxide thickness after an exposure time of 1 h in air at 298 K–373 K was reported to be approximately 10 Å [13]. In the case of the specimen of this study, the calculated thickness value of 0.22 Å is too small and about 10−2 orders of magnitude less than the thickness of the oxide formed after an exposure time of 1 h at 298 K–373 K. Threshold load of approximately 1 N was reported to be enough to fracture the aluminum oxide layers, which were stronger than tin oxide [14]. This means that the oxide layer fractured when the test started under an applied load of 0.85 N; and, the actual contact is partially composed of conductive tin, resulting in low contact resistance. Meanwhile, if we consider that the contact is formed by a tin-plated layer on the copper substrate with a thickness of df and a contact radius of a regardless of the tin oxide layer, the total electrical resistance Rt can be calculated as follows. It can be assumed that the electrical resistivity of copper is ρs = 1.65 × 10−8Ωmand df is 10 μm. Rt ¼

1:65  10−8 1:15  10−7  10−5 þ ¼ 1:4  10−4 Ω 4  0:000067 π  ð0:000067Þ2

ð3Þ

Table 1 Summarized FEM results for contact pressure under 0.85 N. Material

Deformation behavior

Contact radius (μm)

Brass/tin film Brass/tin film Tin

Elastic–plastic Elastic Elastic

75.7 33.2 39

Fig. 7. (a) SEM micrograph of the fretting wear surface of the specimens after 5 cycles at 323 K with a span amplitude of 34 μm and (b) magnified surface.

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Fig. 8. SEM micrographs of the fretting wear surface of the specimens at (a) 298 K, (b) 323 K, (c) 348 K and (d) 373 K, respectively, with a span amplitude of 32 μm.

reduction of the resistance. Then, as the fretting continues, the electric resistance increases gradually due to wear and the presence of oxide debris. Fig. 7(a) and (b) show the surface state at the 5th cycle from Fig. 3(b). Referring to the figure, the fracture of the tin-plated layer during the initial stage of the experiment can be confirmed. 3.4. Surface analysis The areas of fretting contact wear at various fretting span amplitude were observed using the SEM. Figs. 8(a), (b), (c) and (d) show the fretting wear surface of the specimen with a fretting span amplitude of 32 μm at temperatures of (a) 298 K, (b) 323 K, (c) 348 K and (d) 373 K, respectively. These figures show the build-up of tin and oxide debris. As the temperature rises, the size of the debris around the damaged area became smaller, and the damaged area showed a smoother surface. This observation is in good agreement with the experimental results, in which it was found that the value of the friction coefficient decreased as the testing temperature increased. The area damaged by fretting wear was analyzed using energy dispersive spectroscopy (EDS) installed on SEM. Fig. 9(a) and (b) show the results of the EDS analysis of the fretting wear surfaces of the specimens at 373 K with a span amplitude of 25 μm, while moving up to the 640th cycle, where the resistance is 1 Ω. Fig. 9(a) shows a wear scar due to surface damage caused by fretting wear and oxide debris around it.

Referring to Fig. 9(b), the component analysis shows approximately 80% of tin and 20% of oxygen where fretting wear did not take place; a very small amount of zinc and copper was detected. This means that the non-fretted area was initially covered with a thin oxide layer. However, in areas of fretting wear, the concentration of tin increased to 35%, while that of oxygen decreased to 55%. This can be explained by the inflow of the oxides of tin to the electric contact area of the specimen; these oxides increased the contact resistance. This observation was similar for the four different temperature conditions. 3.5. Estimation of contact failure lifetime Several criteria can be used to determine the lifetime of electrical contacts. Whitley and Malucci [15] suggested that failure at electric contacts of noble metals like gold should be set to the time when the electric resistance becomes 10 times higher than the resistance value for a clean surface. However, this cannot be applied to connectors of tin alloy due to the initial formation of an oxide layer. Meanwhile, Mroczkowski [16] pointed out that electrical failure should be determined using a set value for each application area. The user of a specific application area needs to determine the value of contact resistance at which an application system stops functioning. Different criteria values have been set for different cases (e.g., contacts for sensor signals, contacts for power supply, etc.) and for different companies.

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Fig. 11. Fretting cycles to 0.01 Ω of tin-plated brass as a function of span amplitude at different temperatures in the range of below 35 μm span amplitude.

Fig. 9. (a) SEM micrographs of the fretting wear surface and (b) EDS results of the specimens at 373 K with a span amplitude of 25 μm.

By assuming that the specimen (tin-plated connector) will be used as a contact for sensor signals, in this study, the electrical contact failure has been set at 0.01 Ω. Fig. 10 shows the number of cycles to failure

under the load applied. Fig. 10 shows that the lifetime is reduced as the temperature increases at a constant contact force of 0.85 N. This is due to the fact that the build-up of oxides on the tin-plated layer, which is thicker formed as the temperature increases, leads to higher contact resistance, shortening the lifetime of the connector. Further, a greater span amplitude also results in reduced lifetime at the same temperature. As shown in Fig. 10, the slope of the plot changes at the span amplitude of approximately 30 μm. This means that the behavior of the degrading mechanism for reaching 0.01 Ω becomes different at a span amplitude of approximately 30 μm. The relationship of the span amplitude and the lifetime has to be divided into the two different regions, before and after 30 μm. However, governing degrading mechanisms in these two regions are not clear at present. The slopes of the plot in the range of over 30 μm span amplitude are not equal, depending on temperature. This implies that the degrading mechanisms are not identical and depend on testing temperature in this range. Meanwhile, Fig. 11 shows the relationship between the lifetime and the span amplitude in the range of below 30 μm span amplitude. As can be seen in this figure, the span amplitude and the lifetime are in inverse proportion to each other, and the slope of the plot is constant at approximately 0.1 for each temperature. This means that the relationship between the span amplitude and the lifetime is affected by an identical degrading mechanism. Therefore, it is possible to predict the lifetime as a function of t5he span amplitude in the range of below 30 μm span amplitude. Based on the above, the following formula is suggested. First, the rate _ due to the degrading of the electric contact is assumed to of damage (D) be inverse-linearly proportional to the number of cycles to failure (Nf). Second, as the temperature increases, the rate of damage due to fretting (dF / dN) of the contact increases as well. Therefore, the Arrhenius equation can be applied. Third, the rate of damage due to the degrading of the contact is assumed to be caused by the fretting (dF / dN). Fourth, as shown in Fig. 12, based on the fact that the rate of damage due to the degrading of the contact can be expressed as the exponent of the span amplitude, the following formula can be developed:   1 dD _ Q ∝ F ¼ Aδn exp − ≡ D_ ¼ Nf dN RT

Fig. 10. Fretting cycles to 0.01 Ω of tin-plated brass as a function of span amplitude at different temperatures.

ð4Þ

In the above formula, Q indicates the activation energy for fretting damage, R is a gas constant (= 8.31 J/mol·K), T is the testing

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surface and reported the activation energy for the growth of tin oxide film [17]. They reported an activation energy in the range of 5.5– 7.7 kJ/mol in a temperature of 298 K–393 K, which values are more or less than half of the present activation energy (=11.5 kJ/mol). The activation energy for the growth rate of an intermetallic compound formed at the tin-copper interface is reported to be 25 kJ/mol [14]. Therefore, a one-to-one relationship cannot be established for the value of activation energy. Meanwhile, the material constant A has been determined to be 9.6 × 10−13. Conclusively, the following formula can be developed:   1 11:5kJ=mole ¼ 9:6  10−13 δ7:8 exp − Nf RT

ð5Þ

Fig. 13 provides a comparison of the predicted and experimental lifetimes in the range of below 35 μm span amplitude. This figure shows that the prediction of the lifetime is in good agreement with the experimental lifetime data within a factor of two. In the future, to develop a more accurate formula for predicting the lifetime, more data need to be secured for more temperature levels and span amplitudes. 4. Conclusion Fig. 12. Calculation of activation energy for rate of fretting damage at span amplitude of 28.5 μm.

temperature expressed as an absolute value, n is an exponent of the span amplitude (in this case, an average value of 7.8, as shown in Fig. 10) and A is a material constant. For the calculation of activation energy, the data in Fig. 11 were used to plot logarithmic against 1000/T at the span amplitude of 28.5 μm. The activation energy was then determined from the slop of the resultant _

d logD straight line, which is equal to −2:3R dð1∕TÞ . As the plot slope is

log4 = 0.602, as can be seen in Fig. 12, the activation energy is found to be 11.5 kJ/mol. At present, the physical characteristics of the activation energy cannot be defined exactly. In the case of SnO2, no published values can be found for the activation energy for the diffusion of oxygen into tin. Recently, Tamai et al. investigated the oxidation of a tin plated

With respect to the fretting corrosion of electric connectors, experiments have been conducted to analyze the effects of span amplitude at 298 K, 323 K, 348 K and 373 K, while applying a constant force of 0.85 N. For the experiment, riders and flats made of 0.3 mm-thick brass sheet were coated with 10 μm of tin. The electric resistance was measured by applying constant and repetitive displacement to the upper riders. As the number of fretting cycles increased, the electric resistance decreased. Initially, the resistance was very low, afterwards, however, the resistance increased rapidly and intermittently. When measured by setting the electrical failure at 0.01 Ω, it was found that the lifetime _ due was reduced as the temperature increased. The damage rate (D) to the degrading of the contact below span amplitude 35 μm was inverse-linearly proportional to the number of cycles. When the Arrhenius equation is applied, the activation energy is 11.5 kJ/mol. A formula  for lifetime prediction was developed as follows: 1 N ¼ 9:6  10−13 δ7:8 f  expð−11:5kJ=mole RT Þ. By using this formula, the lifetime can be predicted within a factor of two in the range of below 35 μm span amplitude. Acknowledgments This study was supported by the Research Program funded by SeoulTech (Seoul National University of Science & Technology). References

Fig. 13. Comparison between observed lifetime (Nf) and predicted lifetime (Np) at different temperatures in the range of below 35 μm span amplitude.

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