Corrosion of Magnesium, Aluminum, and Steel Automotive Sheet Metals Joined by Steel Self-Pierce Rivets

THESIS

Presented in Partial Fulfillment of the Requirements for the Degree Master of Science in the Graduate School of The Ohio State University

William E. Weimer

Graduate Program in Materials Science and Engineering

The Ohio State University

2015

Master's Examination Committee:

Dr. Gerald Frankel, Advisor

Dr. Rudolph Buchheit

William E. Weimer

2015

Abstract

The automotive industry is investigating advanced light-weight materials in an effort to increase the fuel efficiency of vehicles. High-strength steels, aluminum and magnesium alloys, and carbon fiber and polymer composites are of interest to automobile manufacturers around the globe. One of the major problems facing the widespread implementation of such materials, especially aluminum and magnesium alloys, is their unique corrosion susceptibility. Not only is the corrosion performance of aluminum and magnesium alloys different from steel that is typically used in automobile manufacturing, but when used in combined, mixed-material systems, galvanic corrosion becomes a significant concern.

The United States Department of Energy, in conjunction with Ford Motor

Company, General Motors, and Fiat-Chrysler of America, has established standards that will be enacted to increase the corporate average fuel economy, or fleet-wide average fuel economy, of vehicles to be sold in the United States. This standard is intended to inspire automobile manufacturers to increase the fuel efficiency of vehicles weighing less than

8,500 lbs.

The Materials Technology Subprogram of the Vehicle Technologies Office is responsible for the investigation into new advanced materials that will enable technologies to increase fuel efficiency. The United States Automotive Materials

Partnership, LLC. has been established with funding from the Department of Energy in

order to design a magnesium-intensive front-end substructure. The first phase of the

project resulted in a 44.5% weight reduction and parts-count reduction from 110 to 47.

Now in its third phase, the Partnership has enlisted The Ohio State University in

an effort to create a magnesium-intensive front-end demonstration structure, consisting of advanced high-strength steels, aluminum alloys, and magnesium alloys of interest to the original equipment manufacturers participating in the project. The work performed in this thesis has contributed by investigating a case study of an AZ31 cleaning procedure for the MagPASS® conversion coating. Fourier-transform infrared spectroscopy was enlisted to analyze organic compounds on the surface of AZ31 magnesium alloy sheet metal that were interfering with pretreatment uptake. Polymerization of the organic compounds occurred during a warm-forming procedure, causing them to undergo a structural transformation that rendered the cleaning procedure ineffective.

The hydrogen embrittlement of hardened-steel rivets coupled to Mg panels was then studied. Considering their close proximity to magnesium alloys, which generate copious hydrogen during corrosion, it was proposed these rivets (RC 47) could potentially experience degradation of mechanical properties. A unique slow-strain-rate hoop-stress-test was designed in an effort to elucidate changes in mechanical properties of the rivets after hydrogen charging. Unfortunately, the non-ideal geometry of the rivets made hydrogen charging ineffectual, and the hoop-stress-test results were inconclusive.

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Finally, changes in lap-shear strength of riveted joints as corrosion dissolved the

magnesium surrounding the rivets over time was investigated. It was found that the

riveted joints sustained a significant portion of their lap-shear strength despite significant

material loss in the adjacent magnesium. It was also found that the industry-standard

Zn/Sn-coating for the hardened steel rivets was superior to aluminum-coatings in mitigating the galvanic corrosion. Student-t statistical analysis was employed to establish the statistical significance of the results.

Dedication

This thesis is dedicated to all of those who have allotted their attention and invaluable time towards helping me to complete this work. I would also like to dedicate this work to

the United States Department of Energy Vehicle Technology Office Materials

Technology Program for initiating this research effort towards achieving better fuel economy in vehicles. Finally, this work is dedicated to the Ford Motor Company for their

commitment to sound business practices and intelligent technological foresight.

Acknowledgments

First, I owe a deep debt of gratitude towards Dr. Gerald Frankel, my advisor. He has been a wealth of knowledge and a bastion of consistency for me and everyone in his group. His indelible commitment to his own responsibilities, and the vast array of those of each of his students, has been an inspiration to me during my tenure with the Fontana

Corrosion Center. As my mentor and role model, Dr. Frankel has continuously exhibited the level of excellence I hope to one day exude.

Second, I would like to extend a special thanks to all of the excellent students with whom I have had the pleasure to study. I would like to especially thank those who worked closest with me including PhD candidate Jiheon Jun, Dr. Belinda Hurley, Kerrie

Holguin, Brandon Lynch, PhD candidate Jinwook Seong, PhD candidate Pitichon

Klomjit, Dr. Severine Cambier, Sara Grieshop, Sara Cantonwine, Sean Morton, PhD candidate Zhicao Feng, PhD candidate Xiaolei Guo, Jermaine Onye, Daniel Kaminski, and Dan Campbell.

Third, I would like to acknowledge the help and support of both the people and organizations affiliated with the United States Automotive Materials Partnership. I would like to especially thank Dr. Robert C. McCune of Robert C. McCune and Associates,

LLC, who proved to be a patient, practical, and personable resource to me throughout the project. Thank you to Joy Hines Forsmark of Ford Motor Company who was consistently helpful and willing to extend herself and her resources to help me achieve my goals.

Thank you to Jeff Stalker and John Love of PPG, who were always friendly,

supportive, and willing to send materials and advice without hesitation. Finally, thank

you to Chris Jurey of Luke Engineering and Bruce Davis of Magnesium Elektron for

their support and advice.

Fourth, I would like to thank the dedicated human resources support staff, and innovative building and laboratory staff of the Department of Materials Science and

Engineering. Thank you to Mei Wang, Mark Cooper, and Megan Daniels for all your advice on the logistics of travel and purchasing, and for your flexibility in helping me through extenuating circumstances. A special thanks to Kenneth Kushner, Steven Bright, and Ross Baldwin, who proved time-and-again to be three of my most valuable colleagues during the pursuit of both my undergraduate and graduate degrees. All three afforded me extensive access to their department resources, knowledge, and expert advice, while maintaining a friendly and approachable demeanor.

Fifth, I would like to thank the faculty of The Ohio State University College of

Engineering for their substantial support and expert advice. Special thanks go to

Associate Professor Dr. Mark Ruegsegger of the Biomedical Engineering department. Dr.

Ruegsegger helped me with several meetings, much endearing advice, and access to his

Fourier Transform Infrared Spectrometer at no cost. Thank you to Professor Rudolph

Buchheit for always being helpful, encouraging, and willing to allow me access to his vast knowledge of the corrosion field; thank you also, to Dr. Buchheit, for his role as the second member of my thesis committee. Thank you to Professor David Tomasko of the

Department of Chemical and Biomolecular Engineering for his advice, encouragement,

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and resources. Also, thank you to Dr. Alison Polasik of the Department of Materials

Science and Engineering for advice and encouragement.

Sixth, I would like to acknowledge the incredible amount of personal and

professional support I received from those in the Human Resources Department of the

Ohio State University College of Engineering. Thank you, especially, to Amy Franklin

and Winifred Sampson for their continuous encouragement, emotional support and

advice, and willingness to listen; I truly could not have made it without them.

Finally, I would like to thank my parents. Without their unconditional love and support throughout my pursuit of degrees, I would not be able to undertake the next step: my pursuit of happiness.

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Vita

2001-2005 ...... Pickerington High School Central

2006-2011 ...... Soil Testing and Engineering, Ltd, Quality

Control/Quality Assurance Technician

2011-2012 ...... Material Advantage, Treasurer, The Ohio

State University

2011...... Engineers’ Council, Treasurer, The Ohio

State University

2012...... Engineers’ Council, President, The Ohio

State University

2012...... Engineering Career Expo, Student

Coordinator, The Ohio State University

2009-2012 ...... B.S. Materials Science and Engineering, The

Ohio State University

2012-present ...... Graduate Research Assistant, Materials

Science and Engineering, The Ohio State

University

Fields of Study

Major Field: Materials Science and Engineering

Table of Contents

Abstract ...... ii

Dedication ...... v

Acknowledgments...... vi

Vita ...... ix

Fields of Study ...... ix

Table of Contents ...... x

List of Tables ...... xv

List of Figures ...... xvii

Chapter 1: Literature Review ...... 1

Background ...... 1

Light-weight Materials ...... 4

Magnesium as a Structural Material ...... 6

Corrosion Mechanisms of the Mg-H2O System ...... 7

Galvanic Corrosion ...... 9

Protecting Against Galvanic Corrosion ...... 12

Magnesium Surface Treatment ...... 13

Self-Pierce Riveting ...... 14

Hydrogen Embrittlement ...... 14

Conclusions ...... 14

Tables and Figures ...... 16

Chapter 2: AZ31 Cleaning Procedure for MagPASS Conversion Coating ...... 23

Introduction ...... 23

Experimental Procedures ...... 24

Results and Analysis ...... 28

Simulation of the warm-forming procedure ...... 29

Cleaning and pretreatment ...... 30

Discussion ...... 31

Conclusions ...... 32

References ...... 33

Figures ...... 34

Chapter 3: Hydrogen Embrittlement of Hardened Steel Rivets ...... 43

Introduction ...... 43

Experimental Procedures ...... 44

Rivets ...... 44

Hydrogen Charging ...... 44

Galvanic Corrosion Charging ...... 45

Explanation of mechanical test ...... 45

Fixtures ...... 46

Test parameters ...... 47

Results and Analysis ...... 47

Hydrogen-Charged rivets ...... 48

Discussion ...... 50

References ...... 53

Figures ...... 54

Chapter 4: Effect of Corrosion on Lap-Shear Strength of AM60B Magnesium to 6082-T4

Aluminum Joints formed by Self-Pierce Steel Rivets ...... 61

Introduction ...... 61

Rivet coating ...... 63

Sheet-metal Lay-up Material Selection ...... 63

Material electrochemistry ...... 64

Corrosion Phenomenology of SPR in ASTM B117 ...... 65

Effect of Adhesive ...... 66

Stress state - adhesive versus non-adhesive bond ...... 66

Rivet-only Joints ...... 66

Rivet & Adhesive Joints ...... 67

Lap-shear Testing, Magnesium Failure and Variance Analysis ...... 67

Analysis of Statistical Variance in Lap-shear testing ...... 67

xii

Error Variance in Rivet-only Population ...... 68

Error Variance in Rivet + Adhesive Population ...... 69

Treatment Variance ...... 69

Welch-Aspin Test ...... 70

Experimental Procedures ...... 70

Sample Preparation ...... 70

ASTM B117 Exposure ...... 71

Lap-shear testing...... 71

Baseline testing ...... 71

Results and Analysis ...... 72

Baseline lap-shear data ...... 72

Failure in Rivet-only samples ...... 73

Failure in Rivet & Adhesive Samples ...... 73

ASTM B117 Salt-spray exposure ...... 73

Summary of t-statistics ...... 75

ANOVA ...... 77

Discussion ...... 78

Conclusions ...... 79

References ...... 81

Tables and Figures ...... 82

xiii

Summary and Suggestions for Future Work ...... 100

Comprehensive list of references ...... 102

xiv

List of Tables

Table 1: Chemical potentials of Mg compounds at 25°C.41,42 ...... 16

Table 2: Nominal composition (wt %) of magnesium and aluminum alloys used in this

study.4,5...... 82

Table 3: General mechanical properties of AM60B magnesium alloy and aluminum alloy

6082-T4.4,5 ...... 83

Table 4: Mg/Al 2-plate assemblies designated by letter and number code. Each assembly

was sectioned into four individual lap-shear coupons...... 84

Table 5: B117 salt spray exposure schedule for sectioned, top-coated samples. 500 hr

exposures were conducted at PPG’s laboratory...... 88

Table 6: Lap-shear results for unexposed samples ...... 90

Table 7: Ultimate load and failure mechanism of coupons exposed to B117 for 168 hours.

...... 92

Table 8: Ultimate load and failure mechanism of coupons exposed to B117 for 336 hours.

...... 93

Table 9: Ultimate load and failure mechanism for samples exposed to B117 for 500 hours

in salt spray chambers at PPG’s facility...... 94

Table 10: Summary of t-statistics regarding significance of mean variance - pass/fail

(green/red) ...... 95

Table 11: Error variance as determined by lap-shear testing of unexposed coupon

populations ...... 95

Table 12: Scoring system for performance of coupon set in B117 exposure. A t-stat

“score” indicates that B117 exposure significantly altered the mean lap-shear strength.

An F-ratio “score” indicates that B117 exposure significantly altered the variation in the lap-shear strengths. A lower score indicates better performance...... 95

xvi

List of Figures

Figure 1: Galvanic Series in seawater.43 ...... 17

Figure 2: Pourbaix diagram for magnesium in water.44 ...... 18

Figure 3: Experimental setup showing 2-liter beakers with solution – omitting the

glycolic step here (top) and coupons prior to being cleaned (bottom)...... 34

Figure 4: Flow chart of coupon treatment schedule including the step to simulate the

warm-formed condition...... 35

Figure 5: as-received AZ31B-H24 from MENA with “mill-finish” surface condition.

Coupon on left was received about 6 months prior to that on right. Material on right was

used to acquire data for this report, as its processing history and shelf life was known,

while that on the left had an unknown history...... 36

Figure 6:. FTIR spectra of as-received mill-finished (MF) and as-received warm-formed

(WF) upper rail coupons prior to cleaning. Boxes indicate absorption signal characteristic

of C-H bond bending (735-770 cm-1), C-C aromatic bond stretching (1500-1600 cm-1)

-1 and CO2 in gas phase (2350 cm )...... 37

Figure 7: FTIR spectra comparing warm-formed upper rail coupon to mill-finished coupon exposed to 300°C for 15 minutes, prior to cleaning. Boxes indicate absorption signal characteristic of C-H bond bending (735-770 cm-1), C-C aromatic bond stretching

-1 -1 (1500-1600 cm ) and CO2 in gas phase (2350 cm ) ...... 38

xvii

Figure 8: FTIR spectra comparing warm-formed upper rail coupon to mill-finished coupon exposed to 350°C for 15 minutes, prior to cleaning. Boxes indicate absorption signal characteristic of C-H bond bending (735-770 cm-1), C-C aromatic bond stretching

-1 -1 (1500-1600 cm ) and CO2 in gas phase (2350 cm ) ...... 39

Figure 9: FTIR spectra comparing as-received mill finished, as-received warm-formed

upper, warm-form simulated at 300 and 350°C after exposure to phosphoric acid etching

cleaning procedure. Boxes indicate absorption signal characteristic of C-H bond bending

(735-770 cm-1), P=O stretching (1140-1320 cm-1), C-C aromatic bond stretching (1500-

-1 -1 1600 cm ) and CO2 in gas phase (2350 cm )...... 40

Figure 10: FTIR spectra comparing as-received mill finished, as-received warm-formed

upper, warm-form simulated at 350°C after exposure to glycolic nitrate pre-etch and

phosphoric acid etch cleaning procedure. Boxes indicate absorption signal characteristic

of C-H bond bending (735-770 cm-1), P=O stretching (1140-1320 cm-1), C-C aromatic

-1 -1 bond stretching (1500-1600 cm ) and CO2 in gas phase (2350 cm )...... 41

Figure 11: FTIR spectra comparing as-received mill finished, as-received warm-formed

upper, warm-form simulated at 300 or 350°C, and Salzgitter “pristine” AZ31 after

exposure to glycolic nitrate pre-etch and phosphoric acid etch cleaning procedure. Boxes

indicate absorption signal characteristic of C-H bond bending (735-770 cm-1), P=O

-1 -1 stretching (1140-1320 cm ), C-C aromatic bond stretching (1500-1600 cm ) and CO2 in gas phase (2350 cm-1). Notable differences exist between Salzgitter and other AZ31. ... 42

Figure 12: Microstructure of rivet steel. The darkened phase is martensite and lighter

phase is ferrite...... 54

xviii

Figure 13: Schematic of the hoop-stress test punch tool design...... 55

Figure 14: Photograph of rivet press tool...... 55

Figure 15: Photograph of rivet press tool in load frame, ready to be compressed...... 56

Figure 16: Force versus displacement curves generated on as-received rivets by rivet

punch test. Three distinct regions are separated by dashed lines and numbered...... 57

Figure 17: Hoop-stress test of uncoated rivets in as-received, H-charged, and corrosion- charged conditions...... 58

Figure 18: Hoop-stress test of Al-coated rivets in as-received, H-charged, and corrosion-

charged conditions...... 59

Figure 19: Hoop-stress test of Sn/Zn-coated rivets in as-received, H-charged, and

corrosion-charged conditions ...... 60

Figure 20: Sheet-metal lay-ups were sectioned along the red dotted lines into four

individual lap-shear specimens...... 86

Figure 21: coated samples with masked edges...... 87

Figure 22: uncoated samples ...... 87

Figure 23: Schematic of shimmed lap-shear setup...... 89

Figure 24: Lap-shear force vs. displacement curves of Al/Mg joint with rivet + adhesive

(blue) and rivet only (red)...... 91

Figure 25: Schematic of Mg galvanic corrosion. Dissolving magnesium creates hydroxyl ions, increasing the local pH. This passivates the Mg near the rivet and causes the majority of dissolution to occur at a slight distance away from the rivet. This has implications for the fracture mechanics of the Mg failure...... 96

xix

Figure 26: Sn/Zn-coated rivet samples after 500 hours ASTM B117 exposure. Visually all Sn/Zn-coated rivets displayed the least corrosion of all three coupon groups...... 97

Figure 27: IVD-Al coated rivets with no adhesive after 500 hours ASTM B117 exposure.

Note the large amount of Mg dissolution surrounding the rivet...... 98

Figure 28: IVD-Al coated rivets with adhesive after 500 hours ASTM B117 exposure.

Visually the corrosion was very similar to the IVD-Al coated rivets with no adhesive. .. 99

Chapter 1: Literature Review Background

Metal alloys of lower density than steel are of increasing interest to the automotive industry.1 Steel has long been a mainstay material for the auto industry because of its

superior mechanical properties and low cost relative to other structural metals. The

impetus for investigation into lighter metals comes from regulation which is being

incrementally enacted by two federal agencies: the National Highway Traffic Safety

Administration (NHTSA) and the Environmental Protection Agency (EPA).

The NHTSA is in the midst of setting corporate average fuel economy (CAFE)

standards which are expected to result in an industry-wide fleet average fuel efficiency

improvement to 35 mpg-equivalent by 2020.1 The Energy Policy and Conservation Act

(EPCA), as amended by the Energy Independence and Security Act (EISA), is the

legislative mandate to be enforced by the NHTSA and EPA. A Presidential Memorandum

was issued by President Obama on May 21, 2010 requesting the NHTSA and the EPA to

further improve the fuel efficiency and reduce emissions of light vehicles for the years

2017-2025.1 This is an incremental improvement following the success of the first phase of the CAFE program, which was enacted from 2012-2016. One of the statutes of the

legislation requires that the NHTSA only extrapolate fuel economy standards forward

five calendar years at a time.

Despite the large cost of research and development of technologies that result in

greater fuel efficiency, the program is expected to stimulate the United States economy

by a net amount of $326 to $451 billion over the course of the 2017-2025 second phase.1

This benefit will be enjoyed mainly by consumers, who can expect net savings of $3,400 to $5,000 over the course of the lifetime of their vehicle. This amount is presented as a net amount because the technology implemented to reduce fuel consumption will require a larger up-front cost to the consumer. This upfront cost will be mitigated by the money saved by increased fuel efficiency.

The environmental benefits of the program will be quantified by measuring the reduction in the carbon footprint of light vehicles.1 Not only is implementation of the

second phase expected to save 4 million barrels of oil, but it is expected to reduce

greenhouse gas-emissions (GSG) by the equivalent of 2 billion metric tons. NHTSA

projects the second phase will result in an average industry fleet-wide fuel efficiency of

48.7-49.7 miles per gallon (mpg) by year 2025.1,2

According to the United States Department of Energy (DOE) Office of Energy

Efficiency and Renewable Energy's (EERE) Vehicle Technologies Office (VTO),

“advanced materials are essential for boosting the fuel economy of modern automobiles

while maintaining safety and performance.”1,2 A 10% reduction in gross vehicle weight

results in a 6-8% fuel efficiency improvement. Not only does weight reduction result in

direct fuel savings, it allows for auto manufacturers to incorporate advanced emission-

control equipment, safety devices, and integrated electronic systems which would

otherwise limit the benefits enjoyed by adding such technology. Additionally, more

efficient combustion can be achieved by the use of advanced materials in the motor, and

further research and development of such materials can mitigate the increased cost of

incorporating materials that are more expensive than the steel components that are being

replaced.

The Materials Technologies subprogram (MTS) of the VTO is investigating

replacing cast-iron and other traditional steel components with advanced high-strength

steel, aluminum and magnesium alloys, as well as carbon-fiber and other polymer composites.1,2 This next generation of light-weight automotive materials could potentially

reduce the weight of vehicles by as much as 50%. Over 170 billion gallons of fuel are

consumed by more than 240 million vehicles every year in the United States, alone.

Vehicle weight-reduction and high-efficiency motors, achieved by the incorporation of

advanced materials in one quarter of the U.S. fleet, could save 5 billion gallons of fuel

per year by 2030.

In 2010, it was established that the goal of the MTS was “to validate a cost-

effective weight reduction of 50% in body and chassis systems by 2015 while

maintaining safety, performance, and reliability comparable to 2002 vehicles. For

passenger vehicles, the goal is to validate a cost-effective weight reduction of 50% in body and chassis systems by 2015 while maintaining safety, performance, and reliability comparable to 2002 vehicles. For commercial vehicles, the goal [was] to develop the

materials needed to improve heavy-duty engine efficiency to 55% by 2012 while meeting

U.S. Environmental Protection Agency (EPA) emissions standards.”1,2

In collaboration with universities, national laboratories, and the automotive

industry, VTO material technology research and development (R&D) focuses on two

strategies. One is “Developing advanced engine and powertrain materials that enable

more efficient propulsion systems,” and the other is “Lightweighting (reducing the

weight of) body and chassis components through use of advanced materials.1,2”

Light-weight Materials

Several challenges face designers and manufacturers of automobiles that will

incorporate new materials.1,2 To improve combustion efficiency, materials must be able

to perform in more extreme conditions, such as higher temperatures, pressures, and

corrosiveness. For lightweighting efforts, materials will be subject to these performance

demands, but also new manufacturing methods must be developed that will be capable of

producing new materials in a way that will keep costs low enough for consumers.

The VTO has resulted in significant progress towards commercialization of

lighter-weight vehicles.1,2 In 2014, Ford Motor Company introduced to consumers the

first aluminum-chassis light pickup truck (2015 Ford F-150 full-size pickup truck). The

pickup is 700 lbs lighter than its 2014 predecessor. In line with the objectives of the

VTO, the Ford pickup achieves 19 miles per gallon (mpg) in city driving and 26 mpg

during freeway driving. This is a 22% increase in overall fuel efficiency versus the 2014

model.

6,000-series aluminum alloys and advanced high strength steels (AHSS) are two

examples of the new materials implemented in the 2015 F-150 to achieve the increased

fuel efficiency.3 In order to push the envelope farther, magnesium alloys are also being

investigated as part of the MTS. Aluminum is about a third of the density of steel, while

magnesium is about a fifth. Magnesium alloys have been shown to be viable structural

materials.

The combined R&D expertise of universities, national laboratories, and

automotive industry scientists and engineers resulted in a magnesium alloy engine cradle

for the Chevy Z06 Corvette (2006).2,4 The magnesium engine cradle was 65% lighter

than a steel cradle and 35% lighter than an aluminum cradle. For the successful

implementation of magnesium alloys, many problems must be addressed. In order to

illuminate and solve these problems, the VTO has established government/industry

partnerships. The resulting efforts are the “U.S. DRIVE Partnership,” focusing on light-

duty vehicles; the “21st Century Truck Partnership,” focusing on heavy-duty vehicles; and the US Automotive Materials Partnership, LLC. (USAMP).

Ohio State's involvement is with the USAMP portion of the government/industry partnerships. In 2004, a three-country meeting between China, Canada, and the United

States resulted in a collaborative effort to research technologies that would enable the design of a “magnesium-intensive vehicle front-end substructure” (MIVFES) The first phase of the project, magnesium front-end design and development (MFEDD), proved the viability of designing the substructure using magnesium alloys.2,4-7 The project

5 resulted in an MIVFES that was 44.5% lighter than the steel predecessor. Additionally, a part-count reduction from 110 to 47 was achieved.

Now called magnesium front-end research and development (MFERD), the project began its third phase on June 1, 2012. The project is funded by a DOE directive to design a magnesium-intensive demonstration structure.5-7 The Fontana Corrosion Center at Ohio State is participating in the Task 5 Corrosion and Surface Treatment portion of the MFERD project, which consists of nine task teams. The objectives of Task 5 are as follows (language taken directly from the third phase OEM Kickoff Meeting (July 30,

2012):

1. Explore commercial and developmental corrosion protection ‘systems’ for

magnesium alloys and their integration with vehicle assembly and painting.

2. Explore and validate measurement methods to compare different protection systems.

3. Explore protection methods for steel fasteners to limit galvanic attack.

4. Develop “pilot” scale production facilities for painting “demonstration” structures.

5. Conduct various OEM cyclic testing on assembled ‘demo’ structures.

Magnesium as a Structural Material

Mg alloys have attractive properties including good formability, good recyclability, and a high strength/weight ratio.8 The aerospace and automobile industries have a particular interest in Mg alloys as an alternative to high-strength aluminum and steel in structural applications.8 This review will focus on the electrochemical nature of 6

Mg alloys in environments that would be encountered in manufacturing and in service in the automobile industry. Comparisons will be made to steel as it is the industry standard material in automobile structures. According to some, widespread acceptance of Mg alloys will not occur until the auto industry embraces its use.8-10

The main challenge facing the infiltration of Mg alloys into the automobile manufacturing process lies in the fact that Mg is an electrochemically active material.7-10

With a standard electrochemical potential of -2.363 VSHE, Mg is one of the most anodic metals in the electromotive series. The mechanisms by which Mg and its alloys corrode are complicated and disputed; there is a need for quantification of corrosion mechanisms and documentation of intrinsic corrosion rates in order to establish a foundation by which existing and new Mg alloys may be characterized. The electrochemical nature of pure Mg metal will now be reviewed in order to form a thermodynamic foundation of Mg stability.

Corrosion Mechanisms of the Mg-H2O System

This section will consider the fundamental electrochemical interactions of pure

Mg with water. This will elucidate the complexity of Mg corrosion phenomenology and form a basis for understanding the thermodynamic and kinetic processes that govern corrosion in Mg alloys. As mentioned previously, Mg is a thermodynamically unstable material in most environments; passivating surface oxides are stable only in very alkaline

10-14 solutions (pH>10.5). This is illustrated by the E-pH diagram for the Mg-H2O electrochemical system (Figure 2). The basic Mg oxidation reaction is:

0 Mg+2H2O = Mg(OH)2 + H2 ΔG = -359 kJ/mol (1)

The standard Gibbs free energy (ΔG0) for the above reaction is strongly negative,

thus Mg readily oxidizes, even in pure water. Actually, water is essential to corrosion of

Mg because it supplies the environmental elements to sustain the cathodic reactions.

Mg(OH)2 is the most thermodynamically stable Mg corrosion product. In most aerated

solutions, the equilibrium oxidized Mg surface composition will consist mainly of

14 Mg(OH)2. This is further supported by Table 1, which gives the chemical potential of

Mg and its compounds at 25°C.

Chemical potential (μ0) is inversely correlated with the thermodynamic stability of the compound. Thus, Mg(OH)2 is the most stable Mg compound in an aerated solution.

Interestingly, if oxygen is not present, hydrogen is stable in Mg as MgH2, indicating

hydrogen and Mg will form a MgH2 solid solution at ambient conditions. This has

implications for the use of Mg as a hydrogen storage material – Mg will tend not to

spontaneously release H2.

Many more possible chemical and electrochemical reactions may occur in the

Mg-H2O system. Most of the magnesium and water corrosion reactions have negative

standard potentials and serve as anodic half-cells driving oxygen reduction or

electrochemical hydrogen processes in Mg corrosion.

As alluded to previously, water reduction is the typical cathodic reaction in Mg

corrosion. In an aqueous environment, the following reactions can be cathodic processes

to promote the dissolution of Mg.13,14 8

+ - 0 H2 = 2H + 2e E = 0 VSHE (2)

- - 0 H2 + 2OH = 2H2O + 2e E = -0.826 VSHE (3)

Reactions 2 and 3 have positive standard potentials relative to the standard

potential of magnesium, implying they may serve as cathodic reactions in Mg corrosion.

In reality, other parameters may control Mg dissolution such as solution

chemistry, temperature, hydrogen concentration and availability, and Mg surface

composition, to name a few. Oxygen concentration and oxygen diffusivity in solution

come into play for cathodic reactions during galvanic corrosion involving coupling

materials such as aluminum alloys and steels.

Galvanic Corrosion

Galvanic corrosion is a well-known phenomenon in the corrosion field. There are

three requirements for it to occur:15

1. Presence of dissimilar materials (usually metals)

2. An electrical connection between the dissimilar materials

3. An ionic connection between the dissimilar materials

From this point forward, “dissimilar materials” will be assumed to be metals. For corrosion to occur, there are two governing physiochemical forces that determine the rate

of the reactions that occur on the metal surface, usually resulting in dissolution of at least

part of the metal.15 The first force is the “driving force,” the other limits the rate of the

9 corrosion reactions. The driving force is derived from thermodynamics. It is dependent upon the chemical composition of the metal and its thermodynamic stability in the ambient environment. “Kinetics” is the term commonly used to describe the phenomenon, or phenomena, that limit the rate of the corrosion reactions.

In order to quantify the thermodynamic driving force for corrosion, the electrochemical potential of the metal is established. The electrochemical potential is related to the chemical potential of the metal and its proclivity to exchange electrons with the ambient environment.15 In corrosion science, the electrochemical potential of a metal varies vastly depending on the environment.

However, a useful tool for engineers and scientists to quickly “predict” the interaction of dissimilar metals is called the Galvanic Series (Figure 1). This provides a prediction of the electromotive driving force for galvanic corrosion.15 The Galvanic

Series is commonly established for metals fully immersed in seawater. Therefor the position of metals on the series also depends on the rate of the cathodic reaction, which is usually limited by oxygen diffusion through the seawater to the reacting surface.

Seawater is extremely corrosive due to its electrolytic nature, and is essentially a “worst- case-scenario” for the occurrence of corrosion in nature.

The galvanic series is a quantified scale of the electromotive driving force for corrosion to occur between dissimilar materials.15 The scale provides a ranking of the electrochemical potential of metals, from “noble” to “active.” Most galvanic series

extend from the most noble material, graphite (usually the only non-metal on the scale),

to magnesium, the most active material.

Depending on its position on the galvanic series, a prediction can be made as to the likelihood for self-corrosion to occur.15 Quite simply, the more active the metal, the

more likely it is to corrode. However, the actual corrosion rate will depend on the kinetics

of the electrochemical reactions, as well as the area ratio and ohmic potential drops in the

environment. The kinetics of the corrosion reactions occurring on or near metal surfaces are typically related to the presence of oxygen. In fact, the only reason most metals remain metal and do not immediately revert to their respective oxides, is directly related to the spontaneous formation of a very thin (10-6-10-9 m) oxide layer on the metal

surface.16-18 This oxide layer provides a kinetic barrier to further interaction between the

ambient environment and the metal microstructure. The Gibb's Free Energy of reactions

between water (hydrogen and oxygen) and most metals is on the order of kJ.16-18

Thus, metals should literally be explosively reactive with air. However, the oxide

layer that forms on the metal surface prevents oxygen from permeating throughout the

microstructure.16-18 Metals form oxide-layers of varying density and stability. Ultimately,

the corrosion-resistance of a metal (except for gold and platinum, which are

thermodynamically stable in most ambient environments) is derived directly from the

stability of this oxide layer.

Additionally, if dissimilar metals are present and in intimate contact (electrically

and ionically connected) – their respective location on the galvanic series has

implications for the nature of their galvanic interaction.15,19-20 Assuming intimate contact

between dissimilar metals, the galvanic series provides a quantified measurement of the

electromotive driving force for galvanic corrosion to occur. The more active of the two

materials will corrode at an “accelerated” rate, which depends on the “electromotive overpotential” resulting from the natural tendency of the chemical potentials of the two metals to equilibrate when electrically connected.

Protecting Against Galvanic Corrosion

Since most metals are metastable in ambient environments, exposed surfaces must be treated to maintain their metallic integrity. Typical treatment methods are procedural in nature, involving surface treatments such as chemical or physical cleaning and abrading, pretreatment, conversion coating, and finally painting.15-23 One of the

difficulties with using magnesium as a structural material comes from the complexity of

these procedures in Mg systems, especially in systems involving other materials.

Considering the Pourbaix diagram (Figure 2) for the Mg-H2O system, one can see that

Mg corrodes readily in solutions below a pH of about 10.5.

There exists a wide variety of surface treatments for steels that cater to specific applications.24 Steels are commonly supplied with a metallurgical coating such as a zinc

layer (galvanized steel) that protects the steel substrate (cathodic protection). A newer

application for metallurgical coating provides a galvanic “buffer” between steel and more

active materials.24 For example, a Zn/Sn coating on rivets can depress the electrochemical

potential of the rivets to make them suitable for joining more active materials, such as

aluminum and magnesium alloys. 12

Surface treatments for aluminum alloys are commonly either chemical or electrochemical in nature.25 Chemical treatments are procedural and typically involve a degreasing step using an organic solvent to remove oil and grease. The surface can then be either etched in acidic or caustic media to roughen the surface, or the surface may be pickled in a caustic solution. Electrochemical surface treatments for aluminum alloys usually involve anodizing,25 which uses electrolytic oxidation processes to alter the oxide layer in some desirable way, for example, to induce protective or functional properties.25

Due to magnesium’s narrow thermodynamic stability range, many typical treatment procedures for other metals will dissolve magnesium beyond its usefulness.26,27

This is not a major problem if magnesium can be treated independently. However, to maintain cost-effectiveness in the automotive industry, magnesium will have to be joined to other metals such as aluminum and steel prior to treatment. Thus, treatment procedures must be developed that are effective and cheap for mixed-metal systems.

A review of the literature has revealed that little progress has been made in this regard. The majority of research regarding this effort remains proprietary in nature. Thus, the following is a review of literature pertaining mainly to surface-treating magnesium, alone, rather than as-joined in a mixed-metal system.

Magnesium Surface Treatment

In order for the use of magnesium alloys in the automotive industry, methods must be developed to prepare the magnesium surface for the paint shop. Chemicals used to treat magnesium alloy surfaces typically have a solution pH greater than 10.5 because magnesium is thermodynamically unstable in solutions with pH below 10.5, as is

illustrated by Figure 2.25-30 This pH dependence, which is unique to magnesium, complicates the manufacturing process. Many surface treatments for aluminum alloys

and steels involve chemicals with a pH below 10.5, which will corrode the magnesium

prior to finishing. Ideal metal treatment procedures should be simultaneously effective on

steels, aluminum alloys, and magnesium alloys, thus parts can be joined prior to

undergoing surface treatments, necessarily saving time and money during manufacture.

Self-Pierce Riveting

Self-pierce riveting (SPR) entered the automotive industry as a joining technology in the early 1990’s.31It has many advantages as compared to other joining methods,

including quick (1-4 seconds per rivet) and quiet operation, the ability to join dissimilar

materials, the ability to join multiple material stacks, and it is a non-thermal process.32 It

is impossible, however, to avoid an electrical connection between dissimilar materials in

an SPR stack-up. Thus, galvanic interaction between the adjacent materials and any

fastener material is a concern.

Hydrogen Embrittlement

High-strength, high-hardenability steels are being increasingly implemented as part of the automotive industry’s lightweighting efforts. Unfortunately, environmental hydrogen can enter the steel microstructure, interact with dislocations and other trap sites,

and cause unwanted degradation of mechanical properties. When hardened steels are used

in conjunction with magnesium, which evolves copious hydrogen during corrosion, the

risk of deleterious environmental hydrogen embrittlement is increased.34-41

Conclusions

In order to successfully incorporate advanced materials in the automotive industry, many obstacles must be overcome. As mentioned above, magnesium alloys are extremely electrochemically active. Understanding the corrosion of magnesium alloys and the electrochemical phenomenology associated with galvanic corrosion between magnesium and other materials must be better understood. The following work is put forth in order to add to the existing body of knowledge regarding these topics.

Tables and Figures

Table 1: Chemical potentials of Mg compounds at 25°C.41,42

Species Oxidation State State μ0 (kcal/mol)

Mg 0 Solid 0 Mg+ +1 Ion -61 Mg2+ +2 Ion -109

Mg(OH)2 +2 Solid -199 MgH -1 Gas +34

MgH2 -2 Solid -8 MgO +2 Solid -136

Figure 1: Galvanic Series in seawater.43

Figure 2: Pourbaix diagram for magnesium in water.44

References

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2. Laboratory, National Renewable Energy. "Materials Technologies: Goals, Strategies, and Top Accomplishments." Ed. Energy, U.S. Department of Energy: Office of Energy Efficiency and Renewable2010. Print.

3. Sherman, Don. "In Depth with the 2015 Ford F-150's Aluminum, Presented in an Alloy of Facts and Perspective." (2014). Web.

4. Vehicle Technologies Program: Goals, Strategies, and Top Accomplishments. (2010, December 1). Retrieved January 9, 2015, from http://www1.eere.energy.gov/vehiclesandfuels/pdfs/pir/vtp_goals-strategies- accomp.pdf

5. (n.d.). Retrieved from http://www.uscar.org/guest/partnership/1/us-drive

6. McCarty, E., & Luo, A. (2008, February 28). USAMP AMD 603 – Magnesium Front End Design and Development. Retrieved from http://energy.gov/sites/prod/files/2014/03/f10/merit08_mccarty_10.pdf

7. Luo, A. (2012, January 1). Magnesium Front End Research and Development (MFERD). Retrieved from http://www1.eere.energy.gov/vehiclesandfuels/pdfs/merit_review_2012/lightweight_ materials/lm008_luo_2012_o.pdf

8. W.A. Monteiro, S.J. Buso and L.V. da Silva (2012). Application of Magnesium Alloys in Transport, New Features on Magnesium Alloys, Dr. Waldemar A. Monteiro (Ed.), ISBN: 978-953-51-0668-5, InTech, DOI: 10.5772/48273. Available from: http://www.intechopen.com/books/new-features-on-magnesium-alloys/application-of- magnesium-alloys-in-transport

9. Friedrich, Horst E.; Mordike, Barry L. (2006). Magnesium Technology - Metallurgy, Design Data, Applications. (pp: 2). Springer - Verlag. Retrieved from http://www.knovel.com/web/portal/browse/display?_EXT_KNOVEL_DISPLAY_bo okid=1284&VerticalID=0

10. Song, Guang-Ling (2011). Corrosion of Magnesium Alloys. (pp: 3-4). Woodhead Publishing. Online version available at:

11. "Melting, Alloying and Refining." Magnesium Technology. Springer Berlin Heidelberg, 2006. 109-43. Print.

12. Ralston, KD, G Williams, and Nick Birbilis. "Effect of Ph on the Grain Size Dependence of Magnesium Corrosion." Corrosion 68.6 (2012): 507-17. Print.

13. Bohni, H. "Localized Corrosion of Passive Metals." Uhlig's Corrosion Handbook, Second Edition, Ed. RW Revie, John Wiley and Sons, Inc., New York (2000): 173-90. Print.

14. Cao, Fuyong, et al. "Corrosion of Ultra-High-Purity Mg in 3.5% Nacl Solution Saturated with Mg(OH)2." Corrosion Science 75.0 (2013): 78-99. Print.

15. CORROSION: GALVANIC CORROSION. (n.d.). Retrieved from http://www.ssina.com/corrosion/galvanic.html

16. http://www.cartech.com/techarticles.aspx?id=1566

17. http://www.delstar.com/passivating/stainless-steel-passivation.html

18. http://www.euro-inox.org/pdf/map/Passivating_Pickling_EN.pdf

19. Frankel, Gerald S. "Techniques for Corrosion Quantification." (2002). Print.

20. C.A. Matzdorf, W.C. Nickerson, B.C. Rincon Troconis, G.S. Frankel, Longfei Li, and R.G. Buchheit, “Galvanic Test Panels for Accelerated Corrosion Testing of Coated Al Alloys, Part 1: Concept,” Corrosion 2013 69:12, 1240-1246

21. Feng, Zhicao, and G. S. Frankel. "Galvanic Test Panels for Accelerated Corrosion Testing of Coated Al Alloys: Part 2—Measurement of Galvanic Interaction." Corrosion 70.1 (2013): 95-106. Print.

22. Friedrich, Horst E, and Barry L Mordike. "Corrosion and Surface Protections." Magnesium Technology: Metallurgy, Design Data, Applications (2006): 431-97. Print.

23. Pardo, A, et al. "Corrosion Protection of Mg/Al Alloys by Thermal Sprayed Aluminium Coatings." Applied Surface Science 255.15 (2009): 6968-77. Print. 20

24. Corrosion protection of Steel. (2011, January 1). Retrieved from http://www.keytometals.com/page.aspx?ID=CheckArticle&site=kts&NM=300

25. Surface Treatment of Aluminum and Aluminum Alloys. (2003, January 1). Retrieved from http://www.keytometals.com/page.aspx?ID=CheckArticle&site=ktn&NM=67

26. Corrosion of Magnesium and Magnesium Alloys. (n.d.). Retrieved from http://www.keytometals.com/Article19.htm

27. "Corrosion Behaviour of Magnesium/Aluminium Alloys in 3.5 Wt.% Nacl." Corrosion Science 50.3 (2008): 823-34. Print.

28. Senf, J.: Broszeit, E., M. Gugau, and C. Berger. "Corrosion and Galvanic Corrosion of Die Casted Magnesium Alloys." Magnesium Technology 2000. Warrendale, PA: The Minerals, Metals, and Materials Society, 2000. 137-42. Print.

29. Jia, J. X. and Atrens A. and Song G. and Muster T. H. "Simulation of Galvanic Corrosion of Magnesium Coupled to a Steel Fastener in NaCl Solution." Materials and Corrosion 56.7 (2005): 468--74. Print.

30. Williams, G., N. Birbilis, and H. N. McMurray. "The Source of Hydrogen Evolved from a Magnesium Anode." Electrochemistry Communications 36.0 (2013): 1-5. Print.

31. Pushing the boundaries of riveting technology. (n.d.). Retrieved from http://www.henrob.com/GB/automotive.php

32. What is HSPR? (2015, January 1). Retrieved from http://www.henrob.com/GB/what- is-hspr.php#

33. Luo, H. and Dong C. F. and Liu Z. Y. and Maha M. T. J. and Li X. G. "Characterization of Hydrogen Charging of 2205 Duplex Stainless Steel and Its Correlation with Hydrogen-Induced Cracking." Materials and Corrosion 64.1 (2013): 26--33. Print.

34. Frankel, G. S., A. Samaniego, and N. Birbilis. "Evolution of Hydrogen at Dissolving Magnesium Surfaces." Corrosion Science 70.0 (2013): 104-11. Print.

35. Seok-Jae Lee and Joseph, A. Ronevich and George Krauss and David K. Matlock. "Hydrogen Embrittlement of Hardened Low-Carbon Sheet Steel." Isij International 50 (2010): 294--301. Print.

36. Evers, Stefan, Ceylan Senöz, and Michael Rohwerder. "Hydrogen Detection in Metals: A Review and Introduction of a Kelvin Probe Approach." Science and Technology of Advanced Materials 14.1 (2013): 014201. Print.

37. Ferreira, PJ, IM Robertson, and HK Birnbaum. "Hydrogen Effects on the Interaction between Dislocations." Acta Materialia 46.5 (1998): 1749-57. Print.

38. Koutsi, H, et al. "Corrosion-Induced Hydrogen Embrittlement in Aluminum Alloy 2024." Corrosion Science 48.5 (2006): 1209-24. Print.

39. Kappes, Mariano, Mariano Iannuzzi, and Ricardo M Carranza. "Hydrogen Embrittlement of Magnesium and Magnesium Alloys: A Review." Journal of The Electrochemical Society 160.4 (2013): C168-C78. Print.

40. Robinson, MJ, and PJ Kilgallon. "Hydrogen Embrittlement of Cathodically Protected High-Strength, Low-Alloy Steels Exposed to Sulfate-Reducing Bacteria." Corrosion 50.8 (1994): 626-35. Print.

41. G.G. Perrault, The potential-pH diagram of the magnesium-water system, Journal of Electroanalytical Chemistry and Interfacial Electrochemistry, Volume 51, Issue 1, 25 March 1974, Pages 107-119, ISSN 0022-0728, http://dx.doi.org/10.1016/S0022- 0728(74)80298-6.

42. G.G. Perrault, Encyclopedia of Electrochemistry of the Elements vol. VIII, A.J. Bard ed, Marcel Dekker N.Y., (1978)

43. Marine Cathodic Protection. (n.d.). Retrieved January 7, 2015, from http://www.nstcenter.biz/writeup.aspx?title=Marine Cathodic Protection&page=NavyCommunityAppExtShipCathodicProtection.html

44. Magnesium Alloys Encyclopedia: Corrosion Behavior. (n.d.). Retrieved January 7, 2015, from http://www.magnesium.com/w3/data-bank/index.php?mgw=217

Chapter 2: AZ31 Cleaning Procedure for MagPASS Conversion Coating Introduction

USAMP found from the Magnesium Front End Research and Development

(MFERD) Project, Phase II demonstration structure that the MagPass® conversion

coating “could not be effected” on the warm-formed AZ31B-H24 magnesium sheet alloy acquired from Magnesium Elektron of North America (MENA).1 The sheet material was

provided in the “mill finished” condition, which exhibited a dark gray or blackened surface discoloration. It was speculated this surface condition could result from a combination of the following: 1) Oxidation of the surface during the direct chill (DC) slabbing and sheet-forming process; 2) Carbonization of residual surface lubricant during the warm-forming temperature excursion.

The warm-forming operation was performed to produce the upper rail halves for the demo structure. This procedure involved exposing the material to 300-350°C for 10-

15 minutes in the furnace before the forming step. Prior to warm-forming there was no attempt to remove the mill finish condition, including any residual lubricant. MENA recommended a cleaning procedure to prepare the mill-finished surface for MagPass®

(license provided to Luke Engineering, Wadsworth, OH via AHC Oberflachentechnik,

Germany); the details are in the Experimental Procedures section below.

The oxidation and temperature excursion of the mill-finished sheet during

warming resulted in a “carbonized” lubricant layer that rendered the de-oxidation step of

the MagPass® conversion coating process ineffective and then in “smutting,” which 23

appeared during the post de-ox stop-etch step. This smut film prevented uptake of a

subsequent active-layer deposition (permanganate-based chemical bath with inhibitor

such as vanadate). Prior attempts to remove smut from the sheet material at Luke

Engineering showed the layer could only be removed mechanically, for example by

brushing, and that it was not water-soluble. These results were confirmed by preliminary experiments in our lab.

The experiments undertaken for this work originally planned to apply the

MagPass® conversion coating at Luke to samples specially cleaned at OSU. However, the

smut-containing surface was understood not to be suitable for uptake of the active

compounds in the MagPass® process. Therefore, the goal of this work was to confirm

(visually or otherwise) the presence of smut, as was described by engineers at Luke, on

the Mg surface.

Experimental Procedures

Coupons of various surface conditions were exposed to the MENA-prescribed cleaning procedure. To test the hypothesis that the warm-forming procedure exacerbates a tenacious carbonaceous layer, the following experiments were performed:

• Attempt to reproduce the warm-formed surface condition by exposing as-received

AZ31B-H24 sheet material from MENA to temperatures of 300 or 350°C for 15

minutes.

• Process coupon matrix with cleaning procedure as prescribed by MENA.

• Use visual observation and Fourier Transform Infrared Spectroscopy (FTIR) to

evaluate the surface condition prior to and immediately following cleaning.

More elaborate analytical techniques besides FTIR were not investigated in this study

primarily due to project considerations that both warm-formed AZ31B-H24 magnesium

alloy and the MagPass® conversion coating process would not be used in construction of

the Phase III demonstration structures, so that a prolonged study of this phenomenon was

not warranted, other than for closure in understanding the nature of the process failure.

Furthermore, most analytical techniques are generally expensive and

technologically cumbersome. One of the motivations of this research was to elucidate the

viability of a simple, quick quality assurance/quality control (QA/QC) procedure that

could be implemented by a technician with limited technical background. Other

procedures such as Auger electron spectroscopy and X-ray diffraction generally require

expensive equipment and longer tests, whereas handheld FTIR devices are readily

commercially available.

FTIR is a vibrational spectroscopy which provides a fingerprint of the molecular

bonding of a substance by measuring its characteristic IR absorption frequencies. In this

case study, the absorption spectrum was used to gain information about an organic

surface contaminant. The evolution of the surface contaminant was monitored throughout

the cleaning and warm-forming simulation process in order to assess what contaminants might be present on the surface. These contaminants were suspected to subsequently affect the integrity of the MagPass® conversion coating.

The surface contaminant was, in fact, a proprietary MENA lubricant residual from

the sheet-rolling process. AHC, the licensor of MagPass®, expressed concerns upon visual inspection of the MENA mill-finish condition, apparently having encountered

problems with MagPass® uptake on this type of surface in the past. The main limitation

of FTIR that was experienced in this case study was difficulty in analytically quantifying

the exact composition of the surface contaminant without some knowledge of the original

composition of the rolling lubricant.

Although FTIR spectra are characteristic of specific molecular bonds, there are many bonding possibilities at any specific position along the infrared continuum

(spanning wave numbers from 400-4000 cm-1). Without some prior knowledge of

chemical composition, it is only possible to monitor bonding changes qualitatively,

confirm whether a substance is present, and determine if it has been removed by cleaning.

Thus, unless information is provided as to the original chemical composition of the object

being analyzed, FTIR falls short of providing a truly quantitative method. Otherwise, a

handheld FTIR device meets the aforementioned requirements as a tool to assess the

surface condition of metal sheet material with organic contamination.

Coupons were defined as follows:

• “Pristine” AZ31: as-received via Luke Engineering from Salzgitter Magnesium-

Technologie GmbH (SZMT) used as a control for application of MagPass® without

cleaning should the experiment progress to conversion coating. The Saltzgitter AZ31

magnesium was highly reflective and showed no visual evidence of the darkening

seen on the MENA mill finished AZ31 magnesium sheet.

• MENA **AZ31B-H24 mill-finished: as-received condition sent from Bruce Davis at

MENA and received by OSU 22 July, 2013.

• MENA AZ31B-H24 upper rail component: warm-formed by USAMP contractor,

Troy Tooling, received 15 January, 2013. This material was from a different lot than

the as-received MENA AZ31B-H24

• Warm-forming simulation: as-received MENA **AZ31B-H24 mill-finished coupons

exposed to 300 or 350°C to model temperature excursion of the warm-forming

process. These temperatures span the range used during the warm-forming procedure.

Both temperature excursions lasted for 15 minutes, in air, in a ceramic furnace.

**AZ31B-H24 mill-finished sheet from MENA definition printed on one of the sheets as: MAGNESIUM SHEET COIL AZ31B H24 ASTM B90/98 LOT 840001

GAUGE 0.079

Coupons were prepared to the dimensions 10.2cm×15.2cm by 2mm in thickness. A

6.35mm diameter hole was drilled in the top center for use during dip treatment.

Coupons were processed by two cleaning procedures, the difference being the addition of a glycolic nitrate pre-etch step. The procedure recommended by MENA, which was adapted to the lab-bench (Figure 3), is summarized in Figure 4 and in the following:

1. Mild alkali cleaning

a. Boiling 2.5% Na3PO412H20 sodium phosphate tri-basic 12-H for 5-15 minutes

2. DI water rinse

3. Glycolic (acid) nitrate pre-etch pickle (nonstandard step added in some tests)

a. 20 vol% of 70% glycolic acid with 5 vol% sodium nitrate (NaNO3) - room

temperature – 30s or more if needed

4. Phosphoric acid etch

a. 75% acid (reagent stock) room temperature – 30s or more as needed

5. Alkali stop-etch

a. 10% NaOH solution – room temperature – dip

6. DI water rinse

7. Forced/heated air dry (only forced air was used in this experiment, however, drying

was virtually immediate under the high pressure air flow)

After cleaning, coupons were inspected visually for the presence of smut and sealed

in poly-bags upon drying. The coupons were then analyzed with a Thermo Nicolet FTIR-

ATR at a resolution setting of 8 and 100 scan iterations. FTIR scans were collected to

monitor the evolution of the surface compounds during various steps of the procedure

(Figure 4).

Results and Analysis

Prior to and after coupons were processed through the cleaning procedure, they

were subjected to visual inspection and FTIR analysis. Prior to cleaning, the mill-finished

material exhibited a dull, blackened surface finish (Figure 5) – this is typical of MENA

mill-finished AZ31B-H24 magnesium alloy. FTIR of the as-received mill-finish and as-

received warm-formed upper rail coupons prior to cleaning revealed absorption peaks

characteristic3,4 of carbon-hydrogen (C-H) bond bending (735-770 cm-1), indicating the

presence of a hydrocarbon compound, and carbon-carbon (C-C) aromatic bond stretching 28

(1500-1600 cm-1) (Figure 6). The C-H bond bending absorption peak is possibly due to

the presence of a di-alkyl ortho arene compound (R2C6H4). Magnesium could be bonding

with this hydrocarbon electrophile through a nucleophilic reaction called the Grignard

Reaction.5-8 Heat treating may further exacerbate the extent of this bonding. The peak

centered at 2350 cm-1, visible in all spectra, indicates the presence of carbon dioxide and

is simply an artifact picked up by testing in air.3,4 Without knowledge of the original

composition of the lubricant, it is not possible to quantify exactly what chemical

interactions are occurring between the lubricant and the magnesium alloy.

Simulation of the warm-forming procedure

Both coupons were slightly more reflective and lighter in color after heat

treatment. FTIR of the coupons revealed an absorption profile similar to the as-received

warm-formed upper rail coupon for the 300°C sample (Figure 7). The 350°C coupon also

had similar peaks3,4, but the intensity of the C-H bond bending (735-770 cm-1) peak was

substantially higher (Figure 8). This could be due to the carbonation being exacerbated by the higher temperature and/or enhanced specular reflectance of the heat-treated coupon.

Based on visual assessment and FTIR analysis, the temperature excursion experiment approximately reproduced the surface condition of the warm-formed upper rail component. The differences can probably be attributed to the age of the upper rail component. It was in storage for an unknown length of time in a facility of naturally varying temperature and humidity, whereas the mill-finished sheet from MENA had been produced much more recently. The natural ageing of the warm-formed component may

have caused thickening of the oxide film or some further complexation of the surface

chemistry due to ambient exposure.

Cleaning and pretreatment

Coupons of as-received mill-finished, as-received warm-formed upper rail, as-

received “pristine” from Salzgitter, and warm-form simulated (300 and 350°C) samples

were processed via the two cleaning procedures outlined above. FTIR spectra of the coupons after cleaning were very similar for all samples. The C-H bending signal was reduced in intensity, but the peak was broadened, indicating the possible complexation in solution of the carbonaceous surface compound. It is suspected the carbonaceous oxide scale was etched off the surface, dissociated in solution, but re-deposited on the Mg surface as a further complexed mixture of carbonaceous “smut.” An additional high- intensity peak appeared indicating phosphorous and oxygen (P=O) double bonds (1140-

1320 cm-1); 3,4 this is a product of the deposition of a phosphate structure, which is a

necessary pretreatment for uptake of active ions in a subsequent permanganate bath.

Addition of the glycolic nitrate pre-etch step did not affect the FTIR signal, suggesting it did not improve the smutting problem. Smut was visually apparent on the samples after both cleaning procedures, as well. The Salzgitter AZ31 was also processed through the cleaning procedure to obtain a reference spectrum for a material that was accepting of the MagPass® procedure. Figure 11 shows FTIR spectra comparing the mill finished, warm-formed, and Salzgitter materials after the cleaning procedure. Noticeable differences exist between the spectra of the Salzgitter and other materials. Namely, the difference in absorption intensity that occurred across the entire spectrum, and the

relatively smaller C-H bending signature visible in the Salzgitter spectrum, set it apart.

This reinforces the capability of FTIR to differentiate between surface conditions.

Discussion

The cleaning procedure suggested by MENA to remove the mill-finish from

AZ31B-H24 Mg alloy was attempted and found to be ineffective in reducing or eliminating a smutting reaction accompanying the phosphoric acid deoxidation and

NaOH “stop,” which are precursors to the incorporation of the active conversion coating anions of the MagPass® process. The remnant reacted layers on the surface of the mill- finished and warm-formed AZ31 magnesium sheet were of a carbonaceous nature, and were changed but not removed by the cleaning procedure.

The addition of a glycolic nitrate pre-etch step did not affect the final surface chemistry. FTIR, performed after cleaning, revealed a broadening of the peak indicative of hydrocarbon bonding. It is suspected that the carbonaceous surface oxide dissolved and dissociated in solution, only to be re-deposited on the surface of the metal in a complexed form. Some variables that have a major influence on the effectiveness of any metal cleaning procedure are pH, concentration of dissolved metal ions (Mg2+ in this

case), temperature, etchant chemistry, contamination, exposure time, convection by gases

evolved from metal surface, etc. Failure to investigate the interaction of any or all of

these variables could limit the effectiveness of the cleaning procedure – protecting

magnesium alloys from corrosion is a complicated problem that has yet to be investigated

thoroughly.

Ionic magnesium is extremely reactive and its compounds and complexes generally have low solubility in solution, which implies re-deposition is a possibility.

FTIR is a useful technology for characterizing molecular bonding, but some information as to the chemical nature of the substance must already be known, or it must be used in conjunction with some other analysis technique. It does, however, provide the capability to verify the presence of a surface compound, albeit of unknown overall composition.

Conclusions

• Cleaning procedure suggested by MENA was found to be ineffective.

• Remnant (after warm-forming and cleaning procedure) reacted layers of mill-finished

surface were shown by FTIR to be carbonaceous in nature.

• Addition of glycolic nitrate pre-etch step did not affect the surface chemistry after

cleaning

• FTIR is a quick, simple, and relatively cheap method for characterizing molecular

bonding and organic surface compounds on metals. However, information as to the

general compound composition is necessary for it to be a quantitative as hundreds of

thousands of organic compounds reside within a range of less than 4000 wave

numbers. This may lead to false-positives in results.

References

1. U.S. Department of Energy. USAMP. Task 5 - Corrosion and Surface Treatment: Statement of Work. Office of Energy Efficiency and Renewable Energy Vehicles Technologies Program: Award: DE-EE0005660, 2012.

2. Griffiths, Peter, and James Haseth. Fourier Transform Infrared Spectrometry. 2nd. Hoboken: John Wiley & Sons, 2007. Print.

3. http://www2.ups.edu/faculty/hanson/Spectroscopy/IR/IRfrequencies.html

4. http://ftirsearch.com/default2.htm

5. Smith, Michael B., and Jerry March. March's Advanced Organic Chemistry. 6th. Hoboken: John Wile & Sons, 2007. 1300. eBook.

6. http://www.organic-chemistry.org/namedreactions/grignard-reaction.shtm

7. http://www.britannica.com/EBchecked/topic/246137/Grignard-reagent

8. http://chemistry2.csudh.edu/rpendarvis/grignard.html

Figures

Figure 3: Experimental setup showing 2-liter beakers with solution – omitting the glycolic step here (top) and coupons prior to being cleaned (bottom).

Figure 4: Flow chart of coupon treatment schedule including the step to simulate the warm-formed condition.

Figure 5: as-received AZ31B-H24 from MENA with “mill-finish” surface condition.

Coupon on left was received about 6 months prior to that on right. Material on right was used to acquire data for this report, as its processing history and shelf life was known, while that on the left had an unknown history.

Figure 6:. FTIR spectra of as-received mill-finished (MF) and as-received warm-formed

(WF) upper rail coupons prior to cleaning. Boxes indicate absorption signal characteristic of C-H bond bending (735-770 cm-1), C-C aromatic bond stretching (1500-1600 cm-1)

-1 and CO2 in gas phase (2350 cm ).

-1 -1 (1500-1600 cm ) and CO2 in gas phase (2350 cm )

Figure 9: FTIR spectra comparing as-received mill finished, as-received warm-formed

upper, warm-form simulated at 300 and 350°C after exposure to phosphoric acid etching

cleaning procedure. Boxes indicate absorption signal characteristic of C-H bond bending

(735-770 cm-1), P=O stretching (1140-1320 cm-1), C-C aromatic bond stretching (1500-

-1 -1 1600 cm ) and CO2 in gas phase (2350 cm ).

Figure 10: FTIR spectra comparing as-received mill finished, as-received warm-formed upper, warm-form simulated at 350°C after exposure to glycolic nitrate pre-etch and phosphoric acid etch cleaning procedure. Boxes indicate absorption signal characteristic of C-H bond bending (735-770 cm-1), P=O stretching (1140-1320 cm-1), C-C aromatic

-1 -1 bond stretching (1500-1600 cm ) and CO2 in gas phase (2350 cm ).

Figure 11: FTIR spectra comparing as-received mill finished, as-received warm-formed

upper, warm-form simulated at 300 or 350°C, and Salzgitter “pristine” AZ31 after

exposure to glycolic nitrate pre-etch and phosphoric acid etch cleaning procedure. Boxes

indicate absorption signal characteristic of C-H bond bending (735-770 cm-1), P=O

-1 -1 stretching (1140-1320 cm ), C-C aromatic bond stretching (1500-1600 cm ) and CO2 in gas phase (2350 cm-1). Notable differences exist between Salzgitter and other AZ31.

Chapter 3: Hydrogen Embrittlement of Hardened Steel Rivets Introduction

Steel components are susceptible to hydrogen embrittlement when hardened.

According to Fastenal, steels below HRC 35 on the Rockwell Hardness scale are generally not susceptible to hydrogen embrittlement. Steels above HRC 40, however, are at risk of embrittlement by environmental hydrogen.1 Unfortunately, knowledge of this phenomenon is not emphasized in many engineering disciplines, leading to unexpected, dangerous and costly failures. The risk of hydrogen embrittlement of hardened steel

Henrob Self-Pierce Rivets (SPRs) used in joining automotive sheet material was investigated in this work.

Self-pierce rivets have been rapidly adopted by the automotive industry as a cheap and rapid sheet-metal joining technology.2 One of the advantages of the technology is its capability to join dissimilar materials. However, the piercing action of the rivets makes it impossible to electrically isolate the joined materials. Thus, SPRs can introduce a galvanic corrosion risk if, for example, aluminum or magnesium sheet is joined with steel rivets. Within the scope of the USAMP project, steel SPRs are being investigated for joining alloys of aluminum to aluminum as well as aluminum to magnesium. In any arrangement, galvanic corrosion is a concern because steel rivets are electrochemically cathodic to alloys of Mg and Al.

The risk of embrittlement exists because the rivets were found to have been hardened to HRC 47, which is in the range of susceptibility to hydrogen embrittlement of 43 steel. The rivet steel is ferritic-martensitic in nature. Figure 12 is a micrograph of the rivet-steel microstructure; the dark phases are martensite and the lighter phases are ferrite. As the cathode in a galvanic couple with either Al or Mg, the steel rivets will be the site of cathodic reactions, which might include hydrogen evolution. Furthermore, anodically corroding aluminum and magnesium tend to generate a copious amount of hydrogen gas even during anodic dissolution.3 Thus, if the inherently electrically connected bi- or tri-metal joints are ionically connected by mutual exposure to a corrosive environment, hydrogen should be generated immediately proximal to the hardened steel rivets.

In this work, samples were tested mechanically after cathodic charging or exposure to B117 salt spray while galvanically connected to AM60B magnesium plate material.

Experimental Procedures

Rivets

The SPRs were 8 mm in height and 5 mm in diameter with a countersunk head.

They were supplied by Henrob Corporation in uncoated and coated conditions. The base metal was AISI 10B37 hardened steel. Two coating options were investigated: a barrel electroplated 70-30 Zn-Sn alloy, and ion-vapor assist deposited aluminum.

An experiment was designed to induce a “worst-case scenario” hydrogen embrittlement effect in the SPRs. It was attempted to saturate the steel rivets with hydrogen to force embrittlement.

Hydrogen Charging

Hydrogen ingress into the base metal was promoted by cathodically polarizing the

rivets. Electrical connection was made to uncoated rivets by spot-welding a steel wire to the sidewall of the rivets. Spot welding to the coated rivets was problematic so connection was made by drilling a very small hole in the base and inserting a steel wire painted with conductive silver paint. The wire was protected with non-reactive epoxy.

2 Charging was performed at 1 mA/cm in 500 mL of 1 N H2SO4 + 0.25 g/L As2O3, which

is a hydrogen recombination poison that promotes entry of atomic hydrogen into the

metal.4 Hydrogen charging time was estimated by the basic diffusion distance equation:

< >= (1) 2 where is𝑥𝑥 the diffusion𝐷𝐷𝑡𝑡 distance into the rivet, D is the hydrogen diffusion coefficient in

steel and𝑥𝑥 t is time. The diffusion distance was estimated to be 1 mm, as the SPR sidewalls are about 2 mm thick, and hydrogen entered from both sides during charging. depends on microstructure and was estimated to be 10-7 cm2/sec. This calculation yielded𝐷𝐷 about 28

hours for through-thickness hydrogen ingress. Hydrogen charging was carried out as outlined below for at least this long.

Galvanic Corrosion Charging

Rivets were force-fitted into pre-drilled holes in uncoated AM60B magnesium plate material 3 mm in thickness. The holes had a diameter about 50 µm smaller than the rivet diameter to ensure electrical connection during testing. The assembly was then placed in B117 salt spray for 7 days to expose the rivets to the hydrogen evolved during magnesium corrosion.

Explanation of mechanical test

Typically the mechanical effects of hydrogen embrittlement are investigated by a standardized test such slow strain rate testing. However, the SPR geometry is not conducive to creating a standardized specimen. Thus, a unique rivet punch tool was designed to assess the hoop stress capacity of the rivets. A computer controlled hydraulic load frame was used to impose a slow strain-rate on a highly polished tungsten carbide sphere that was pushed into the open cylindrical end of the rivets. The downward force from the upper punch was translated through the sphere into roughly radially-oriented tensile hoop stress in the rivet. The diameter of the sphere was chosen such that it caused the sidewalls of the rivet to flay radially outward. The stress state during the deformation test was complex and dynamic. The rivet eventually failed by splitting, which then resulted in irreproducible applied stresses.

Fixtures

The fixtures to hold the rivets and sphere in place were machined from D2 tool steel, and customized to fit an 80kN MTS Landmark Servohydraulic Test Frame. Figure

13 is a schematic of the design, a photograph of which is Figure 14. To avoid deformation of the sphere by the rivet, a 7 mm diameter, HRC 77 tungsten carbide sphere was used to contact the rivet (Figure 15). This size sphere was chosen by trial and error of several diameters. Its use resulted in consistent data up to the point of the initial fracture of the rivet, and there was no evident plastic deformation of the sphere by the rivet. A circular depression was machined into the lower fixture for snug seating of the rivet base, which prevented it from sliding during deformation. A conical pocket was machined into

46 the center of the upper punch to constrain the sphere from lateral slippage during deformation.

Prior to testing, the load frame was manipulated manually until the components just made contact. Care was taken not to impart force on the rivet prior to running the test protocol. It is also important to start each test from the same position so displacement data are comparable between tests.

Test parameters

A strain-rate-controlled test protocol was programmed on the Landmark Station

Manager software. A strain rate of 0.5 mm/min was imposed over a total displacement of

2 mm, and the force-versus-displacement data were recorded.

Results and Analysis

Figure 16 shows the force versus displacement curves for the different as-received rivets. The test was repeated three times for each condition. For all of these experiments, the force versus displacement curve exhibited three distinct regions of loading behavior.

Region 1 was roughly linear, from zero displacement up to about 0.4 mm. The point of transition between region 1 and 2 corresponded to the displacement at initiation of a fracture in the sidewall of the rivet. Despite its resemblance to the linear-elastic region of a tensile stress-strain curve, this is not the same effect. Elastic strain is exhausted over only a few microns.

Beyond region 1, the utility of the test decreased because the fracture (location, shape, number of fractures, etc.) was not controlled. In standardized tests, such as the

Charpy impact test, for example, the sample is pre-notched to control the location and

type of fracture. The behavior in the 2nd and 3rd regions exhibited less reproducibility than

region 1, and it is not indicative of any obvious trend, expected or otherwise. Therefore,

the analysis will focus on region 1.

The initial displacement behavior was not entirely consistent between samples.

From zero to about 0.05 mm of displacement, a low slope on the force versus displacement curve was exhibited, transitioning sharply into a steeper linear region at about 0.05 mm. This behavior indicates a lower initial loading capacity and is probably due to the fact that the loading in this test initiates at the very end of the rivets, which are tapered to a sharp edge so they can pierce the sheet material being joined. Because of the

reduced cross-sectional area of this tapered portion compared to the full-thickness sidewall, the initial load-bearing capacity is lower than the hoop load capacity of the sidewalls.

Once the sphere engaged the full sidewall, the force-displacement curve transitioned to the maximum slope. This portion of the curve corresponds to plastic yield in the rivet steel as the sphere was pushed into the shank. From about 0.05 to 0.4 mm, the sidewalls were plastically deformed in a complex stress state that was a combination of compressive and radial hoop stresses. There was a significant amount of hoop stress during the test, as the rivets failed by tensile cracking rather than a compressive buckling mechanism. This supports the notion that this test elucidates changes in mechanical strength.

Hydrogen-Charged rivets

Several uncoated, Al-coated, and Sn/Zn-coated rivets were hydrogen charged as

described above and tested with the hoop stress test. Figure 17 and Figure 18 show the

results for the uncoated and Al coated samples.

No noticeable change was observed in the first region (0-0.4 mm displacement) of the curve for either the uncoated or the Al-coated conditions, suggesting that hydrogen charging by cathodic polarization or corrosion exposure did not have a deleterious effect on the hoop-stress loading capacity of the rivets. This was unexpected because of the high hardness of the rivets.

There does seem to be a consistent reduction in properties in Al-coated rivets beyond 0.5 mm of absolute displacement. This may be due to corrosion of the coating that resulted in a reduction in load-bearing cross-sectional area. However, due to the uncontrolled nature of deformation, it may not be possible to draw any conclusions from this portion of the curve.

Hydrogen charging did seem to have an effect on the Sn/Zn-coated rivets. The hoop stress test caused fracture to initiate at a lower stress and smaller absolute displacement in the hydrogen charged samples than for the other conditions, as shown in

Figure 19. However, this is probably not due to hydrogen embrittlement of the steel.

Unwarranted corrosion of the Sn/Zn coating occurred during hydrogen charging. This

“unwarranted corrosion” was caused by the imposition of an anodic current, rather than what was intended to be a cathodic current, and was simply due to operator error. This resulted in an effective reduction in the load-bearing cross-sectional area in the rivet,

which was reflected in the data. It is unclear why the rivets exposed to B117 seemed to

exhibit a significantly higher stress for fracture initiation.

Discussion

It was attempted to induce “worst-case scenario hydrogen embrittlement in hardened steel rivets. In addition to concerns about the high hardness, the rivets will be used to join magnesium and aluminum alloys, which could result in hydrogen generation immediately proximal to the rivet during corrosion. Hydrogen charging was conducted by cathodically polarizing in a solution infused with a hydrogen recombination poison.

Hydrogen charging was also attempted by B117 salt spray exposure of the rivets in intimate electrical and ionic contact with AM60B plate material. A unique test was designed to assess the hoop-stress load-bearing capacity of the rivets and to attempt to evaluate changes in mechanical properties caused by hydrogen uptake.

It can be concluded that the hoop-stress test has limited utility. It was reproducible from 0 to about 0.4 mm of absolute displacement. Fracture initiated in most samples at about 0.4 mm of displacement. There was no control over location, type, or number of fractures, so beyond 0.4 mm of displacement the test loses utility. The strain-rate of 0.2 mm/min may have been too rapid to allow hydrogen to diffuse through the microstructure to interact with dislocations and other traps. Hydrogen embrittlement tests should take hours and even days.5

Analysis of the hoop-stress test data suggests hydrogen charging did not have a

deleterious effect on the hoop-stress load-bearing capacity of uncoated or aluminum

coated rivets. Hoop-stress test data for the Sn/Zn-coated rivets did show a reduction in

stress and displacement to initiate fracture, but this could be due to unwarranted corrosion

of the coating, which resulted in a reduction in the load-bearing cross-sectional area in

the rivet. The unexpected corrosion of the rivets during hydrogen charging was most

likely due to a poor electrical connection at the lead wire/rivet interface. Despite the unwanted corrosion, this does provide a positive result in that it shows the hoop-stress test is sensitive to changes in mechanical properties.

More analysis is required to validate this hoop stress test. Hydrogen diffusivity in steels can vary widely depending on microstructure; it is important to use a representative value in estimating diffusion time. For example, if the diffusivity were one order of magnitude larger, the required charging time would have been 280 hours rather than 28 hours.

In order to increase the utility of the test, it would be necessary to further constrain the deformation of the rivet somehow. Also, it might be interesting to use a pre- notched rivet to have a defined fracture location. The noise in the curves could have been caused by instabilities in the load-cell, and might have been eliminated with better calibration. Finally, the H diffusivity should be known so that full charging of the rivet shank could be assured.

Integration of the area under the force/displacement curve could be relevant to evaluating hydrogen embrittlement. Hydrogen embrittlement tends to reduce the energy required to cause failure. Integration of the area under the force/displacement curve is the way to quantify energy-to-failure. More testing would be required to establish statistical significance, but this could be an interesting future effort.

Conclusions

• Hoop-stress test has limited utility without extensive testing, which is outside the

scope of this project.

• It cannot be safely concluded whether hydrogen charging had an effect on the

mechanical properties of the rivets due to the limited utility of the newly designed

hoop-stress test.

• Rivets were of a non-ideal geometry for effective hydrogen charging and hydrogen

embrittlement testing. It proved to be difficult to effectively polarize the rivets during

hydrogen charging. In fact, it appears as though the Sn/Zn-coated rivets were not

polarized at all during charging. This is probably due to a poor electrical connection

between the potentiostat electrodes and the rivet.

• More effective hydrogen embrittlement testing should be conducted on the rivet-steel

itself with a standardized and established method. A lower strain-rate should also be

imposed.

References

1. https://www.fastenal.com/content/feds/pdf/Article%20-%20Embrittlement.pdf

2. http://www.henrob.com/GB/automotive.php

3. A. Samaniego, N. Birbilis, X. Xia, and G.S. Frankel (2014) Hydrogen Evolution During Anodic Polarization of Mg Alloyed with Li, Ca, or Fe. Corrosion In-Press.

4. Luo, H., C.F. Dong, Z.Y. Lui, M.T.J. Maha, and X.G. Li. "Characterization of hydrogen charging of 2205 duplex stainless steel and its correlation with hydrogen- induced cracking." Materials and Corrosion 64: 26-33. Print.

Figures

Figure 12: Microstructure of rivet steel. The darkened phase is martensite and lighter phase is ferrite.

Figure 13: Schematic of the hoop-stress test punch tool design.

Figure 14: Photograph of rivet press tool.

Figure 15: Photograph of rivet press tool in load frame, ready to be compressed.

Figure 16: Force versus displacement curves generated on as-received rivets by rivet punch test. Three distinct regions are separated by dashed lines and numbered.

Figure 17: Hoop-stress test of uncoated rivets in as-received, H-charged, and corrosion- charged conditions.

Figure 18: Hoop-stress test of Al-coated rivets in as-received, H-charged, and corrosion- charged conditions.

Figure 19: Hoop-stress test of Sn/Zn-coated rivets in as-received, H-charged, and corrosion-charged conditions

Chapter 4: Effect of Corrosion on Lap-Shear Strength of AM60B Magnesium to 6082-T4 Aluminum Joints formed by Self-Pierce Steel Rivets Introduction

Reducing mass has proven to be crucial to improving automobile fuel efficiency.

For example, “cutting 1 kg (24.2 lb) from each of the 70 million light-vehicle engines produced in 2011 could save up to 908 million liters (240 million gallons) of refined fuel or nine million barrels of crude.”1 In 2012, the U.S. Federal Government announced aggressive corporate average fuel economy (CAFE) standards, imposing the requirement that automakers’ fleet average carbon emissions be reduced to 163 grams of carbon per mile by 2025. This translates to 54.5 miles per gallon (mpg) of gasoline in modern combustion motors. These standards must be met for the automaker to continue operating in the United States.2

The automotive industry is facing a major challenge. Consumers are willing to buy fuel-efficient cars, but are not willing to sacrifice comfort, space, performance, safety, or endure major price increases. Even switching to fully electric motors encompasses substantial carbon emissions due to the electricity being generated by coal and natural gas-fired power plants. By reducing the mass propelled by the motor, carbon emissions can be directly reduced. One route to mass reduction is to replace dense metals, like steels, with lighter ones, like aluminum and magnesium alloys.

Aluminum alloys are actively being considered, but the use of magnesium and its alloys as structural automotive materials has been limited by many factors including cost,

formability, durability, strength, and corrodibility. The aforementioned future federal fuel

standards have generated a renewed effort to incorporate magnesium alloys in

automobiles. Under the support of the United States Automotive Materials Partnership

(USAMP), Ohio State is exploring the corrodibility of magnesium alloys.

In this work, the galvanic interaction of mixed-metal lay-ups of cast AM60B magnesium alloy plate joined with steel rivets to 6082 Al alloy sheet has been investigated. The following focus points have been investigated:

Galvanic corrosion of the riveted structure has been investigated by exposing mixed-metal lay-ups to ASTM B117 salt spray for 1, 2, or 3 weeks. The samples were

then pulled in lap-shear to monitor changes to joint strength as a function of exposure

time. The corrosion phenomenology and failure mechanisms have been assessed. A

limited number of samples were exposed; a larger number of uncoated samples were

available to be pulled in lap-shear to establish a statistically significant baseline value.

Exposing joints consisting of material combinations of interest to ASTM B117

salt-spray and then measuring changes to the lap-shear strength over time is a practical materials selection method. In this work, ASTM B117 exposure is used in conjunction with lap-shear testing to discriminate between an industry-standard Sn/Zn and a novel

ion-vapor-deposited aluminum (IVD Al) rivet coatings.

Several mixed-metal lay-ups were provided to OSU with an additional adhesive

between the lap-region of the aluminum and magnesium sheets. Despite the electrical

connection between Al and Mg due to the rivet, it is possible that benefit (i.e. water

ingress prevention, isolation, etc.) is gained by using the adhesive.

Rivet coating

A barrel-plated 70-30 %wt Zn-Sn coating is an industry standard metallic rivet coating for steel rivets, when used to join non-ferrous metals. The coating is designed to depress the more noble electrochemical potential of the rivet, bringing it closer to that of the more active adjacent material. Coating steel rivets is a cost-effective method to take advantage of the electrochemical nature of metals that are typically too expensive to justify using for the entire rivet. Ion-vapor-deposited (IVD) aluminum coatings have been suggested to provide a better electrochemical buffer than Zn-Sn for rivets used to join aluminum and magnesium alloys.

Sheet-metal Lay-up Material Selection

For coupon-level investigation of corrosion phenomenology in mixed-metal systems, sheet metal “lay-ups” were fabricated by Ford Motor Company (Figure 1 -

Appendix). The lay-ups considered in this experiment consisted of a die-cast AM60B magnesium alloy, joined with steel self-pierce rivets, to 6082 aluminum alloy sheet.

Table 2 gives the composition of the alloys.

In Phase I and Phase II of the USAMP project, several combinations of metal pretreatments, rivet coatings, and top-coats were investigated by the USAMP project team. The pretreatment of choice for the USAMP coupons was Alodine 5200. Alodine

5200 (also known as BONDERITE M-NT 5200) is a chromium-free product of Henkel

Corporation, specifically formulated for treatment of non-ferrous alloys. The pretreatment converts the metal surface into an ideal base for adhesives or organic

finishes.6 It was optionally applied immediately preceding and/or immediately following

the self-pierce riveting procedure.

Material electrochemistry

The electrochemical potential and galvanic coupling current of the AM60B Mg

plate and the uncoated and coated rivet material were measured in 3.5% NaCl by Wilson

et al. at North Dakota State University7. The aluminum sheet does not significantly affect the electrochemistry of the tri-metal couple, as the relatively cathodic nature of the steel and anodic nature of the AM60B will dominate the driving forces for corrosion.

The Zn/Sn-coated rivet had a corrosion potential (Ecorr) of about -1 VSCE in 3.5%

7 NaCl, compared to an Ecorr of about -1.55 VSCE for the AM60B. The IVD aluminum-

coated rivets exhibited an Ecorr of about -680 mVSCE, which was only slightly lower than

the uncoated rivet Ecorr of about -620 mVSCE. Although the kinetics, and therefore the

practical rate of corrosion, cannot be determined by this information, this is indicative of

the driving force for corrosion. This would suggest the driving force for corrosion will be

smaller between the Zn/Sn-coated rivets and the AM60B plate than between the IVD Al-

coated rivets and the AM60B.

The galvanic coupling current in 3.5% NaCl is a measure of the corrosion rate of

the magnesium when the rivets are electrically connected to the AM60B. The coupling

current between the IVD Al-coated rivet and AM60B was very similar to that between

uncoated rivets and the AM60B.7 The galvanic coupling current between Zn/Sn-coated

rivets and AM60B was about half of that which occurs between the IVD Al-coated rivets

and AM60B. The full-immersion corrosion potential and galvanic coupling current

presented above both point to the Zn/Sn rivet coating being the superior coating, from an

electrochemical standpoint.

Corrosion Phenomenology of SPR in ASTM B117

It is appropriate to introduce the chemical reactions occurring on the sample

surface during ASTM B117 salt-spray exposure, so the implications can be further

discussed in the results section. As was discussed in the Material Electrochemistry

section, above, broadly speaking, there is a coupling current that occurs between AM60B

magnesium and steel rivets when they are exposed to 3.5% NaCl. The solution used in

ASTM B117 is 5% NaCl, but it should produce a similar situation. The reactions

producing this current result in the dissolution of magnesium near the rivet (Schematic in

appendix). Water reduces on the cathode (rivet), producing hydrogen and hydroxide ions,

accompanied by a local increase in pH. Magnesium near the rivet dissolves, supplying

the electrons consumed by the water reduction reaction. The increase in pH, however,

results in passivation of the magnesium, and thus dissolution occurs slightly farther from the rivet than would be expected by the implications of the IR drop in solution – which, in theory, should result in electrical exchange across the shortest possible distance.

- - 2H2O + 2e = H2 + 2OH (aq)

Mg = Mg2+(aq) + 2e-

The fact that the magnesium dissolution occurs slightly away from the rivet has

significant implications regarding the geometry of the load-bearing material cross- section, and thus the fracture mechanics associated with failure. This will be further discussed in the results section.

Effect of Adhesive

It has already been shown by the USAMP Joining Team that the addition of an

adhesive increases the strength of Al/Mg joints in lap-shear significantly.9 It also

effectively renders the adhesively-bonded samples incomparable to the non-adhesively- bonded samples, at least in-so-far as lap-shear testing is concerned. If the joint could be achieved with only adhesive and no rivet, aluminum-magnesium joints would maintain

high strength, the steel cathode would be eliminated, and the electrical connection

between the sheets would be eliminated. However, if the rivet is to be used, it is of

interest to know whether the cheaper, barrel-plated Zn-Sn coating can be justifiably replaced by the novel IVD aluminum plating technique.

Stress state - adhesive versus non-adhesive bond

The stress evolution during loading and failure initiation locations are consistently different between riveted samples with and without adhesive. The failure will be further discussed in the results section, but the differences in the evolution of the stress state will be discussed here.

Rivet-only Joints

The rivet serves as a stress-concentrator in the joined assembly. Failure in the riveted assemblies is expected to occur in the AM60B magnesium sheet, as it is weaker and less tough than the 6082 aluminum sheet. These characteristics are confirmed by the lower shear strength, ultimate tensile strength, and elongation-to-failure exhibited by

AM60B in Table 3.

The strength of the joint is also expected to decrease with increased B117

exposure time, as the magnesium immediately surrounding the rivet should be

preferentially corroded, reducing the load-bearing cross-sectional area of the sheet.

Rivet & Adhesive Joints

During loading, stress will be distributed across the entire lap-area that is

adhesively bonded. There will not be a stress concentrator until decohesion occurs, thus

corrosion near the rivet should not affect the strength of the joint as dramatically as in the

rivet-only joints.

Lap-shear Testing, Magnesium Failure and Variance Analysis

Variance analysis will be useful for interpreting the results of lap-shear testing

because it allows for the separation and quantification of variance that arises due to non-

random changes in sample populations. In non-statistical terminology – this means that it

allows quantification of the effect of corrosion on lap-shear test data. In this study, that

ability will be used as a materials selection mechanism in order to shed light on the

galvanic interaction between AM60B Mg and the two different metallurgical rivet-

coating options mentioned above.

Analysis of Statistical Variance in Lap-shear testing

Lap-shear testing is proposed in this work to be a materials selection method, even though it met resistance due to the large amount of variance that occurs in the data resulting from the many variables associated with lap-shear testing.8 In order for this to

be a useful materials selection tool, the variables will be economized and quantified.

OriginPro 8 was the statistical software package used for calculations.

The variance analysis is based on an F-ratio (F for Fisher) where the numerator is

the treatment variance, and error variance is the denominator. Thus, the F ratio indicates

the ratio of the treatment variance to the error variance. This is useful, because the

random variation in results that occurs due to the complicated stress state in lap-shear testing can be lumped into “error variance”. The error variance will be determined by lap- shear testing a population of 12 unexposed rivet-only and 12 unexposed rivet & adhesive coupons. Any additional variation in the results will then be assumed to be due to treatment variance. Thus, the F-ratio gives a simple quantified value for the overall performance of the lay-up materials, topcoat, etc.

In this study, two independent populations are investigated: rivet-only and rivet & adhesive samples. Additionally, the rivet-only samples can be further subdivided into two independent populations and designated by either A (Sn/Zn-rivet coating) or B (IVD-Al- rivet coating). Coupons with rivet & adhesive are all designated C. No coupons were fabricated with Sn/Zn-coated rivet & adhesive, only with the IVD-Al coating.

Error Variance in Rivet-only Population

In the rivet-only samples, the “error variance” arises from uncontrollable variation in the way the self-pierce rivet ultimately forms its bond, and also from variations in the cross-sectional area of the lap-shear coupons as they were sectioned by hand with a band- saw. The data are much less sensitive to the latter, as was confirmed by conducting lap- shear tests on rivet-only samples with extremely varying cross-sectional area. The maximum loads from these experiments with varying cross section were in close agreement with the coupons with a regular cross-sectional area.

Error Variance in Rivet + Adhesive Population

The error variance is much more sensitive to differences in band-saw-induced variations in the cross-sectional area in coupons with rivet & adhesive. While the joint is adhesively bonded, the entire bonded surface area is load-bearing, and the load should be

distributed in some way across the surface. Changes anywhere in the cross-sectional –

and thus the bonded - surface area, will have a substantial effect on the load-bearing capacity of the adhesive joint. This is reflected in the results and explains the larger variance in the un-exposed adhesively bonded joints. Random variation in the way the rivet forms the joint also occurs here, but this study is more concerned with the initial decohesion event as it controls the maximum lap-shear stress in coupons with rivet & adhesive. It should be mentioned that there was significant variation measured in the

“secondary” failure of coupons with rivet & adhesive, and that interpretation of the differences could be an interesting future effort.

Treatment Variance

For the rivet-only A and B coupon groups, the treatment variance will be used to determine which rivet-coating performs better in accelerated salt-spray exposure tests.

For the rivet & adhesive C coupon group, the treatment variance will be quantified and

analyzed, but not used for comparison as there are no Sn/Zn-coated-rivet & adhesive coupons.

As an additional screening procedure, the significance of the F-ratio that results from the variance analysis will be tested by a Chi-Squared test. This statistical analysis tool assigns a probability based on a specified confidence interval as to whether two

populations are independent. Thus, if the populations are independent, the error and

treatment variance can be separated, and the F-ratio is significant.

Welch-Aspin Test

This statistical theory will be used in the results section, but is introduced here.

The Welch-Aspin Test can be used to check that two unequal (in number) populations

with equal or unequal variance (in results) have equal means. In other words, this

statistical theory provides the probability that the results being measured are significant,

despite being from a very small population.

Welch’s t-test defines the statistic t as:

= 𝑋𝑋1 − 𝑋𝑋2 + 𝑡𝑡 2 2 𝑠𝑠1 𝑠𝑠2 � 𝑁𝑁1 𝑁𝑁2 Where , , are the sample mean, sample variance and sample size, 2 𝑡𝑡ℎ 𝑖𝑖 𝑖𝑖 𝑖𝑖 respectively. This𝑋𝑋 𝑠𝑠is where𝑁𝑁 the Welch’s𝑖𝑖 t-test departs from the Student’s t-test – the

Student’s t-test’s denominator is based on an estimate of the pooled variance between the

samples, where Welch’s t-test accounts for independent variation. Both the Welch-Aspin

and Student’s t-tests are part of the two-sample t-test function in the OriginPro 8 software

package, which was used for computation in this study.

Experimental Procedures

Sample Preparation

Ford Motor Co. supplied Mg/Al 2-plate coupons, which were sectioned with a

band saw into four individual lap-shear specimens, as shown in Figure 20. The

experimental matrix developed from sectioned samples is shown in Table 4. Lay-ups were provided with either a Sn/Zn or IVD-Al coated rivet. A limited number of uncoated samples were provided in order to establish baseline lap-shear strength values. The uncoated samples were produced only with IVD-Al coated rivets. The rivet coating will not have an effect on baseline lap-shear strengths, so this group was used to provide baseline values for both Sn/Zn-coated and IVD-Al coated sample populations, which were then exposed to ASTM B117. The cut edges on samples with the electro-coat that were to be exposed to B117 were masked with a non-conductive, protective epoxy so edge-corrosion would not occur, the drainage thereof altering the electrolyte chemistry at the rivet/stack-up interface (Figure 21). Uncoated samples are shown in (Figure 22).

ASTM B117 Exposure

Electro-coated samples, namely A4, A7, B22, B24, C22, C23, C27, and C30, were exposed to ASTM B117.6 Table 5 details the exposure schedule for the sectioned

samples. Upon removal from the chamber, corrosion product was removed from samples

by washing with DI water and light brushing. Samples were then immediately dried by

forced-air for several minutes and stored in a desiccator to limit additional corrosion.

Lap-shear testing

Lap-shear tests were conducted on an 80kN MTS Landmark Servohydraulic load

frame. A 0.2 mm/min crosshead speed was used to pull samples to failure in tension.

Figure 23 is a schematic of the lap-shear setup. Shims were implemented to reduce

bending during the test.

Baseline testing

Samples B155 through B160 and C2 through C7 were provided with no e-coat,

were not exposed to B117, and were thus used to establish a statistical baseline strength

value, and to quantify the error variation. Lay-ups were sectioned as described above.

Two samples of each lay-up were pulled in lap-shear to establish a baseline value for samples with and without adhesive.

Results and Analysis

Failure in all but one coupon ultimately occurred by brittle fracture in the

AM60B. One coupon failed by a fracture of the rivet. Corrosion damage, which caused extensive dissolution of magnesium and deep pitting in some cases, resulted in large flaws that initiated fractures at lower loads. Lap-shear data can shed light on the changes in the joint due to corrosion – and will be used in conjunction with ASTM B117 salt- spray as a statistically significant and quantified material selection method. Coupons consisting of materials that resulted in more corrosion in ASTM B117 subsequently showed a greater degradation in lap-shear strength – suggesting lap-shear testing following corrosion exposure can be used as a practical measure of material compatibility for joints that may be at risk of corrosion.

Baseline lap-shear data

Lap-shear results for unexposed samples are given in Table 6: Lap-shear results for unexposed samples. The B-group samples were joined with only a SPR; the C-group samples were joined with a SPR and adhesive. The average B-group baseline load was

6,642.5 N with a standard deviation of 400 N. The average C-group baseline load was

10,885.1 N with a standard deviation of 691 N. These average lap-shear strength results

correspond with fatigue-life data compiled in USAMP Phase I, in that the adhesive creates a stronger joint, but results in more scatter in the data. The increased scatter associated with the adhesively bonded & riveted joint is evidenced by the higher C-group

standard deviation.

Failure in Rivet-only samples

Samples consistently failed by fracture in the AM60B, with the fracture initiating at the rivet-hole. AM60B being the weaker and less ductile of the three metals, it makes sense that this would be the common failure mechanism.

Failure in Rivet & Adhesive Samples

Failure in the adhesive samples was more complicated than in the rivet-only samples. Prior to B117 exposure, coupons invariably display a “primary” failure and a

“secondary” failure (Figure 24). The adhesive sustained a higher load than the rivet, initially. There was then a decohesion event, releasing 3-5,000 N, and the load versus displacement curve began to follow a curve similar to that of rivet-only samples, eventually “failing again” by the typical fracture in the AM60B. The decohesion event occurred at about 0.5 mm total displacement, which is about a mm less than the riveted joint can endure. As one might predict, prior to corrosion, the “secondary” failure event occurred at about the same strength and displacement as the failure in rivet-only samples.

After corrosion, however, the joints sometimes failed by fracture above the rivet. This will be further discussed later in this section.

ASTM B117 Salt-spray exposure

The coupons designated A4, A7, B22, B24, C23, and C27 were sectioned prior to exposure (as described in Experimental Procedures). Sectioned coupons were prepared for salt-spray exposure by masking cut-edges with non-conductive silicone. It was observed prior to exposure that edge corrosion did not occur on any of the samples. The exposure schedule conducted at OSU, as described in Table 5, only involved 1 and 2 week exposures.

Three-week exposures were carried out by PPG Industries (Pittsburgh, PA). Thus coupons A15, A32, A35, A36, C22, and C30 were sectioned following exposure. This is one difference between the sample sets that is an unfortunate bi-product of having very limited coupon availability. It is assumed that this had a negligible effect on the corrosion damage.

168 Hour Exposure Lap-shear Test Data

Twelve total coupons (A4, A7, B22, B24, C23, and C30 samples 1 and 2) were exposed for 168 hours in ASTM B117 at OSU. Visually, A4 and A7 displayed substantially less corrosion than B22, B24, C23 and C30. Table 6 shows the lap-shear data.

The A-group samples (no adhesive, Sn/Zn-coated rivet) had an average lap-shear strength of 6,220 N. This is 94% of the baseline average of 6,643 N. B-group samples (no adhesive, IVD-Al-coated rivet) had an average lap-shear strength of 6,692.5 N, which is

100% of the original strength. The C-group samples (adhesive, IVD-Al-coated rivet) had an average lap-shear strength of 12,652.5 N. This is 116% of the baseline average of

10,885 N. Having only 4 samples in each population, the Welch-Aspin adaptation of the

Student’s t-test was used to test the statistical significance of the results at a 95%

confidence level.

336 Hour Exposure Lap-Shear Test Data

Twelve total coupons (A4, A7, B22, B24, C23, and C30 samples 3 and 4) were

exposed for 336 hours in ASTM B117 at OSU. Again, A4 and A7 displayed substantially

less corrosion than B22, B24, C23 and C30. Table 6 shows the lap-shear data.

After 336 hours ASTM B117 exposure, A-group samples had an average lap- shear strength of 6,327.5 N. This is 95% of the baseline average of 6,643 N. B-group samples had an average lap-shear strength of 6,145.8 N, which is 92.5% of the original strength. The C-group samples had an average lap-shear strength of 9,890 N. This is

90.9% of the baseline average of 10,885 N. Again, having only 4 samples in each population, the Welch-Aspin adaptation of the Student’s t-test was used to test the statistical significance of the results at a 95% confidence level.

500 Hour Exposure Lap-shear Test Data

The 500 hour exposures were conducted in the salt-spray chamber at PPG. As described previously, the coupons were then sent to OSU, where they were sectioned and lap-shear tested. The results are in Table 9:

Summary of t-statistics

By the Welch-Aspin t-test with a 95% confidence interval, all A-group samples,

168 hour B-group samples, and 500 hour C-group samples were “passes,” meaning B117 exposure could not be determined to significantly affect the average lap-shear strength of the coupon sets. ASTM B117 exposure did significantly change the lap-shear strength in

336 hour B-group, and 168 and 336 hour C-group coupon sets. Lap-shear strength

decreased in the 336 hour B and C-group sample sets, but increased in the 168 hour C-

group sample set. It remains unknown why the lap-shear strength would increase after a week of ASTM B117 exposure in samples with adhesive. A-group coupons all had

Sn/Zn-coated rivets, while the B and C-group coupon sets had IVD-Al-coated rivets.

These results are summarized by Table 10.

If the difference of the population means is significant at the 0.05 level, this test

conveys with 95% confidence that the strength of the joint changed after ASTM B117

salt-spray exposure, suggesting corrosion activity affected the lap-shear strength of the

joint.

Curiously, lap-shear strength increased in C-group samples after 168 hours B117

exposure. The reason for this is unclear. In a study on the mechanical strength and

corrosion behavior of interstitial-free steel self-pierce riveted to an aluminum alloy,

Ioannou10 introduced an explanation for a similar phenomenon in lap-shear results. It is

suggested that minor buildup of corrosion product in the lap causes an increase in the

frictional force resisting rivet pull-out, however the load-transfer mechanism in

adhesively bonded assemblies is different, and thus this explanation is probably invalid.

The reason for the strength increase at this time is unknown.

Lap-shear strength in B and C-group samples after 336 hours of ASTM B117

exposure decreased significantly. By inspection of the joint and corrosion

phenomenology, joint strength loss probably occurred due to galvanic corrosion between

the aluminum-coated rivet and the adjacent magnesium, resulting in reduction in the

magnesium thickness.

ANOVA

The significance of the variance in each population is first tested by the Chi-

Squared test. Once independence has been decided, variance within the relevant

populations can then be analyzed.

Error Variance Determination

The error variance was established for A and B-group coupons by lap-shear

testing a 12-coupon population of unexposed rivet-only assemblies. It is relevant to

compare unexposed coupon strength of coupons with either Sn/Zn or IVD-Al coatings

prior to ASTM B117 exposure. The lap-shear strength is assumed to be independent of

the rivet coating until exposure to B117. However, following exposure, the A and B

groups will be subdivided and treated as independent populations due to major visual

differences in the corrosion phenomenology. The error variance for C-group coupons will

be established by testing a 12-coupon population of unexposed rivet & adhesive

assemblies.

The variance (Table 11) for each population set at each exposure time will be rigorously compared to this error variance with the Chi-square test. If at the 0.05 level, the variances are determined to be significantly different, the difference will be assumed to be due to treatment variance. This will then be used to establish the F-ratio for each population.

Treatment Variance Determination

To determine significance of additional variance in data, a two-sample Chi-

squared test for variance was conducted with the following hypothesis:

H0: µ1=µ2 (the variance of group A and baseline are the same)

Ha: µ1≠µ2 (the variance of group A and baseline are different)

Two population subsets generated significant treatment variance. The 168 hr B-

group subset F-ratio was 1.18. The 500 hr C-group subset F-ratio was 4.97.

Discussion

The statistical approach presented above will now be combined to rank the performance of the coupon populations. A simple scoring method will be used to rank the populations. If a population subset’s mean was confirmed by the t-test to be significantly different, it is assumed that the lap-shear strength was affected by B117 exposure, and that particular population subset will get one point. If the population subset’s F-ratio was calculated, the F-ratios for that population will be accumulated and presented as evidence that the variance in the results of the population was affected by exposure to B117.

Table 12 gives the t-stat “score” established by the previously described method.

If the sample received a t-stat “score,” this indicates that B117 exposure significantly altered the mean lap-shear strength. An F-ratio “score” indicates that B117 exposure significantly altered the variation in the lap-shear strengths. A lower score indicates better performance. A score of “0” indicates that the exposure did not result in a significant change in lap-shear strength of the joint.

To summarize, the population set with the lower score (A) is the better performer, and the population set with the higher score, (C), is the worst performer. This approach

78 clearly illustrates that the Sn/Zn rivet coating is the superior coating for maintaining joint integrity after ASTM B117 exposure. Both B and C populations, which had the IVD-Al rivet coating, also visually displayed substantially more corrosion after exposure to

ASTM B117. Examples are shown in Figure 26, Figure 27, and Figure 28 of corrosion surrounding rivets after 500 hours ASTM B117 exposure time.

Despite extreme dissolution of the AM60B plate surrounding the SPR, all joint populations maintained an average of at least 90% of the original lap-shear strength. The reason for this is related to the geometry of the corrosion phenomenology. Figure 25 is a schematic of the possible phenomenological basis for the geometry of magnesium dissolution.

When Mg dissolves, the local pH increases. Once it reaches about pH 10.5, which happens within minutes, the magnesium passivates. This results in a pH gradient that pushes the dissolving magnesium farther from the cathode (rivet), despite the tendency for anodic and cathodic reactions to occur as close together as possible (in order for electrons to achieve the shortest path). Since the Mg is dissolving slightly away from the rivet, the magnesium cross-section in intimate contact with the rivet is maintained. Any fracture that occurs in the magnesium will tend to initiate at the rivet/Mg interface, as the rivet is a stress-concentrator. Since the magnesium cross-section is largely maintained, the lap-shear strength is in turn largely maintained.

Conclusions

Through a combination of lap-shear testing, ASTM B117 exposure, and a simple statistical analysis, a study was conducted to investigate the corrosion of self-pierce

79 riveted joints of AM60B magnesium and 6082 aluminum. Coupons were supplied with either Sn/Zn or IVD-Al-coated plain steel rivets. Coupons with Sn/Zn-coated rivets performed better overall in ASTM B117 salt spray exposure testing. Significant changes in lap-shear strength and significant increase in variation in the lap-shear results occurred in both IVD-Al coated rivet-only and IVD-Al coated rivet & adhesive coupon sets, suggesting the IVD-Al-rivet coating results in substantially more joint degradation when exposed to ASTM B117.

A rigorous statistical approach was used in order to establish significance in results despite having a very limited population set. This was used as evidence that B117 exposure caused treatment variance, which was attributed to corrosion degradation in lap- shear strength.

Severe dissolution of the magnesium surrounding the rivet occurred, yet the average lap-shear strength for all population sets was largely maintained. Regardless of the rivet coating and whether adhesive was used, the average lap-shear strength, even after 500 hours of ASTM B117 exposure, remained greater than 90% of baseline average lap shear strength.

References

1. http://www.dupont.com/industries/automotive/articles/lightweighting.html

2. http://usatoday30.usatoday.com/money/autos/story/2012-08-29/fuel- standards/57383050/1

3. http://www.usatoday.com/story/money/cars/2014/07/22/ford-f150-engines/12997489/

4. http://www.magnesium-elektron.com/data/downloads/ds475diecastingalloys.pdf

5. http://www.aalco.co.uk/datasheets/Aluminium-Alloy-6082-T4-Extrusions_147.ashx

6. http://www.henkelna.com/product-search-1554.htm?nodeid=8797997531137

7. McCune, R., Forsmark, J., Battocchi, D., & Upadhyay, V. (n.d.). Characterization of Coatings on Steel Self-Piercing Rivets for Use with Magnesium Alloys. Magnesium Technology 2015.

8. http://www.henrob.com/GB/what-is-hspr.php#

9. http://energy.gov/sites/prod/files/2014/03/f11/lm008_luo_2011_o.pdf

10. Ioannou, J. (n.d.). Mechanical Behaviour and Corrosion of Interstitial-Free Steel - Aluminium Alloy Self-Piercing Riveted Joints. Retrieved from http://uhra.herts.ac.uk/handle/2299/4611?show=full

Tables and Figures

Table 2: Nominal composition (wt %) of magnesium and aluminum alloys used in this study.4,5

Al Mg Cu Fe Mn Ni Si Zn Cr Ti Other

5.5- <= <= 0.25- <= <= <= <= AM60B 94 6.5 0.01 0.005 0.60 0.002 0.10 0.22 0.20

Al Alloy 95.2- 0.6- <= <= 0.40- 0.70 <= <= <= <= 6082-T4 98.3 1.2 0.10 0.50 1.0 -1.3 0.20 0.25 0.10 0.15

Table 3: General mechanical properties of AM60B magnesium alloy and aluminum alloy 6082-T4.4,5

Shear Density UTS Elongation strength AM60B 1.8 g/cc 140 MPa 220 MPa 6-8%

Al Alloy 6082-T4 2.7 g/cc 170 MPa 260 MPa 19%

Table 4: Mg/Al 2-plate assemblies designated by letter and number code. Each assembly was sectioned into four individual lap-shear coupons.

Des. SPR ADHESIVE PRETREAT PRETREAT TOPCOAT BEFORE AFTER JOINING JOINING

A4 Zn/Sn no no Alodine 5200 ecoat

A7 Zn/Sn no no Alodine 5200 ecoat

A15 Zn/Sn no no Ecoat

A32 Zn/Sn no Ecoat

A35 Zn/Sn no Ecoat

A36 Zn/Sn no Ecoat

B22 IVD Al no no Alodine 5200 ecoat

B24 IVD Al no no Alodine 5200 ecoat

B155 IVD Al no no No none

B156 IVD Al no no No none

B157 IVD Al no no No none

B158 IVD Al no no No none

B159 IVD Al no no No none

B160 IVD Al no no No none

C2 IVD Al yes Alodine 5200 No none

C3 IVD Al yes Alodine 5200 No none

C4 IVD Al yes Alodine 5200 No none

C5 IVD Al yes Alodine 5200 No none

Continued

Table 4: Continued

C6 IVD Al yes Alodine 5200 No none

C7 IVD Al yes Alodine 5200 No none

C22 IVD Al yes Alodine 5200 Alodine 5200 ecoat

C23 IVD Al yes Alodine 5200 Alodine 5200 ecoat

C27 IVD Al yes Alodine 5200 Alodine 5200 ecoat

C30 IVD Al yes Alodine 5200 Alodine 5200 ecoat

Figure 20: Sheet-metal lay-ups were sectioned along the red dotted lines into four individual lap-shear specimens.

Figure 21: coated samples with masked edges.

Figure 22: uncoated samples

Table 5: B117 salt spray exposure schedule for sectioned, top-coated samples. 500 hr exposures were conducted at PPG’s laboratory.

Des. Sample 1 Sample Sample Sample 2 3 4

A4 168 hr 168 hr 336 hr 336 hr

A7 168 hr 168 hr 336 hr 336 hr

A15 500 hr 500 hr 500 hr 500 hr

A32 500 hr 500 hr 500 hr 500 hr

A35 500 hr 500 hr 500 hr 500 hr

A36 500 hr 500 hr 500 hr 500 hr

B22 168 hr 168 hr 336 hr 336 hr

B24 168 hr 168 hr 336 hr 336 hr

C22 500 hr 500 hr 500 hr 500 hr

C23 168 hr 168 hr 336 hr 336 hr

C27 168 hr 168 hr 336 hr 336 hr

C30 500 hr 500 hr 500 hr 500 hr

Figure 23: Schematic of shimmed lap-shear setup.

Table 6: Lap-shear results for unexposed samples

Sample 1 Sample 2 Failure Des. Load (N) Load (N)

B155 6416 6633 AM60B at rivet B156 6706 6199 AM60B at rivet B157 6716 6622 AM60B at rivet B158 6618 6720 AM60B at rivet B159 6453 7797 AM60B at rivet B160 6519 6311 AM60B at rivet C2 9993 10360 Decohesion C3 10182 11197 Decohesion C4 10285 10680 Decohesion C5 10764 10530 Decohesion C6 11968 11197 Decohesion C7 12141 11324 Decohesion

Figure 24: Lap-shear force vs. displacement curves of Al/Mg joint with rivet + adhesive

(blue) and rivet only (red).

Table 7: Ultimate load and failure mechanism of coupons exposed to B117 for 168 hours.

Sample Sample 1 Failure Des 2 Load Load (N) Mechanism (N)

A4 5850 6197 Fracture in Mg

A7 5980 6853 Fracture in Mg

B22 7711 7335 Rivet fail

B24 6298 5426 Fracture in Mg

C23 13117 12859 Decohesion

C30 12865 11769 Decohesion

Table 8: Ultimate load and failure mechanism of coupons exposed to B117 for 336 hours.

Sample Sample 3 Failure Des 4 Load Load (N) Mechanism (N)

A4 6398 6049 Fracture in Mg

A7 6437 6426 Fracture in Mg

B22 5981 6241 Fracture in Mg

B24 6122 6239 Fracture in Mg

C23 10289 9396 Decohesion

C30 10224 9650 Decohesion

Table 9: Ultimate load and failure mechanism for samples exposed to B117 for 500 hours in salt spray chambers at PPG’s facility.

Sample 1 Sample 2 Sample 3 Sample 4 Failure Des Load (N) Load (N) Load (N) Load (N) Mechanism

Fracture in A15 6986 6737 6674 Mg

Fracture in A32 6743 6054 6336 Mg

Fracture in A35 6609 6196 6531 Mg

Fracture in A36 6743 7213 7241 Mg

C22 8311 12102 11302 11029 Decohesion

C30 11813 11075 10158 8024 Decohesion

Table 10: Summary of t-statistics regarding significance of mean variance - pass/fail

(green/red)

Des. 168 336 500

B N/A

Table 11: Error variance as determined by lap-shear testing of unexposed coupon

populations

Group Des. N Mean Error Variance

A 12 6643 477000

C 12 10885 160000

Table 12: Scoring system for performance of coupon set in B117 exposure. A t-stat “score” indicates that B117 exposure significantly altered the mean lap-shear strength. An F-ratio “score” indicates that B117 exposure significantly altered the variation in the lap-shear strengths. A lower score indicates better performance.

Des. t-stat F-ratio Rank

A 0 0 0

B 1 1.18 2.18

C 0 4.97 4.97

Figure 26: Sn/Zn-coated rivet samples after 500 hours ASTM B117 exposure. Visually all Sn/Zn-coated rivets displayed the least corrosion of all three coupon groups.

Figure 27: IVD-Al coated rivets with no adhesive after 500 hours ASTM B117 exposure. Note the large amount of Mg dissolution surrounding the rivet.

Figure 28: IVD-Al coated rivets with adhesive after 500 hours ASTM B117 exposure. Visually the corrosion was very similar to the IVD-Al coated rivets with no adhesive.

Summary and Suggestions for Future Work Advanced materials are essential to achieve United States Department of Energy

Corporate Average Fuel Economy Standards. Much work is required regarding research

and development of corrosion prevention and protection techniques of such materials in

order for their implementation to be cost-effective and safe for consumers.

This work was put forth in an effort to add to the existing body of knowledge regarding the corrosion of magnesium and aluminum alloys, high strength steels, and the joining techniques that will be implemented to enable the incorporation of such materials in automobiles.

To effectively use magnesium alloys in automobiles, work must be done to evaluate the corrosion phenomenology of magnesium. Galvanic corrosion of magnesium and aluminum alloys, and high-strength steels will be an ongoing concern, and more work must be done in this area, to evaluate the long-term characteristics of corrosion in complex systems. Not only will automobiles implementing these kinds of materials have to meet all the safety, cost, manufacturability, and consumer-oriented demand requirements of automobiles on the market today, but automakers will also require the ability to fore-see the demands of “tomorrow” that arise from using new advanced materials.

Environmental hydrogen degradation of high-strength steels in automobiles is a concern, and methods must be developed in order to investigate the complex material microstructures that will arise as a result of forming technologies. Self-pierce riveting, for 100 example, relies on cold-working of hardened steels that will have an effect on the hydrogen-trapping capacity of such steels. Tests must be developed that are capable of reproducing the resulting microstructures in order for high-strength steels to be safely implemented in automobiles.

Finally, statistical analysis techniques using small population sizes in complex arrays of materials must be evaluated. This way the cost of research and development of new materials can be reduced as statistical analysis can be a powerful tool for material selection. If done intelligently and creatively, minimal amounts of material will be required in order to produce the maximum amount of information.

101

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