lost foam casting of periodic cellular materials

LOST FOAM CASTING OF PERIODIC CELLULAR MATERIALS

WITH ALUMINUM AND MAGNESIUM ALLOYS

by

Samson Shing Chung Ho

A thesis submitted in conformity with the requirements

for the degree of Master of Applied Science

Graduate Department of Materials Science and Engineering

University of Toronto

Copyright by Samson Shing Chung Ho (2009)

ii

Lost Foam Casting of Periodic Cellular Materials with Aluminum and Magnesium Alloys

M.A.Sc. , Samson Shing Chung Ho (2009)

Department of Materials Science and Engineering

University of Toronto

This study investigates the possibility of fabricating periodic cellular materials (PCMs)

via the lost foam casting (LFC) process using aluminum alloy A356 and magnesium alloy AZ91.

This approach combines the structural efficiency of PCM architectures with the processing

advantages of near-net-shape LFC. An initial feasibility study fabricated corrugated A356 panels.

This was followed by a study of casting variables such as pattern design, vacuum assistance, and

alloying additions in order to improve the fillability of the small cross-section struts. Finally,

integrated pyramidal sandwich panels having different relative densities were subjected to

artificial aging treatments and subsequently tested in uniaxial compression. The A356 PCMs

experienced a continuous increase after yielding while the AZ91 PCMs exhibited strut fracture

after peak strength. The results showed the compressive yield strengths of this study are

comparable with those previously reported PCMs produced by different fabrication methods.

iii

Acknowledgements

I would like to express the deepest appreciation to Professor Hibbard and Professor Ravindran for

their encouraging guidance and endless support during the course of my work. Their mentoring has

provided insight into ongoing scientific research and a rapidly changing industry. Without their guidance

and persistent help, my dedication and best efforts for this work would not been possible.

I wish to thank my colleagues Francesco D’Elia, Abdallah Elsayed, Ken Lee, Sophie Lun Sin,

Lukas Bichler and Brandon Bouwhuis for their friendship. The technical support from Alan Machin,

Joseph Amankrah and Sal Boccia is greatly appreciated.

I would like to thank the Natural Science and Engineering Research Council of Canada and the

University of Toronto for their financial support.

Finally I am grateful to my parents for their never-ending care, my brother for his friendship and

my girlfriend Joey for her endless support, throughout the years.

iv

Table of Contents

Abstract

Acknowledgements

Published Work

List of Tables

List of Figures

List of Symbols

List of Abbreviations

ii

iii

vi

vii

ix

xiii

xiv

Chapter 1: Introduction 1

Chapter 2: Literature Review 6

2.1. Periodic Cellular Materials (PCM) Topologies 6

2.2. Fabrication Methods in Wrought and Cast Periodic Cellular Materials 8

2.2.1. Investment Cast Periodic Cellular Materials 9

2.3. Types of Commercial Casting Methods 12

2.4. Lost Foam Casting (LFC) Process 13

2.4.1. Lost Foam Casting Advantages 15

2.4.2. Defects in Lost Foam Casting 16

2.4.3. Challenges in Magnesium Alloy Thin-Walled Lost Foam Casting 19

Chapter 3: Experimental 22

3.1. Foam Core Preparation 22

3.1.1. Corrugated Core Fabrication 22

3.1.2. Pyramidal Core Fabrication 23

3.1.3. Integrated Sandwich Fabrication 24

3.2. Lost Foam Casting 26

3.3. Post-Casting Characterization 28

3.4. Experimental Plan 30

Chapter 4: Results and Discussion 31

4.1. Initial Feasibility Study 31

4.2. Effect of Casting Conditions on LFC PCMs 34

4.2.1. Sprue and Runner Designs 35

4.2.2. Vacuum Assisted LFC 38

4.2.3. Alloying Additions 43

v

4.2.4. Preliminary Compressions Test Results 51

4.3. Integrated Pyramidal Sandwich 55

4.3.1. A356 PCMs Compression Test Results 56

4.3.2. AZ91 PCMs Compression Test Results 61

4.3.3. Lower and Upper Bound Model 66

Chapter 5: Conclusions 74

References 75

Appendix 80

A.1. Schematic Diagram 81

A.2. Phase Diagrams 82

A.3. Grain Fineness Number 83

A.4. Integrated Pyramidal Sandwich Compression Test Results 84

A.5. Previous Studied PCMs and Metal Foams 86

vi

Published Work

This thesis is based on the following publications:

a) Ho, S., Bichler, L., Hibbard, G.D. and Ravindran, C. 2008. Synthesis-Structure

Relationships in Cast Magnesium Periodic Cellular Materials, Magnesium Technology, TMS

2008: New Orleans, L.A.

b) Ho, S., Ravindran, C. and Hibbard, G.D. 2008. Lost Foam Casting of Magnesium Periodic

Cellular Materials, Processing and Fabrication of Advanced Materials XVII: New Delhi,

India, (Paper AC-136), pp.309-317.

c) Ho, S., Ravindran, C. and Hibbard, G.D. 2009. Production of Magnesium Thin-Wall

Cellular Castings Through Lost Foam Casting, AFS Transactions, Vol. 117, (Paper 09-051),

pp. 857-865.

d) Ho, S., Ravindran, C. and Hibbard, G.D. 2009. Fabrication of Periodic Cellular Materials

with Lost Foam Casting in Magnesium Alloy AZ91, Scripta Materialia (in preparation).

e) Ho, S., Ravindran, C. and Hibbard, G.D. 2009. Fabrication of Integrated Pyramidal

Sandwich Panels with Lost Foam Casting in Aluminum Alloy A356, Materials Science and

Engineering A (in preparation).

vii

List of Tables

Table 2-1. Fabrication details and compressive properties of various cast PCMs. 11

Table 2-2. Thermal degradation of EPS and PMMA foam pattern [Shivkumar,

1994].

19

Table 3-1. Dimensions of the corrugated sandwich core. 23

Table 3-2. Dimensions of the pyramidal lattice core. 24

Table 3-3. Dimensions of the integrated pyramidal sandwich. 25

Table 3-4. A356.2 Aluminum alloy chemical composition in wt% [House of

Metals, 2008].

27

Table 3-5. AZ91D Magnesium alloy chemical composition in wt%. 27

Table 3-6. The heat treatment schedule for A356 [Davis ed., 1993] and AZ91

[Avedesian and Baker ed., 1999].

29

Table 3-7. Etchants used in microscopic examination of A356 [Keller, 1948] and

AZ91 [Maltais et al., 2004] alloys.

29

Table 4-1. The casting parameters for the corrugated core assembly. 31

Table 4-2. The dimensions of the corrugated core assembly. 31

Table 4-3. Microhardness measurements as a function of distance from the

downsprue.

34

Table 4-4. The casting parameters for the vertical sprue lattice core assembly. 35

Table 4-5. The dimensions of the vertical sprue lattice core assembly. 36

Table 4-6. The dimensions of the vertical sprue sandwich assembly. 37

Table 4-7. The casting parameters for the vertical sprue sandwich assembly. 37

Table 4-8. The casting parameters for the horizontal sprue pattern assembly. 38

Table 4-9. The dimensions of the horizontal sprue pattern assembly. 38

Table 4-10. The casting properties for the narrowed vertical sprue pattern assembly. 39

Table 4-11. The dimensions of the narrowed vertical sprue pattern assembly. 39

Table 4-12. Summary of casting conditions: the vacuum level, pouring temperature,

and number of coatings for the eight different casting conditions.

40

Table 4-13. Summary of porosity and microhardness characterization. 43

Table 4-14. The casting properties for the narrowed vertical sprue pattern assembly. 44

Table 4-15. The dimensions of the narrowed vertical sprue pattern assembly. 44

Table 4-16. Measured chemical composition by spark emission spectroscopy for

each casting condition in wt%.

45

Table 4-17. Summary of casting parameters for integrated sandwich. 55

Table 4-18. The dimensions of integrated sandwich pattern assembly. 56

viii

Table A-1. AFS sieve number and the multiplying factor Mi [Rao, P.N., 1999]. 83

Table A-2. Yield strength and peak strength of A356 and AZ91 PCMs in compressions. 84

Table A-3. Experimental measured yield strength (σ yield PCM) and peak strength

(σ peak PCM) of A356 and AZ91 PCMs in compressions.

85

Table A-4. Previous studies of cast PCMs, summarizing the alloy, architecture, bulk

metal density (ρs), relative density ( ρ~ ), PCM’s density (σ peak PCM),

PCM’s specific peak strength (σ peak PCM / ρ PCM), PCM’s yield strength

(σ yield PCM) and PCM’s specific yield strength (σ yield PCM / ρ PCM).

86

Table A-5. Previous studies of aluminum PCMs, summarizing the alloy, architecture,

bulk metal density (ρs), relative density ( ρ~ ), PCM’s density (σ peak PCM),

PCM’s specific peak strength (σ peak PCM / ρ PCM), PCM’s yield strength

(σ yield PCM) and PCM’s specific yield strength (σ yield PCM / ρ PCM).

87

Table A-6. Previous studies of magnesium metal foams, summarizing the alloy,

architecture, bulk metal density (ρs), relative density ( ρ~ ), PCM’s

density (σ peak PCM), PCM’s specific peak strength (σ peak PCM / ρ PCM),

PCM’s yield strength (σ yield PCM) and PCM’s specific yield strength

(σ yield PCM / ρ PCM).

88

ix

List of Figures

Figure 1-1. The sandwich panel consists of a core between a pair of face sheets

where hfs = thicknesses of the face sheets, hc = thicknesses of the

core, L = the sandwich length, b = the sandwich width and htotal =

the sandwich height [after Ashby, 2005].

1

Figure 1-2. Pin-jointed frame with (A) bending-dominated mechanism and (B)

stretch-dominated mechanism [Ashby, 2005].

2

Figure 1-3. Material-property chart of Young’s modulus and density [Ashby,

2005].

3

Figure 2-1. Honeycomb core structures: (A) hexagonal, (B) square and (C)

triangular shaped [Wadley, 2006].

6

Figure 2-2. Corrugation structures: (A) triangular, (B) diamond and (C) navtruss

[Wadley, 2006].

7

Figure 2-3. Lattice architectures: (A) tetrahedral, (B) pyramidal and (C) kagome

[Wadley, 2006].

7

Figure 2-4. Normalized strength with respect to relative density for different

cellular architectures [Wadley, 2006].

8

Figure 2-5. Investment cast lattice sandwich [Chiras et al., 2002]. 10

Figure 2-6. The series of major steps in Lost Foam Casting process [dos Santos et

al., 2007].

15

Figure 2-7. Average metal velocity window with different defects [Hess, 2004]. 17

Figure 2-8. Lost foam casting defects: (A) cold laps and (B) misruns [Foseco,

1991].

17

Figure 3-1. Schematic diagram of corrugated sandwich core. 22

Figure 3-2. Three-dimension schematic model of pyramidal lattice core. 23

Figure 3-3. Two-dimension schematic diagram of pyramidal lattice core. 24

Figure 3-4. 10 mm-thick foam panel is cut with (A) the hot-wire-cutting template

into (B) pyramidal lattice pattern.

24

Figure 3-5. Schematic diagram of integrated pyramidal sandwich showing the

integration of node thickness to face sheets.

25

Figure 3-6. 20 mm-thick foam block is carved with (A) hot-wire-cutting template

into (B) integrated pyramidal sandwich.

26

Figure 3-7. Schematic representation of experimental plan. 30

Figure 4-1. As-cast corrugated core with pouring temperatures of (A) 710 °C and

(B) 730 °C.

32

Figure 4-2. Cross-sectional node geometry from the as-cast corrugation core (a,b)

poured at 710 °C and (c,d) poured at 730 °C.

32

Figure 4-3. Typical porosity at the centre (A) and edge (B) of the corrugated core 33

x

at a pouring temperature of 730 °C.

Figure 4-4. Typical microstructure of as-cast AZ91D corrugated core and etched. 33

Figure 4-5. Pyramidal lattice core pattern (A) ready for casting and as-cast

samples (B).

35

Figure 4-6. Pyramidal sandwich pattern (A) ready for casting, A356 as-cast

sample (B) and AZ91 as-cast sample (C).

36

Figure 4-7. Pyramidal sandwich assembly with top feeding sprue (A) ready for

casting, AZ91 as-cast sample (B), altered pyramidal sandwich

assembly with tapered top feeding sprue (C) and AZ91 as-cast

sample (D).

38

Figure 4-8. Pyramidal sandwich assembly with narrow and lengthen sprue (A) and

AZ91 as-cast sample (B).

39

Figure 4-9. The particular casting condition for the eight samples (summarized in

Table 4-12) had a significant effect on the extent of fillability.

Complete pattern was obtained in samples S5, S6, and S8.

41

Figure 4-10. Figure 4-10. Percentage of truss core (left) and face sheets (right)

filled for the eight different casting conditions (Table 4-12).

41

Figure 4-11. Casting map summarizing the extent of face-sheet and truss core

filling.

42

Figure 4-12. SEM images showing exposed porosity on polished surfaces of sample

S1 (A), sample S3 (B), sample S6 (C), and sample S7 (D).

43

Figure 4-13. The assembled coated pattern (A) and sandwich panel in close up (B). 45

Figure 4-14. Schematic representation showing the position of thermocouples in the

pattern.

45

Figure 4-15. The final casting of magnesium PCMs with different alloying

additions.

46

Figure 4-16. Temperature (left) and time (right) profiles: T1-T3 (top of panel) and

B1-B3 (bottom of panel).

48

Figure 4-17. Solidification time (left) and cooling rate (right) profiles: T1-T3 (top

of panel) and B1-B3 (bottom of panel).

48

Figure 4-18. Optical micrographs (at same magnification, see scale bar from

sample s6) and thermal analysis of the castings.

49

Figure 4-19. The measured porosity levels with Archmedes principle. 50

Figure 4-20. Hardness of the cast samples. 51

Figure 4-21. Typical compression test coupon (2×3.5 pyramidal truss core cells). 52

Figure 4-22. Typical uniaxial stress strain curves showing the mechanical

behaviour with and without sample preparation edge effects.

52

Figure 4-23. Progressive node/face sheet fracture during compression testing

(image labels correspond to the testing points shown on Fig. 4-22).

53

xi

Figure 4-24. Comparisons between different AZ91 PCMs in compressions. 53

Figure 4-25. Integrated sandwich assembly (A) and AZ91 as-cast sample (C). 55

Figure 4-26. Stress-strain curves (left) and rate of change in stress with strain

(right) for typical A356 PCMs in compression with different

relative density ρ~ .

56

Figure 4-27. Struts morphology initially (A) and after compression loaded (B). 57

Figure 4-28. Stress-Strain curves for typical solutionized (left) and solutionized and

aged (right) A356 PCMs in compression with different relative

density ρ~ .

57

Figure 4-29. A356 PCM’s yield strength (0.2 % offset) with respect to relative

density: as-cast (left), solutionized (middle), and solutionized and

aged (right).

58

Figure 4-30. A356 PCMs’ strains at which they yield at 0.2 % offset with respect to

relative densities after different heat treatments.

58

Figure 4-31. A356 PCMs’ modulus with respect to strains after different heat

treatments.

59

Figure 4-32. A356 PCMs’ modulus with respect to strains and different relative

densities: as-cast (left), solutionized (middle), and solutionized and

aged (right).

59

Figure 4-33. A356 etched microstructures in as-cast (A), solutionized (B), and

solutionized and Aged (C).

60

Figure 4-34. A comparison of different aluminum alloy PCMs and the current

studied LFC PCMs.

60

Figure 4-35. Stress-strain curves (left) and change in strain (right) for typical AZ91

PCMs in compression with different relative densities ρ~ .

61

Figure 4-36. Struts morphology initially (A), after peak (B), expanded view of the

fractured struts after peak (C) and after stress plateau (D).

61

Figure 4-37. Strain-strain curves for typical solutionized (left) and solutionized and

aged (right) AZ91 PCMs in compression with different relative densities

ρ~ .

62

Figure 4-38. AZ91 PCM’s yield strength (0.2 % offset) with respect to relative

density ρ~ : as-cast (left), solutionized (middle), and solutionized and

aged (right).

62

Figure 4-39. AZ91 PCMs’ strains at which they yield at 0.2 % offset with respect to

relative densities after different heat treatments.

63

Figure 4-40. PCMs’ strains at peak stress with respect to relative densities after

different heat treatments.

63

Figure 4-41. PCM’s modulus with respect to strains after different heat treatments. 64

Figure 4-42. PCMs’ modulus with respect to strains and different relative densities: 64

xii

as-cast (left), solutionized (middle), and solutionized and aged

(right).

Figure 4-43. AZ91 etched microstructures in as-cast (A), solutionized (B), and

solutionized and Aged (C).

65

Figure 4-44. A comparison on yield strength with respect to density between

magnesium foams and present study of AZ91 LFC PCMs.

65

Figure 4-45. Lower bound model based on minimum strut cross-section thickness,

tc.

66

Figure 4-46. Stress-strain curve of A356 square compression blocks with different

heat treatment conditions.

67

Figure 4-47. A356 sample block after compression (A) and magnified surface

feature of the block (B).

67

Figure 4-48. A356 square compression block`s surface with crack propagations. 68

Figure 4-49. Stress-strain curve of AZ91 square compression blocks with different

heat treatment conditions.

69

Figure 4-50. AZ91 sample block’s fractured cross section after compression (A)

and its magnified surface (B) showing transgranular fractures with

parallel plateau and ledges.

69

Figure 4-51. AZ91 square compression block fracture surface after 45° fracture. 70

Figure 4-52. Compressive yield strength of A356 and AZ91 compression blocks

with different heat treatments.

71

Figure 4-53. The comparisons between summarized experimental A356 PCMs’

yield strengths with lower bound model: as-cast (left), solutionized

(middle), and solutionized and aged (right).

72

Figure 4-54. The comparisons between summarized experimental AZ91 PCMs’

yield strengths with lower bound model: as-cast (left), solutionized

(middle), and solutionized and aged (right).

72

Figure A-1. Schematic diagram of corrugation sandwich core with dimensions. 81

Figure A-2. Aluminum-silicon phase diagram (left) and magnesium-silicon phase

diagram (right) [Nayeb-Hashemi and Clark, 1988].

82

Figure A-3. Aluminum-magnesium phase diagram [Nayeb-Hashemi and Clark, 1988]. 82

xiii

List of Symbols

α Mg-Al Phase

Β Mg17Al12 Phase

δ Face Angle of Indenter

θ ° Strut Angle

ξ Metal Loss Coefficient

ρ~ Core Relative Density

ρl Density of Liquid Metal

ρp Density of EPS

ρ PCM Density of the PCM Core

ρs Density of the Bulk Material

ρw Density of Water at Room Temperature

σys Compressive Yield Strength of Bulk Material

σys PCM Compressive Yield Strength of PCM

a Section Thickness of Pattern

a/l Strut Aspect Ratio

A0 Minimum Cross-Section Area of the Strut

Am Cross-Section Area of Pattern

APCM Cross-Section Area of PCM Under Compression

As Cross-Section Area of Sprue

b Sandwich Width

Cl Specific Heat of Liquid Metal

d Mean Diagonal of Indention

dc Node Thickness

dT/dt First Derivative Cooling Curve

E~

Young’s Modulus of the PCM Core

sE Young’s Modulus of the Bulk Material

g Gravitation Constant

h Heat Transfer Coefficient

H Latent Heat of Liquid Metal

xiv

hc Thicknesses of the Core

HE Decomposition Energy of EPS

hfs Thicknesses of the Face Sheets

hs Sprue Height

htotal Sandwich Height

l Sandwich Length

lc Length of Characteristic Chill Zone

Lf Metal Flow Length

Lp Penetration Load of Indenter

Mm Mass of Sample

Mm+w Mass of Water with Sample

' Number of Struts with Each PCM Under Compression

Pp Back Pressure Due to Gaseous EPS

Rc Radius of Curvature

tc Web Thickness

Tm Melting Temperature

To Mould Temperature

∆T Superheat Temperature

VE Metal Velocity

List of Abbreviations

ABS Acrylonitrile Butadiene Styrene

EPS Expanded Polystyrene

LFC Lost Foam Casting

pcf Pounds Cubic Feet

PCM Periodic Cellular Materials

PMMA Polymethyl Methacrylate

PS Polystyrene

Chapter 1: Introduction

1

1. Introduction

Hybrid materials are the combination of a monolithic material with another, or a monolithic

material with open space; they are designed to have their own set of properties that neither component can

offer alone [Wadley, 2006]. Hybrids include composites (fibrous or particulate), sandwich materials,

lattice structures and segmented structures [Ashby, 2005], and can reach new regions of material-property

space. The present study focuses on sandwich panels with lattice structure core.

A sandwich panel (as shown in Fig. 1-1) is an example of a hybrid with a combination of a

specially designed core that is layered between a pair of face sheets. Typically, the face sheet material is

required to be stiff and strong since it must carry most of the applied load. The core material can either be

a monolithic material or a hybrid material itself that is lightweight but strong and stiff enough to resist the

applied stresses [Ashby, 2005]. The purpose of the core is to minimize the mass near the sandwich

panel’s centroid, which contributes to higher specific bending stiffness and strength.

Figure 1-1. The sandwich panel consists of a core between a pair of face sheets where hfs = thicknesses of

the face sheets, hc = thicknesses of the core, L = the sandwich length, b = the sandwich width and htotal =

the sandwich height [after Ashby, 2005].

A variety of cellular architectures can be used as the sandwich core. The most commonly seen

architectures are metal foams and honeycomb structures. Recently, lattice truss or periodic cellular

materials (PCMs), have been studied as potential sandwich cores. These materials have internal struts

oriented in a regular three-dimensional architecture. In metal foams, the stiffness is relatively low because

their geometries permit bending of the cell edges which collapse when the externally applied load is


2

transmitted through the internal joints. This can be seen in the pin-jointed frame analogy in Fig. 1-2A.

Metal foams are therefore called bending-dominated structures [Ashby, 2005]. In contrast, lattice trusses

have higher structural efficiency than metal foams. In Fig. 1-2B, the lattice truss analog has a transverse

strut between the pin-jointed frames. As the load is applied, the force causes tension along the horizontal

strut before the cell collapses. These lattice trusses are often called stretch-dominated structures. For the

same material and density, lattice trusses will have a higher stiffness compared to the foam structure. This

performance difference is shown in Fig. 1-3, where the stretch-dominated structures retain higher stiffness

than bending-dominated structures with decreasing relative density.

Figure 1-2. Pin-jointed frame with (A) bending-dominated mechanism and (B) stretch-dominated

mechanism [Ashby, 2005].

With less than 20 % of the volume occupied by metal, the open space in PCMs can serve

additional functions such as: heat exchange by fluid flow within the core [Lu, 1999], thermal insulation

by isolating heat transfer between the face sheets [Wadley, 2006], acoustic damping by absorbing noise

and vibration within the core [Wadley, 2006], and impact resistance by absorbing energy during

architecture collapse [Zhang and Ashby, 1992].


3

Figure 1-3. Material-property chart of Young’s modulus and density [Ashby, 2005].

The potential applications for hybrids can be found in transportation industries and storage

systems. Karmann GmbH has developed the technology to produce sandwich body panels with cost-

effective performance that offer as much as an order of magnitude higher stiffness than similar steel body

parts with half the weight [Ashby et al., 2000]. The sandwich body panel consists of outer aluminum

skins bonded to titanium-hydride-expanded aluminum foam core. This structure can be used as firewalls,

roof panels and trunk panels in the car body shell. Another type of aluminum foam core (ALPORAS),

developed by Shinko Wire Company Ltd, provides sound absorbing capability that reduces highway

noise [Ashby et al., 2000]. Another type of open-celled aluminum foam (ERG DUOCEL) has been

developed for pressures tanks that requires constant temperature and pressure in moving systems [Ashby

et al., 2000]. The aluminum foam can prevent the tank from dimension change and baffle motion fluid

from affecting normal operations.


4

Several methods have been developed to fabricate PCMs. From the solid state, PCMs can be

fabricated by metal fabric layup [Sypeck and Wadley, 2001; Queheillalt and Wadley, 2005] and the

deformation forming of expanded or perforated metal sheets [Sypeck and Wadley, 2002; Bouwhuis and

Hibbard, 2006; Kooistra and Wadley, 2007]. PCMs can also be fabricated with template solidification

from the liquid state by investment casting [Deshpande and Fleck, 2001; Wallach and Gibson, 2001;

Chiras et al., 2002; Zhou et al., 2004; Li et al., 2008]. The newly developed lost foam casting (LFC) has

been utilized in manufacturing industries, and may provide an alternative PCM fabrication approach to

investment casting.

LFC was initially developed in 1962 [Goria et al., 1986] and has been commercially available

since the 1980s [Rodgers, 1988], providing a niche casting approach for products that are not ideal for

other casting methods. LFC is a near-net-shape method because it can produce components with complex

geometries and open cavities. In this process, molten metal is poured over the expanded polymeric foam

pattern, the heat from the molten metal degrades the foam into liquid and gas products by endothermic

reaction. The degradation product escapes through foam’s ceramic coating to the surrounding sand. The

molten metal fills the cavity and shape of the pattern as it solidifies; the metal velocity is an important

parameter since it determines the overall casting quality.

With increased interests in developing lightweight and high-performance materials for structural

applications, magnesium becomes a good candidate because of its lowest density among the structural

metals. While much research has been done to develop the LFC technology for aluminum alloys,

magnesium alloys have only been cast successfully in research environment [Bichler et al., 2003; Marlatt

et al., 2003]. The fillability of magnesium alloys with LFC becomes a challenge because of magnesium’s

lower density and heat content compared to aluminum, which results in lower metal velocity during

casting. Aluminum alloys offer high castability, while magnesium alloys may offer enhanced weight

specific material properties.


5

The main objective of this thesis was to use the lost foam casting method to produce magnesium

and aluminum alloy PCMs for the first time. Since PCMs have never been fabricated by LFC, the first

objective was to conduct an initial feasibility study. The second objective was to adjust the processing

parameters to enhance the PCM casting quality. The final objective was to determine the effect of relative

density and precipitation hardening on the mechanical properties of aluminum and magnesium alloy

PCMs.

This thesis is organized as follows: Chapter 2 presents a literature review of periodic cellular

materials and lost foam casting; Chapter 3 provides experimental procedures in foam core fabrication, lost

foam casting preparations and post-casting characterization; Chapter 4 presents the results and discussion

of the present study; Chapter 5 concludes the findings from this thesis.

Chapter 2: Literature Review

6

2. Literature Review

2.1 Periodic Cellular Architectures

There are three broad classes of architectures that have been studied in periodic cellular metals.

They are honeycomb, prismatic (corrugation) and lattice truss. The closed-cell honeycomb configurations

can be hexagonal, triangular, and square shaped (examples shown in Fig. 2-1). Their unit cells are

repeated in two dimensions that are oriented normal to the face sheets. If triangular configuration in a

honeycomb is rotated by 90° about the face sheet’s axis, a prismatic structure (corrugation) is then formed

as seen in Fig. 2-2. The cell structure is partially open in one direction. Diamond and navtruss

configurations are also possible by alternating the orientation and plate width. These architectures

contribute to anisotropic mechanical behavior; the stiffness of the panels is different when loading in

longitudinal or transversal directions. The open space volume in PCM can be further increased by

aligning slender beams instead of plates in different configurations as the core. Fig. 2-3 shows tetrahedral,

pyramidal and kagome lattice architectures. These structures have free flowing channels in two or three

directions. They become very mechanically efficient as the relative density decreases [Wadley, 2006].

Figure 2-1. Honeycomb core structures: (A) hexagonal, (B) square and (C) triangular shaped [Wadley,

2006].


7

Figure 2-2. Corrugation structures: (A) triangular, (B) diamond and (C) navtruss [Wadley, 2006].

Figure 2-3. Lattice architectures: (A) tetrahedral, (B) pyramidal and (C) kagome [Wadley, 2006].

The strength of the hybrid structure with respect to its bulk material (structural efficiency) is

largely determined by the particular cellular architecture, seen in Fig. 2-4. The structural efficiency

improves from foams, to corrugations, to honeycombs and finally to lattice trusses by moving towards the

top left corner of the figure. It is important to note that the structural efficiency of pyramidal solid trusses

matches that of square honeycomb as the relative density (structural density divided by material density)

reduces. The stiffness of the structure with respect to its density for stretching-dominated structure is

∝

s

PCM

s

PCM

E

E

ρρ

Eqn. 1-1,

and bending-dominated structure is

2

∝

s

PCM

s

PCM

E

E

ρρ

Eqn. 1-2,

where PCME = Young’s modulus of the structure, sE = Young’s modulus of the bulk material, PCMρ =

density of the structure and sρ = density of the bulk material [Ashby, 2005]. Overall, there is less

reduction in the cellular architectures’ stiffness with decreasing density for lattice trusses than for

conventional metal foams.


8

Figure 2-4. Normalized strength with respect to relative density for different cellular architectures

[Wadley, 2006].

2.2 Fabrication Methods in Wrought and Cast Periodic Cellular Materials

There are several methods that have been developed for creating cellular materials. These can be

classified into vapour, liquid or solid state approaches. The PCM’s architecture depends on the type of

fabrication methods and (precursor or sacrificial) materials used. The most common methods for foam

fabrications are foaming solidification with gas evolution or gas injections. The most common methods

for lattice truss periodic cellular materials are metal fabric layup and deformation forming of expanded or

perforated sheets. In the metal fabric layup method, individual metal mesh layers are stacked together and

joined to face sheets with transient liquid phase (brazing) method [Sypeck et al., 2002; Tian et al., 2007].

In the pressed expanded metal method, a rolled metal sheet is punched in different lattice architectures

and plastically deformed with a die into a three-dimensional architecture [Sypeck and Wadley, 2001;

Bouwhuis and Hibbard, 2006; Bouwhuis et al., 2008]. Most recently, electrodeposition on rapid-


9

prototyped polymer cores has been used to create nanocrystalline periodic cellular lattice materials

[Gordon et al., 2009]. Thus far, the only cast method developed for PCM fabrication is templated

solidification with investment casting; a polymer lattice is used as a sacrificial template and removed to

create the negative mould for subsequent investment casting [Chiras et al., 2002; Li et al., 2008].

2.2.1 Investment Cast Periodic Cellular Materials

In investment casting, the sacrificial pattern is created in wax or polymer and joined with a runner

and gating system. The whole pattern is then coated multiple times with ceramic casting slurry and heated

to remove the wax before filling with liquid metal. Cast periodic cellular materials have been fabricated

(see Fig. 2-5) using rapid-prototyped acrylonitrile butadiene styrene (ABS) or injection moulded

polystyrene (PS) as sacrificial lattice patterns in tetrahedral [Deshpande and Fleck, 2001; Chiras et al.,

2002], pyramidal [Wallach and Gibson, 2001], and kagome [Wang et al., 2003] architectures. Table 2-1

summarizes the cast PCMs studies conducted to date. Note that in certain cases, the physical properties of

the sacrificial materials and process parameters (injection moulded polystyrene and injection moulded

polymer) were not clearly explained in the literature as they were supplied by third parties such as Meka

Mouldings Ltd. and Jonathan Aerospace Materials Corporation (JAMCORP). The types of casting alloys

consist of copper, aluminum and titanium. High fluidity casting alloys such as Al-Si and Cu-Be are used

with higher than ideal relative density ( ρ~ < 5 %) because the casting process requires higher pressure

feeding of molten metal before solidification occurs. This becomes increasingly difficult as the strut

cross-section area is reduced and increases the possibility for casting defects.


10

Figure 2-5. Investment cast lattice sandwich [Chiras et al., 2002].


11

Table 2-1. Fabrication details and compressive properties of various cast PCMs.

# Note: the core architecture is not sandwiched between solid face sheets but perforated face sheets.

* Note: the core relative density stated is calculated from the overall relative density (includes face sheet) given from the source in parentheses.

‡ Note: the value was not given but is calculated based on the given strut dimensions from the source.

Ψ Note: the strut aspect ratio is the diameter to length ratio of the strut.

Authors Alloy Process Sacrificial

Material Architecture

Strut

Angle

Strut

Aspect

RatioΨ

Core

Relative

Density

Compressive

Strength

Yield Peak

θ° a/l ρ~ (%) σ PCM (MPa)

Deshpande and

Fleck, 2001

Al-7Si-0.3Mg

Investment

Cast

Injection

Moulded

Polystyrene

Tetrahedral#

55°‡ 0.071 8

‡

4.2 7.3

Cu-4Si-1.4Zn 4.2 10.5

Wallach and

Gibson, 2001 (443) Al-4Si-0.2Fe

Injection

Moulded

Polymer

Pyramidal# 45° 0.059 6.2

* (10) 13.5 18

Chiras et al., 2002

Cu-2Be

Rapid-

Prototyped

ABS

Tetrahedral

55°‡

0.041

2

4.0 4.8

Wang et al., 2003 Kagome 0.045 3.8 4.8

Zhou et al., 2004

(516.1) Al~3Mg~1Si-1Fe~1Mn Injection

Moulded

Polymer

Pyramidal#

55°‡ 0.053

‡ 6.5

* (13)

2.9 5.9

(518.0) Al~8Mg-1.8Fe-0.35Si 3.8 5.6

(A356.0) Al~7Si~0.4Mg 3.5 5.8

Li et al., 2008

Ti-6Al-4V

Vacuum

Investment

Cast with

Hot

Isostatic

Pressing

Wax 51°‡

0.047‡ 5.4

* (13)

19 22

0.052‡ 5.5

* (16)

29 31

Ti-6Al-2Sn-4Zr-2Mo 35 41


12

2.3. Types of Commercial Casting Methods

Some of the commonly used casting processes in the foundry industry include sand, permanent

mould and die casting. The preferred casting process for a component depends on various restrictions,

including a component’s geometry, production volume, design tolerance and alloy characteristics. These

constraints help to determine the most economical and feasible process for manufacturing.

Sand casting uses bonded sand to create drag and cope (top and bottom halves of the mould) with

removable core for pattern cavity [Monroe, 1992]. After liquid metal filling the space between sand

moulds and solidified, the as-cast product is removed by breaking the sand mould and machined to finish.

This is an economical process that has both low capital and maintenance cost for mould making [Monroe,

1992]. Since the initial cost is small for sand mould making, the quantity can be flexible as well.

However, the method can only provide rough surface finish and simple geometry within the mould, and

extensive machining is generally required.

Permanent mould casting uses a steel mould that is machined to the desired shape of the casting.

Casting defects can be minimized by heating the steel mould during casting. The as-cast product is

ejected from the mould with hydraulic power. If the design requires an open cavity within the product,

multiple moulds are necessary to create multiple casting components that are assembled later. This

process provides excellent surface finish, but yet it requires high initial cost in steel machined mould and

design limitations [Monroe, 1992].

Die casting uses a steel mould with liquid metal injection at high pressure. This process can

produce thin-walled castings having excellent surface finish that are difficult to fabricate with

conventional casting methods [Avedesian and Baker ed., 1999]. This process is suitable for high-level

production because of the reusability of the mould.


13

While there is no single process that is suitable for all requirements, each process has its unique

advantages and disadvantages. In addition to the conventional casting methods discussed above, lost foam

casting (LFC) has recently been developed for manufacturing aluminum components. LFC can be utilized

for niche applications, providing some features that other conventional casting methods cannot provide.

The benefits of this process will be further discussed below.

2.4. Lost Foam Casting Process

Before actual casting of liquid metal can occur in the lost foam process, a series of manufacturing

steps are required as shown in Fig. 2-6. Expanded polystyrene (EPS) is the most commonly used foam

pattern, it can be produced by bead pre-expansion into polystyrene (PS) beads that are ready for

moulding. PS precursors are formed from ethyl benzene through an aluminum catalyst with benzene and

ethylene obtained from crude oil and natural gas [Shivkumar, 1994]. Ethyl benzene is then converted to

styrene at high temperature with nitrogen gas and iron catalysts. It forms polystyrene when exposed to a

peroxide catalyst and polymerized in a water solution [Goria et al., 1986]. These unexpanded beads have

a density of 600 g/l (38 pounds per cubic feet (pcf)) and they are expanded 20~50 times with heat at

100 ºC until the desired density is reached [Kanicki, 1985].

Moulding process occurs after the pre-expanded beads have been stabilized. They are injected

into tooling machine and are ready to form the pattern sections. As they are blown into the tool cavity, air

escapes through a venting position to achieve a proper fill. Steam is then used to soften and expand the

beads again so that they begin to fuse [Goria et al., 1986]. After the steam cycle, the foam is cooled

rapidly and aged to avoid shrinkage. Prototype parts can be carved from foam blocks and assembled by

hand gluing.

Various pattern pieces are joined to create the finish model pattern which is then joined to

specially designed runners and sprue system using hot melt glue. The final model assembly is submersed

into refractory ceramic slurry, spun and air dried between 40~60 ºC, under low-humidity. Ceramic slurry


14

in a water-based solution consists of silica, alumina, zircon, chromite or complex aluminosilicates

[Monroe, 1992]. The purposes are: first to prevent molten metal penetrate through the pattern to the

surrounding sand; second, to allow the foam decomposition to escape; and third, to retain the pattern

geometry without distortion [Monroe, 1992]. Coating permeability is important in controlling the metal

velocity because the rate of filling depends on the backpressure from gaseous foam residue. Increasing the

rate of transport through the coating (air flow rate) can minimize the backpressure. The metal front has a

concave profile with low-permeability coating, transporting liquid products to the side for ease of removal

[Liu et al., 1997]. Coating permeability to gases is important for iron castings that have relatively high

casting temperatures. The metal velocity should be low enough to minimize foam residue entrapment by

controlling coating permeability and thickness [Green et al., 1998]. For aluminum castings, the coating

requires more liquid permeability for liquid foam residue [Monroe, 1992]. A uniform coating is necessary

as coating thickness can affect the casting quality. The primary control method is to determine the dry

coating weight of foam pattern.

The assembly is then embedded within the casting pit with refractory sand and compacted with

vibration to fill the assembly’s cavity. The lost foam process is finally ready for casting. Automatic

pouring is generally used to avoid pour-to-pour variations because if the pouring operation is interrupted,

metal flow will stop advancing and retreat backwards. The rate of pouring should be sufficiently slow to

allow the foam to decompose, but also fast enough to prevent the mould from collapsing [Bast et al.,

2004]. Vacuum can also be applied during casting if necessary; in this case the whole flask is sealed with

a plastic cover layer and connected to a vacuum pump. As the metal fills the mould, the foam collapses at

~100 ºC, and decomposes into liquid residue at 165 ºC [Gallois et al., 1987]. It then depolymerizes at

316 ºC and decomposes into gas at 576 ºC [Gallois et al., 1987]. The plastic residue that is trapped inside

the casting can cause porosity, cold laps or carbon defects [Gallois et al., 1987]. The casting solidifies as

the heat from the liquid metal escapes through the assembly’s coating into the sand. After solidification,

the pattern is taken out and shaken to remove any adhering sand. Finally, the pattern assembly is

sectioned and machined from the runner and sprue system before quality inspection can be done.


15

Figure 2-6. The series of major steps in LFC process [dos Santos et al., 2007].

2.4.1. Lost Foam Casting Advantages

In conventional sand casting, pattern design with internal cavities requires a core and extensive

passage that are removed after casting. The core can also shift during casting, which results in variation of

internal passages and wall thickness. In contrast, LFC only requires a small-hole passage to fill the

internal cavities with unbonded sand. This allows the designer to create complex geometries and, the time

and cost for core production and removals are not necessary. Using unbonded sand also reduces the

amount of cleaning time required [Rodgers, 1988]. This flexibility can reduce some of the design

limitations and cost between [Rodgers, 1988; Rodgers, 1985].

In addition to requiring a core for casting internal cavities, conventional moulding technology

often imparts limits on the mould design and accuracy of sand and permanent mould casting. For

example, the gap between upper and lower moulds during casting produces parting lines and flashes


16

which result in extra machining costs. This parting line from conventional casting methods can also limit

the orientation of the pattern in the moulds, number of castings in flask, placement of gates and risers, and

cause mould misalignments.

LFC is a near-net-shape process because it minimizes post-processing time. The foam patterns

can be glued together from many layers to create complex shapes. For instance, a product that requires ten

die casting parts which are assembled together with fasteners can be replaced with a single LFC casting.

An engine block is a good example that typically requires on the order of 100 different parts; these

different components can be integrated into one final assembled foam pattern for LFC [Donahue, 1990].

Large volume parts that require extensive coring and machining operations show the biggest benefits of

lost foam casting [Monroe, 1992]. Finally, LFC can be used to produce: sand cast components without the

need of bonded sand and core removal; permanent mould cast components without geometry limit and

extensive machining; or die casting components by integrating many parts into one single casting. These

advantages allow LFC to be applicable in producing steel and aluminum components such as brake rotors,

water pumps and engine components.

2.4.2. Defects in Lost Foam Casting

The advantages of LFC are only attainable, however, if the processing parameters are controlled

precisely. These processing parameters can affect the amount of casting defects, which can be minimized

by carefully adjusting the polymer degradation characteristics with polymer foam, foundry sand and

refractory coating properties. These properties are tailored according to the type of alloys used (e.g. steel,

aluminum and magnesium).

In lost foam casting, higher or lower metal velocity does not necessarily mean better quality

castings. There is an optimal range of average metal velocity (Fig. 2-7) for each alloy that can produce a

defect-free casting [Hess, 2004]. If the velocity is slower than optimum, metal would solidify and cause

cold laps or misruns (see Fig. 2-8) and heat from the metal would cause the pattern surface to collapse


17

before any metal filling occur. Cold laps form between two premature solidified flows of metal merging

together; misruns are more severe forms of cold laps where the discontinuity extends throughout the

casting [Beeley, 2001]. In contrast, backpressure and turbulences from the liquid metal would form

porosity and cold folds, and penetration defects when the velocity is too high.

Figure 2-7. Average metal velocity window with different defects [Hess, 2004].

Figure 2-8. Lost foam casting defects: (A) cold laps and (B) misruns [Foseco International Ltd., 1991].

The decomposition products from the foam pre-form can also be a source of casting defects. In

addition to the typically used polystyrene (PS), polymethyl methacrylate (PMMA) has been developed for

lost foam casting to reduce the carbon residue. PS contains benzene side chains that are highly stable and

difficult to thermally decompose [Brenner et al., 1990]. On the other hand, the PMMA monomer unit

contains three less carbon atoms than PS and two oxygen atoms to help carry away carbon residue as

carbon monoxide.


18

The decomposition mechanisms between PS and PMMA are different. PS decomposes by slow

random scission process; it degrades as a liquid and often remains trapped inside the casting [Brenner et

al., 1990]. Bromide-based additive can be used in PS to decompose PS molecular weight quickly and its

liquid residue is less viscous [Molibog, 2002; Hess et al., 2003]. This modified PS has a significant effect

on thin patterns where the surface area to perimeter is low because it is important to decompose quickly

when metal temperature is low and solidification time is short. PMMA decomposes by a rapid unzipping

process; its product residue is in the gaseous state which escapes quickly. Due to the difference in

decomposition mechanism between PS and PMMA, PMMA has only ~30 % viscous residue (rest in

gaseous product) at 750 °C, compared to ~60 % viscous residue in EPS [Shivkumar, 1995]. This viscous

residue can remain after the metal fills the mould, so it is very important in removing viscous residue

rapidly for EPS. In addition to decomposition mechanism, the energy required for heat of

depolymerization of PMMA is 14 kcal/mol compared to 17 kcal/mol in PS, this difference lowers the

thermal gradient between pouring temperature and metal feeding temperature within the casting [Brenner,

1989]. The differences in thermal degradation properties between EPS and PMMA are summarized in

Table 2-2. Even though the glass transition, collapsing and melting temperatures are higher with PMMA,

the starting, peak and end temperatures for volatilization are lower, resulting in a longer period for

volatiles to escape. The heat of degradation required for PMMA is also lower which requires less heat

energy from the molten metal compared to EPS.


19

Table 2-2. Thermal degradation of EPS and PMMA foam pattern [Shivkumar, 1994].

Thermal Degradation Properties EPS PMMA

Glass transition temperature (°C) 80-100 105

Collapsing temperature (°C) 110-120 140-200

Melting temperature for collapsed beads (°C) 160 260

Starting temperature of volatilization (°C) 275-300 250-260

Peak temperature of volatilization (°C) 400-420 370

End temperature of volatilization (°C) 460-500 420-430

Heat of polymerization (J/g) 648 578

Heat of degradation (J/g) 912 842

Gas yield at 750 °C (cm3 (STP)/g) 230 273

Gas yield at 1300 °C (cm3 (STP)/g) 760 804

Density and the level of cell bonding (fusion) are the main control properties of the foam. The

density is largely influenced by the size of air cells [Monroe, 1992]. Low-density foams are more

desirable for the casting process because it minimizes foam residue trapped within casting which can lead

to porosity and carbon defects. But at low density, the stiffness of foam is low, allowing higher chance for

distortion.

2.4.3. Challenges in Magnesium Alloy Thin-Walled LFC

In order to minimize casting defects, it is important to understand and control the metal velocity.

Foam properties, pattern dimensions and alloy properties influence the metal velocity as shown in

equation 2-1 [Shivkumar et al., 1987; Pan and Liao, 2000]:

Eqn. 2-1

where VE = metal velocity (cm/s), As = cross-section area of sprue (cm2), Am = cross-section area of

pattern (cm2), ξ = metal loss coefficient, g = gravitation constant (cm/s

2), hs = sprue height (cm), Pp =

back pressure due to gaseous EPS (g/cms2), ρl = density of liquid metal (g/cm

3). From a pattern design

perspective, the metal velocity would increase by enlarging the cross-section area ratio of the sprue

relative to pattern, this same would occur by increasing the sprue height. The metal velocity would

−

+=

g

Phg

A

AV

l

p

s

m

sE ρξ

21

1


20

decrease with back pressure due to higher gaseous foam products and the use of lightweight alloy. This

lower metal velocity results in reduced flow length as shown in equation 2-2 [Shivkumar, 1987; Pan and

Liao, 2000]:

Eqn. 2-2

where Lf = metal flow length (cm), ρL = density of liquid metal (g/cm3), a = section thickness (cm), VE =

metal velocity (cm/s), h = heat transfer coefficient (W/cm2°C), Tm = melting temperature (°C), To = mould

temperature (°C), H = latent heat (J/g), Cl = specific heat of liquid metal (J/g°C), ∆T = superheat

temperature (°C), ρp = density of EPS (g/cm3), lc = length of characteristic chill zone (cm) and HE =

decomposition energy of EPS (J/g). The flow length is affected by the physical and thermal properties of

the liquid metal, casting parameters and foam properties. The flow length would decreases: with liquid

metal that has low density, and low latent heat and low specific heat; with foam that has high density and

requires high decomposition energy.

When low-melting temperature magnesium or aluminum alloys are used in lost foam casting, the

foam pattern does not decompose completely into gas residue [Tseng et al., 1992]. The foam decomposes

partly into liquid as the pattern is filled with metal; some of the liquid residue can permeate through the

coating but the rest is trapped during solidification. The mould filling rate is a combination of foam

decomposition rate and foam product removal rate. In aluminum lost foam casting, the higher pouring

temperature does not necessarily increase the fill rate because the metal velocity is reduced by the

increasing backpressure at the thin sections [Lawerence et al., 1998]. The major challenge in casting these

structures is the ability to fill the whole pattern before the fraction of solids increase to the level where the

metal flow stops and solidifies.

Using flow length as an indicator, aluminum alloy A356 was found to be more fluid than

magnesium alloy AZ91 for the same superheat in gravity casting [Sadayappan et al., 2006]. The poor

fluidity of magnesium alloy is attributed to its low density, low heat capacity and low latent heat [Shin et

( )

−∆+

−= E

lc

pf

m

ELf H

l

LTc

H

TTh

aVL

ρ

ρρ

ο1

22~


21

al., 2005]. The results from Sadayappan et al. shows AZ91 with shorter flow length compared to A356 in

either gravity or vacuum lost foam casting. Under gravity casting with the thickness of the pattern

reduced to half of the original, the flow length of AZ91 decreased to one third compared to only one half

in A356. The difference between AZ91 and A356 becomes more significant when increasing the cross-

section-area ratio of gating to sprue system. In addition, AZ91 could only achieve similar flow length to

A356 at the same vacuum level when the thickness of the pattern was doubled.

Chapter 3: Experimental

22

3. Experimental

3.1. Foam Core Preparation

3.1.1. Corrugated Core Fabrication

The starting pre-forms were 10 mm-thick expanded polystyrene (EPS) boards with 0.026 g/cm3

(1.6 pcf) density. A milling machine was used to fabricate the initial corrugated panel. EPS was machined

with a 60° chamfer tool bit, the rotational speed and feeding rate were adjusted to 2700 rpm and

50 mm/min respectively to prevent the foam beads from rupturing. A series of parallel cuts were made on

both sides of the EPS boards to obtain a triangular corrugated structure. Fig. 3-1 shows a schematic

representation of the corrugated structure; the height of the core, hc, was constrained by the starting

thickness of the as-received foam; the thickness of the node, dc, was set equal to the web thickness, tc; and

the panel angle, θ, was chosen to be 60º. Since the average EPS bead size was ~1 mm, a minimum web

thickness of 3 mm was selected to ensure approximately three bead diameters through each segment of

the EPS cross-section area for structural integrity. The core dimensions are summarized in Table 3-1. The

relative density, ρ~ , was determined to be 51 % using the following equation (Appendix A-1):

( )

+

−

−

−=

θθ

θρ

sintan

2

tan1~

2

cccc

cc

tdhh

dh

Eqn. 3-1

Fig. 3-1, Schematic diagram of corrugated sandwich core.


23

Table 3-1. Dimensions of the corrugated sandwich core.

Dimensions

Height of the core, hc 10.0 mm

Thickness of the node, dc 3.0 mm

Web thickness, tc 3.0 mm

Corrugation angle, θ 60 º

Relative Density, ρ~ 51 %

3.1.2. Pyramidal Core Fabrication

A pyramidal lattice design was chosen for the second round of pattern design in order to reduce

the relative density of the as-cast core, see Fig. 3-2. 10 mm-thick EPS boards having density of

0.021 g/cm3 (1.3 pcf) and 10 mm-thick expanded copolymer Probeads 70 boards (30 % PS and 70 %

PMMA) having 0.024 g/cm3 (1.5 pcf) density were used as the starting pre-forms. All the dimensions

were the same as that used in the corrugated patterns with the addition of radius of curvature at the node

(Fig. 3-3) as summarized in Table 3-2. Instead of the milling technique previously used, the hot-wire

cutting method was utilized with a guiding template (Fig. 3-4A). The current in the nickel chromium

resistant wire was adjusted to ~5 A with a (10 A-capacity) power supply in order to provide enough heat

to locally melt but not burn the foam. A series of parallel cuts were made on both sides to obtain a

triangular corrugated structure. The foam panel was then rotated 90º for another series of parallel cuts to

obtain the pyramidal lattice core (Fig. 3-4B). The relative density decreased from 51 % to 22 % by adding

the second dimension of corrugation to form the pyramidal lattice.

Figure 3-2. Three-dimensional schematic model of pyramidal lattice core.


24

Table 3-2. Dimensions of the pyramidal lattice core.

Dimensions


Thickness of the node, dc 3.0 mm

Thickness of the strut, tc 3.0 mm

Strut angle, θ 60 º

Radius of Curvature, Rc 3.2 mm

Relative Density, ρ~ 22 %

Figure 3-3. Two-dimensional schematic diagram of pyramidal lattice core.

Figure 3-4. 10 mm-thick foam panel is cut with (A) the hot-wire-cutting template into (B) pyramidal

lattice pattern.

3.1.3. Integrated Sandwich Fabrication

The pyramidal lattice approach had fillability issues because of the hot melt glue used between

the struts and the face sheets. The integrated pyramidal approach eliminated this problem by fabricating

both the pyramidal lattice and the face sheets from a single block of polymer foam. The dimensions

stayed the same as the previous pyramidal lattice design summarized in Table 3-3. The thickness of the

node, dc was eliminated by integrating the lattice core with face sheet (Fig. 3-5). Four different strut


25

thicknesses were studied; tc had values of 1.5 mm, 2.0 mm, 2.5 mm and 3.0 mm, which led to relative

densities of approximately 11 %, 12 %, 14 % and 16 %.

Figure 3-5. Schematic diagram of integrated pyramidal sandwich showing the integration of node

thickness to face sheets.

Table 3-3. Dimensions of the integrated pyramidal sandwich.

Dimensions


Height of the face sheets, hfs 5.0 mm

Thickness of the node, dc N/A

Thickness of the strut, tc 1.5, 2.0, 2.5, 3.0 mm

Strut angle, θ 60 º

Radius of Curvature, Rc 3.2 mm

Relative Density, ρ~ 11, 12, 14, 16 %

The hot-wire-cutting method was used to section the 33 mm-thick expanded copolymer Probeads

70 foam board, having 0.024 g/cm3 (1.5 pcf) density, to a 20 mm starting thickness. Then a slender rod

was used to pierce through the foam to create room for passing a resistant nickel chrome wire. The hot

wire cut through the foam following the guide template (Fig. 3-6A); this process was repeated on two

orthogonal directions to obtain the integrated pyramidal sandwich structure (Fig. 3-6B).


26

Figure 3-6. 20 mm-thick foam block is carved with (A) hot-wire-cutting template into (B) integrated

pyramidal sandwich.

3.2. Lost Foam Casting

Once the desired foam board pre-forms were prepared, they were ready for coating. The coating

solution (HA Styroshield 6450) was mixed to uniformity with a specific gravity of 45 Baume (water was

added as needed) by a propeller. The pattern was dipped into the coating solution, taken out and excess

solution was spun away. The moist-coated pattern was hung dry in a circulating hot air oven at 60 ºC for

one day. If thermal analysis was necessary, a pin-sized hole was made on the pattern for inserting a K-

type thermocouple.

The coated pattern assembly was then placed below a metal pouring cup in a silo that was filled

with refractory sand. The whole silo was vibrated horizontally at 1 g acceleration for 30 s for compaction

to allow the sand to fill the pattern cavity. If thermal analysis was needed, the K-type thermocouples were

connected to a data acquisition unit.

In this project, aluminum alloy A356 and magnesium alloy AZ91 were chosen for lost foam

casting. A356.2 (‘.2’ refers to recycled) aluminum alloy and is widely used in sand and permanent mould

casting. It has excellent castability, good resistance to hot cracking and shrinkage and also exhibits good

fluidity [Davis ed., 1993]. The chemical composition was measured by the supplier with arc emission

spectroscopy, results given below in Table 3-4. AZ91 magnesium alloy is commonly used in die casting

but can also be used in sand, permanent mould and investment casting. Among different magnesium


27

alloys, it has high purity and good corrosion resistance [Avedesian and Baker ed., 1999]. The exact

chemical composition used in this study was also measured with arc emission spectroscopy, summarized

in Table 3-5.

Table 3-4. A356.2 Aluminum alloy chemical composition in wt% [House of Metals, 2008].

Si Mg Ti Fe Cu Zn Mn Pb Al

7.08 0.41 0.17 0.08 < 0.01 < 0.01 < 0.01 0.001 Balance

Table 3-5. AZ91D Magnesium alloy chemical composition in wt%.

Al Zn Mn Si Cu Fe Ni Be Mg

8.51 0.58 0.23 0.043 0.003 0.062 0.0014 0.0006 Balance

Before casting, the ingots were preheated at 200 ºC for two hours to remove any moisture. A steel

crucible with the dried ingots was placed in an electric resistance furnace for ~1.5 hour until the ingots

melted and the temperature reached ~800 ºC. If the furnace was used to melt magnesium, a CO2 cover gas

was required to avoid any oxidation and ignition. For aluminum melt, a degasser and flux were added at

~50 ºC and ~20 ºC above the pouring temperature.

Once ready for casting, the melt was poured directly into the pouring cup, (casting time was ~10s

for a volume of ~1000cm3). A fume hood was placed over the pouring cup to vent some of the gas foam

pyrolysis products. A data acquisition unit recorded the changes in resistance from the K-type

thermocouples every 0.02 s for 900 s until the casting solidified and cooled. If an applied vacuum was

necessary for casting, a plastic cover was placed over the sand silo and pouring cup with a vacuum pump

connected at the bottom of the silo. The required vacuum level was reached by adjusting the side vent

value to let air enter the silo.


28

3.3. Post-Casting Characterization

Once cooled, the casting was taken out from the sand silo, and loose sand and refractory coating

were removed. The casting was then sectioned from the sprue with a vertical band saw. Samples were

prepared for uniaxial compression testing by machining the outer face sheet surfaces parallel. A cutting

speed of ~265 rpm using a 1” four flute cutter at a feed rate of ~0.001 in/rev was found to be suitable for

machining both the aluminum and magnesium casting.

The porosity of the casting was measured using the Archimedes method [e.g. Bendick, 1995], the

density can be calculated by the substance’s change in weight in a liquid due to buoyant force, which is

equal to the weight of the displaced liquid. The mass of the sample was first measured on a scale in air;

the weight of water in a light container is measured as well; at the end, the weight of both liquid and the

sample submerges mid-level inside by suspending in a fishing wire. The volume of the sample can be

found from the weight difference of the sample with buoyancy force in water. Then the porosity of the

sample can by found with the known theoretical density in the following equation:

)(100

%wwm

wm

s

s MM

Mporosity

−−=

+

ρρ

ρ Eqn. 3-2

where ρs = theoretical density of sample (g/cm3), Mm = mass of sample (g), ρw = density of water at room

temperature (0.9982 g/cm3) and Mm+w = mass of water with sample (g).

Heat treatment was applied to as-cast A356 and AZ91 to determine the effect of microstructure

on the mechanical properties of the PCMs. The heat treatment temperature and time summarized in

Table 3-6 for aluminum alloy A356 and magnesium alloy AZ91 were used in this study. The air furnace

was preheated and adjusted to the required temperature by monitoring the temperature with a K-type

thermocouple in the centre of the furnace. The samples were placed in the centre of the air furnace with a

heat soaking time of 5~10 minutes. When the required heat treatment was over, the samples were taken

out and quenched in moving air.


29

Table 3-6. The heat treatment schedule for A356 [Davis ed., 1993] and AZ91 [Avedesian and

Baker ed., 1999].

Alloy Solutionized Solutionized and Artificially Aged

A356 540 ºC for 12 h 540 ºC for 12 h followed by 155 ºC for 4 h

AZ91 413 ºC for 16 h 413 ºC for 16 h followed by 168 ºC for 16 h

The mechanical properties of the PCMs were tested in uniaxial compression. A Shimadzu AG-1

tensile / compression machine (with a 50 kN load capacity and internal strain gage) was used. The test

was performed at constant strain rate of 1 mm/min and the force was measured every 0.05 s. A unload-

and-reload cycle was performed during testing in order to determine the PCM loading modulus at ~0.75

of the sample yield strength. The force and displacement were converted to stress and strain by the

theoretical PCM area and PCM core height.

Table 3-7. Etchants used in microscopic examination of A356 [Keller, 1948] and AZ91 [Maltais

et al., 2004] alloys.

Alloy Etchant Procedure

A356 2 ml HF (48 %), 3 ml HCl, 5 ml

HNO3 and 190 ml water

Immerse for 8 to15 seconds, wash in warm water

and blow dry.

AZ91 15 ml acetic acid, 10 ml water

and 75 ml ethanol

Immerse for 30 seconds, wash in ethanol and blow

dry.

Rockwell scale E (HRE) and Vickers (HV) hardness measurements were performed on samples

mounted in epoxy and diamond polished surfaces. The hardness value was computed based on the change

in depth or width of the indentation with the preset load. HRE hardness value is measured by using a 1/8”

steel ball indenter with 10 kg minor and 100 kg major load to create an indent on the surface for 10 to

15 s [Voort ed., 1999]. The hardness value is inversely proportional to the indenter penetration depth,

each increment of hardness corresponds to 0.002mm of penetration depth. HV hardness value is measured

by using a pyramidal shaped indenter with 100g load to create a diamond shaped indent on the surface for

10 to 15 s; HV can be calculated from the following [Voort ed., 1999]:

22

8544.1)2/sin(2

d

Lp

d

LpHV ==

δ (Eqn. 3-3)


30

where d = mean diagonal (mm), Lp = penetration load (kgf) and δ = face angle (136º). ASTM standard

E384-05 [Mayer ed., 2007] was followed by spacing each indentation with a minimum 2.5 diagonal

widths apart or with the edge of the sample. A minimum of 20 hardness measurements were completed;

an average and standard deviation values are given in the Results and Discussion Section.

3.4 Experimental Plan

A schematic representation of the Experimental Plan is shown in Fig. 3-7.

Figure 3-7. Schematic representation of experimental plan.

Initial Feasibility Study

Fabrication and Casting of the

Corrugated Core

Pyramidal Sandwich Fillability Study

Fabrication and Casting:

i, Sprue and Runner Designs

ii, Vacuum Assisted LFC

iii, Alloying Additions

Integrated Pyramidal Sandwich

Fabrication and Casting:

i, Effect of PCM Architecture

ii, Effect of Microstructure

Thermal Analysis

Fillability Analysis

Porosity Measurement

Hardness Measurement

Optical Characterization

Metallography

Hardness Measurement

Thermal Analysis

Porosity Measurement

Uniaxial Compression

Lower and Upper Bound

Model

Chapter 4: Results and Discussion

31

4. Results and Discussion

4.1. Initial Feasibility Study

The initial investigation uses a simple triangular corrugated panel to determine the feasibility of

casting PCM structures with LFC. The pouring temperatures of 710 °C and 730 °C, within typical AZ91

magnesium alloy casting temperature, were repeated twice as summarized in Table 4-1. Table 4-2 also

shows the dimensions of the corrugated core assembly. Fig. 4-1 presents the overview of the as-cast

panels, the backside of the panels (pointing into the page) were sectioned from the side-feeding sprues.

Both cases had excellent overall fillability but with minor corner misfills furthest from the sprue. Overall,

this morphology is similar to the typically observed convex profile of the advancing metal front through

the EPS in pressureless LFC [Liu et al., 2002]. There was relatively little difference in the panels between

the two casting temperatures, but the 710 °C cast sample displayed slightly better surface finish and

sharper corners.

Table 4-1. The casting parameters for the corrugated core assembly.

Pouring Temperatures (°C) Foam Properties Alloy

710 EPS with 0.021 g/cm

3 (1.3 pcf) AZ91D

730

Table 4-2. The dimensions of the corrugated core assembly.

Web Thickness tc (mm) Relative Density ρ~ (%) Pattern Size (mm) Sprue Dimensions (mm)

3 51 155 X 150 270 X 50 X 50


32

Figure 4-1. As-cast corrugated core with pouring temperatures of (A) 710 °C and (B) 730 °C.

The panels were sectioned at different locations to identify defects within the casting as seen

in Fig. 4-2. From the cross-section area of the panels, the morphologies of the node were not as sharp

as predicted in the schematics. The curvature of the panels and the cavity found near the node (Fig. 4-

2D) showed the morphology of the foam beads (approximately 1mm in diameter) and the effects of

foam machining. During fabrication, the foam beads were torn off from the foam panel by the rotating

chamfer tool. This resulted in foam beads delaminating from each other or caused dimples on the

surface by extra foam beads shredding off.

Figure 4-2. Cross-sectional node geometry from the as-cast corrugation core (A,B) poured at 710 °C and

(C,D) poured at 730 °C.

The cross-sections were polished to examine the level of porosity within the casting. Typical

examples are shown from the centre of the web (Fig. 4-3A) and at the edge (Fig. 4-3A) from the

730 °C cast sample. There was no significant difference between the two pouring temperatures. By the


33

morphology of the porosity indicated that a large percentage was shrinkage caused by rapid

solidification from low-melting magnesium alloy with incomplete decomposition of foam products.

The permeability of the coating, the foam physical property and the volume-to-surface-area ratio of

the pattern can influence the rate of filling, and foam degradation and removal of the decomposition

products [Lawerence et al., 1998]. These parameters could be adjusted later to minimize the level of

casting porosity.

Figure 4-3. Typical porosity at the centre (A) and edge (B) of the corrugated core at a pouring

temperature of 730 °C.

The polished sample was etched to reveal the typical dendrite morphology within LFC AZ91’s

microstructure as seen in Fig. 4-4. As-cast AZ91 exhibits a eutectic microstructure with lighter regions of

primary α and darker regions of Mg17Al12 within the eutectic microstructure, which matches with what is

found from literature [Avedesian and Baker ed., 1999].

Figure 4-4. Typical microstructure of as-cast AZ91D corrugated core and etched.


34

Microhardness measurements were obtained on the node cross-sections at regular intervals from

the sprue (10 mm, 50 mm, 90 mm, and 130 mm) and values are summarized in Table 4-3. There was no

apparent change in microhardness across the casting, nor any significant difference between the two

casting temperatures 710 °C (72 ± 7 HV) and 730 °C (76 ± 8 HV). Overall, these values were consistent

with those previously reported for as-cast AZ91D [Avedesian and Baker ed., 1999].

Table 4-3. Microhardness measurements as a function of distance from the sprue.

The initial casting study showed that it was possible to produce simple corrugated panels with

magnesium by LFC. It also showed that an alternative foam pattern fabrication technique would be

necessary to avoid damage to the foam beads. There was overall good fillability at both the 710 °C and

730 °C pouring temperatures but the corners of the panels showed the fill limit with the 3 mm web

thickness. Finally, the porosity and microhardness values found from this study indicated similar casting

properties to as-cast AZ91D. From the success in producing simple corrugated panels, the next step was

to create more complex and lower-relative-density PCMs by LFC and to characterize them in terms of

their casting qualities and mechanical properties.

4.2. Effect of Casting Conditions on LFC PCMs

In this study, the effects of sprue and runner designs, vacuum assistance and alloying additions on

the fillability of PCMs were studied. Different sprue and runner designs were developed to determine the

optimum filling of the pattern. The fillability of the pattern can be improved by delivering liquid metal

with minimum turbulence and heat loss. Different vacuum levels were applied during casting to increase

the metallostatic pressure, which improves transportation of foam degradation products and results in

Pouring Temperature (°C) 710 730

Distance from Sprue (mm) 10 50 90 130 Avg 10 50 90 130 Avg

Average Hardness (HV) 72 68 78 69 72 76 72 72 83 76

Standard Deviation (HV) 9 5 8 6 7 9 7 9 6 8


35

improved flow length. Different levels of alloying additions were also studied with AZ91D to widen the

solidification range and improve fluidity which also improves the flow length of the pattern.

4.2.1. Sprue and Runner Designs

The new pyramidal lattice pattern helped to further reduce the relative density (from ρ~ = 51 % to

22 %). The same side feeding with vertical sprue design was used as initial feasibility study, but

additional side runners was implemented to allow flow to lattice pattern in the centre (see Fig. 4-5A). The

casting conditions and pattern assembly dimensions are summarized in Table 4-4 and Table 4-5.

Figure 4-5. Pyramidal lattice core pattern (A) ready for casting and as-cast samples (B).

Since the lattice pattern has many 3 mm-thick struts and there was no direct metal flow from the

runner to each strut, this pattern was considered to be very difficult to fill. Therefore, aluminum alloy

A356 was first used with a casting temperature of 730 °C instead of AZ91 due to A356’s good castability

with LFC. As seen in Fig. 4-5B, the casting failed to fill, the side runners offered insufficient volume for

metal flow before rapid solidification. The combination of small cross-section area runners and 60º turn

struts resulted in lower metal velocity and shortens the flow length. From this failed sprue and runner

design, it was determined that struts would need to be filled by providing more direct metal flow.

Table 4-4. The casting parameters for the vertical sprue lattice core assembly.

Pouring Temperature (°C) Foam Properties Alloy

730 EPS with 0.021 g/cm3 (1.3 pcf) A356


36

Table 4-5. The dimensions of the vertical sprue lattice core assembly.

Web thickness tc (mm) Relative Density ρ~ (%) Pattern size (mm) Sprue dimensions (mm)

3 22 190 X 150 250 X 50 X 50

In order to improve fillability of the lattice pattern, a pair of face sheets was attached to the lattice

core to create a complete sandwich panel. The pair of face sheets can help provide metal flow directly to

each strut and eliminate post casting joining steps of the sandwich core to face sheets. The same sprue and

runner design (see Fig. 4-6A and Table 4-6) was used to determine the improvement in fillability due to

the additional face sheets. A356 was used at the same pouring temperature (730 °C) as AZ91,

summarized in Table 4-7. As Fig. 4-6B shows, there was good overall fillability on A356 sandwich

panel’s face sheets but there were numerous unfilled struts in the lattice core. The same pattern was cast

in AZ91 as well to determine the difference in flow ability between the two alloys (see Fig. 4-6C). AZ91

casting showed rather limited fillability on the face sheets and very low fillability within the core.

Figure 4-6. Pyramidal sandwich pattern (A) ready for casting, A356 as-cast sample (B) and AZ91 as-cast

sample (C).

In order to produce A356 cast sandwich panel, it was necessary to adjust minor casting

parameters to help minimize misfilled struts. Since the contact between face sheets and nodes of the struts

was obtained by hot melt glue, the adhesion was important for fillability but the use of hot melt glue also

decreases the metal velocity by having a lower vapourization rate than foam [Wang et al., 1990; Hess,


37

2004]. In contrast, AZ91 cast sandwich panel might require a different sprue and runner design that

improves pressure feeding within the pattern to solve the major fillability issue.

Table 4-6. The dimensions of the vertical sprue sandwich assembly.

Web thickness tc

(mm)

Face sheet

thickness hfs (mm)

Relative Density

ρ~ (%) Pattern size (mm)

Sprue dimensions

(mm)

3 5 22 150 X 150 250 X 50 X 50

Table 4-7. The casting parameters for the vertical sprue sandwich assembly.

Pouring Temperature (°C) Foam Properties Alloys

730 EPS with 0.021 g/cm3 (1.3 pcf)

A356

AZ91

A horizontal sprue design was developed to provide more direct metal flow to the sandwich panel

with the help of gravity and straight through feeding design. The dimensions of the sprue were the same

as the previous design; the pouring cup attaching on the sprue and the whole design was simply rotated

90º to the top, see Fig. 4-7A. This design was cast with AZ91 at 730 °C in order to keep the other casting

variables constant. The casting conditions and pattern dimensions are summarized in Table 4-8 and Table

4-9. Fig. 4-7B shows that the pouring cup had collapsed on the sprue by its own weight. The impact from

falling liquid metal has also caused the core to collapse and metal penetration through the ceramic shell to

the supporting sand. The collapse was caused by insufficient compaction of the refractory sand because

the sand could not fill the middle of the sandwich core properly from the sides only. The sprue design was

also alternated by tapering towards the sandwich panel, see Fig. 4-7C. The tapered sprue design with

smaller pouring cup was expected to reduce the chance of core collapse and metal penetration through the

ceramic shell. As shown in Fig. 4-7D, although the pouring cup did not collapse onto the pattern, the

tapered sprue design did not prevent pattern collapse within the truss core. The major flaw to this design

was the insufficient flow of refractory sand to the sandwich panel below the horizontal sprue. The

fillability problem was not reduced by this horizontal sprue design either; neither the face sheets nor core

struts were completely filled.


38

Figure 4-7. Pyramidal sandwich pattern assembly with top feeding sprue (A), AZ91 as-cast sample (B),

altered pattern assembly with tapered sprue (C) and AZ91 as-cast sample (D).

Table 4-8. The casting parameters for the horizontal sprue pattern assembly.

Pouring Temperature (°C) Foam Properties Alloys

730 EPS with 0.021 g/cm3 (1.3 pcf) AZ91

Table 4-9. The dimensions of the horizontal sprue pattern assembly.

Web thickness

tc (mm)

Face sheet thickness

hfs (mm)

Relative Density

ρ~ (%)

Pattern size

(mm)

Sprue dimension (mm)

3 5 22 170 X 170 (Square) 250 X 50 X 50

(Tapered) 170 X 20 X 40

From the various sprue and runner designs, vertical sprue produced the castings without pattern

collapses and refractory sand filling issues, a pair of face sheets helped to improve filling to the truss core

by producing direct flow into each strut.

4.2.2. Vacuum Assisted LFC

Since the horizontal sprue design impeded the flow of refractory sand and sand compaction, a

vertical sprue, having smaller cross-sectional area, was used, see Fig. 4-8A. The narrow vertical sprue

tended to minimize the amount of foam on the wall from the hollow sprue, which reduced the amount of

heat extracted from the liquid metal during foam degradation. The casting conditions and pattern


39

dimensions are summarized in Table 4-10 and Table 4-11. The casting, as shown from Fig. 4-8B, did not

show significant improvement in fillability compared to previous casting from Fig. 4-6C. The uneven

fillability between the face sheets could be affected by the amount of hot melt glue used in between each

contact area with the sprue.

Figure 4-8, Pyramidal sandwich assembly with narrow and lengthen sprue (A) and AZ91 as-cast

sample (B).

Table 4-10. The casting properties for the narrowed vertical sprue pattern assembly.

Pouring Temperature Foam Properties Alloy

730°C EPS with 0.021 g/cm3 (1.3 pcf) AZ91

Table 4-11. The dimensions of the narrowed vertical sprue pattern assembly.

Web thickness tc

(mm)

Face sheet

thickness hfs (mm)

Relative Density

ρ~ (%)

Pattern size (mm) Sprue dimension

(mm)

3 5 22 170 X 170 250 X 40 X 40

A range of casting conditions were tested in terms of pouring temperature and applied vacuum

level (Table 4-12) to improve the fillability. Samples S1, S2 and S3 were cast at a pouring temperature of

730 °C and at vacuum levels ranging from 0 to 40 kPa. Samples S4 to S6 were cast at a pouring

temperature of 750 °C with vacuum levels ranging from 20 to 40 kPa. Finally samples S7 and S8 were

cast at 750 °C and vacuum levels of 40 and 50 kPa, respectively, but were double coated with the

refractory slurry.


40

Table 4-12. Summary of casting conditions: the vacuum level, pouring temperature, and number

of coatings for the eight different casting conditions.

Sample Vacuum level (kPa) Temperature (°C) Number of coatings

S1 0

730

1

S2 14

S3 40

S4 20

750

S5 30

S6 40

S7 40 2

S8 50

The particular casting condition had a significant effect on the extent of face sheet and truss core

filling. Samples cast at 730 °C exhibited both incomplete core and face sheet filling, regardless of the

vacuum level (Fig. 4-9: S1, S2, and S3). Increasing the pouring temperature to 750 °C generally increased

the fillability (Fig. 4-9: S4, S5, and S6) and complete filling could be obtained for the 30 kPa (S5) and 40

kPa (S6) vacuum levels. However, applying vacuum also had the effect of inducing metal penetration

through the ceramic shell, which resulted in poor surface finish and sand adhesion. Double coating the

EPS pattern (S7 and S8) was attempted to create a thicker ceramic shell in order to prevent metal

penetration. The core filling of sample S7 was significantly reduced compared to the equivalent single

coated sample (S6). Complete face sheet and core filling could be achieved in the double coated condition

by increasing the vacuum level to 50 kPa (S8). However, the surface finish was only marginally

improved. Fig. 4-10 and Fig. 4-11 summarize the face sheet and core filling for all eight samples. As

would be expected, it was significantly easier to completely fill the face sheets than the pyramidal core.


41

Figure 4-9. The particular casting condition for the eight samples (summarized in Table 4-12) had a

significant effect on the extent of fillability. Complete pattern was obtained in samples S5, S6, and S8.

0

20

40

60

80

100

S1 S2 S3 S4 S5 S6 S7 S8

% F

ille

d T

ru

sse

s

0

20

40

60

80

100

S1 S2 S3 S4 S5 S6 S7 S8

% F

ille

d F

ace

Sh

ee

t

Figure 4-10. Percentage of truss core (left) and face sheets (right) filled for the eight different casting

conditions (Table 4-12).

A B C D

E F G H

S1 S2 S3 S4

S5 S6 S7 S8


42

0

20

40

60

80

100

60 70 80 90 100

% F

ille

d T

ru

sse

s

% Filled Face Sheets

S1: gravity cast at 730°C

S2: 14kPa vacuum at 730°C


S6: 40kPa vacuum at 750°CS5: 30kPa vacuum at 750°C


S7: 40kPa vacuum at 750°C 2 coats

S8: 50kPa vacuum at 750°C 2 coats

Figure 4-11. Casting map summarizing the extent of face-sheet and truss core filling.

Porosity and microhardness measurements were obtained from three different positions in the S1,

S3, S6, and S7 samples. Sample S1 showed a slight increase in porosity as distance increased from the

sprue where the cooling rate was slowest; this is likely related to proper metal feeding in the near sprue

region before interdendritic solidification. The overall porosity values were similar at ~3 % for the S1, S3

and S7 samples (Table 4-13). The fact that sample S7 had a lower porosity than sample S6 may be related

to the extra slurry coating given to S7, which would be expected to reduce the rate of heat transfer to the

surrounding sand. However, sample S6 also exhibited some sand adhesion due to metal penetration

through the ceramic shell, which may have an effect on the measured porosity value. The morphology of

the porosity was studied by scanning electron microscopy of polished surfaces; similar pore

morphologies, which were consistent with interdendritic porosity, were seen for all four samples (Fig. 4-

12). There was no clear evidence of spherical porosity in any of the samples which might indicate gas

porosity.

Microhardness measurements were obtained after following ASTM E92-82 [Mayer ed., 2007] on

face sheet cross-sections at 3 different positions in each of the 4 samples. The values are summarized in

Table 4-13. There was little difference in the average microhardness between samples cast using the four


43

different conditions and the measured values are consistent with the reported hardness value of 70 HV for

as-cast AZ91D [Avedesian and Baker ed., 1999].

Table 4-13. Summary of porosity and microhardness characterization.

Sample Porosity (%) Hardness (HV)

S1: cast at 730°C in gravity 3.4 ± 1.9 66 ± 4.9

S3: cast at 730°C in 40 kPa vacuum 2.8 ± 1.3 66 ± 4.5

S6: cast at 750°C in 40 kPa vacuum 7.0 ± 3.1 65 ± 5.3

S7: cast at 750°C in 40 kPa vacuum with 2 coats 3.5 ± 1.7 65 ± 6.5

Figure 4-12. SEM images showing exposed porosity on polished surfaces of sample S1 (A), sample S3

(B), sample S6 (C), and sample S7 (D).

Fillability was a significant issue in magnesium LFC with PCMs, the reduced cross section of the

struts resulted in rapid solidification. Using higher casting temperatures and with an applied vacuum, it

was possible to completely fill the PCM pattern at the expense of some sand adhesion. This sand adhesion

prevented any accurate mechanical property to be obtained from the samples.

4.2.3. Alloying Additions

Since the vacuum assist LFC did not improve PCM fillability significantly, pure aluminum was

added to improve castability by broadening the solidification range [Avedesian and Baker ed., 1999] or

100µm

(a) (b)

(c) (d)


44

aluminum-silicon master alloy was added to increase fluidity [Sheng et al., 2009]. Probeads 70 foam was

used instead of EPS (see Table 4-14) to improve the fillability by lowering the volatile temperature and

lowering the decomposition energy [Shivkumar, 1994], the pattern size was slightly shortened and the

sprue was slightly lengthen (see Fig. 4-13 and Table 4-15) to minimize misruns as seen previously. In the

first series of experiments, pure aluminum was added to AZ91D to increase the aluminum content by

~3 wt% (close to maximum solubility of α Mg-Al phase, ~11.5 wt% Al) and by ~14 wt% (closer to

eutectic composition of α and β phase, 32 wt% Al). In a second series of experiments, various amounts of

Si (~0.5, ~1.0 and ~1.5 wt%) were added to AZ91D using Al-50%Si master alloy. See Table 4-16 for the

chemical composition resulted from alloying additions, the bold values were altered by the alloying

additions.

Thermal analysis was performed using thermocouples positioned at six locations in the pattern, as

indicated in Fig. 4-14, to study the cooling and solidification characteristics of each alloy. The cooling

curves were recorded for all alloy compositions. The first derivative (dT/dt) of these curves was also

determined (Fig. 4-18). Thermal analysis was also used to determine solidification time and velocity

within the panel.

Table 4-14. The casting properties for the narrowed vertical sprue pattern assembly.


750°C Probeads with 0.025 g/cm3 (1.5 pcf) AZ91

Table 4-15. The dimensions of the narrowed vertical sprue pattern assembly.

Web thickness tc

(mm)

Face sheet

thickness hfs (mm)

Relative Density

ρ~ (%) Pattern size (mm)

Sprue dimension

(mm)

3 5 22 85 X 170 450 X 40 X 40


45

Figure 4-13. The assembled coated pattern (A) and sandwich panel in close up (B).

Table 4-16. Measured chemical composition by spark emission spectroscopy for each casting

condition in wt%.

Mg Al Zn Mn Si Cu Fe =i Be

S1 90.62 8.51 0.58 0.23 0.043 0.003 0.062 0.0014 0.0006

S2 87.10 12.1 0.55 0.21 0.021 0.003 0.0069 0.0013 0.0008

S3 76.06 23.14 0.50 0.19 0.053 0.005 0.0378 0.0017 0.0009

S4 89.38 9.23 0.57 0.23 0.572 0.003 0.0077 0.0013 0.0011

S5 89.57 8.81 0.57 0.21 0.824 0.003 0.0052 0.0013 0.0011

S6 89.11 8.76 0.55 0.18 1.389 0.003 0.0060 0.0012 0.0006

Figure 4-14. Schematic representation showing the position of thermocouples in the pattern (top: T1, T2,

T3 and bottom: B1, B2 and B3).

Fig. 4-15 shows photographs of the as-cast structure for each alloy composition. For the S1

(reference sample), the face sheets were completely filled, while many struts were unfilled in the outer

edge of the casting. With an addition of 3 wt% Al (S2), the casting showed slightly more filled struts on

the outer edge as compared with S1. Further increase in Al content (S3) resulted in the complete filling of

1 2 3

1 2 3

Top

Bottom


46

the casting. However, severe metal penetration through the ceramic shell was seen inside the sandwich

core.

With the addition of 0.6 wt% Si to AZ91D in S4, the casting showed slightly improved filling of

the internal struts as compared with S1. The extent of filling was similar to that seen for S2. As the weight

addition of Si increased from 0.6 to 0.8 wt% (S5), there was no noticeable change in filling. With 1.4 wt%

Si addition (S6), only one corner of the casting did not fill. With the highest Al (15 wt%) or Si (1.4 wt%)

additions, the entire casting was improved significantly in terms of filling.

Figure 4-15. The final casting of magnesium PCMs with different alloying additions.

Fig. 4-16 represents the maximum temperature and the time to reach this temperature as a

function of the location in the casting. The melt was poured at 750 °C and it was observed that the

temperature decreased to ~700 °C as the melt reached the bottom of the sprue (B3), as shown in Fig. 4-16

on the left. As the melt entered from the sprue and flowed into the sandwich panel, the maximum

temperature dropped to ~620 °C along the panel adjoining to the sprue (T2) and (B2). The temperature

decreased even further to ~580 °C as the melt continued to flow through the 5 mm-thick panel and into

the 3 mm-thin struts. The decrease in temperature was attributed to the heat transfer from the metal to the

S3: AZ91D + ~15 wt% Al

S2: AZ91D + ~3wt% Al S1: AZ91D

S4: AZ91D + ~0.6 wt% Si

S5: AZ91D + ~0.8 wt% Si S6: AZ91D + ~1.4 wt% Si


47

surrounding materials and thermal degradation of the foam. The decrease in temperature from T3 to T2

was almost twice that from T2 to T1. Assuming that the heat loss to the surrounding materials was similar

across the panel, and that the distances between the thermocouples were the same, the temperature loss

results were attributed to the fact that a greater amount of heat from the melt between the sprue and

sandwich panel is required to degrade the hot melt glue, as compared with that required to degrade the

foam.

The order of filling at the different locations can be determined by the time taken for each to

reach the maximum temperature (Fig. 4-16 on the right). As the melt was poured, it first entered the

bottom of the sprue (B3) and then filled to the top of the sprue (T3). The melt slowly spread to the panel

between T2 and B2 in 3 s. Then, the melt flowed across the panel to the edge between T1 and B1. The

difference in time needed to reach the maximum temperature with respect to the distance between the top

and bottom parts of the panel corresponded to a melt flow velocity of ~1.66 cm/s at the bottom of the part

(B2 to B1), as compared to ~1.06 cm/s at the top (T2 to T1). The difference in velocity could be due to

the pressure gradient between the top and bottom parts of the panel.

Fig. 4-17 shows the solidification time and cooling rate at different locations in the casting. Due

to the asymmetrical feeding of the sprue design, there was a difference in cooling rate and solidification

time across the sandwich panel because part of the heat from the melt was transferred to the surrounding

material at regions 2 and 3, see Fig. 4-14. As the melt flowed across the panel, the melt temperature

dropped significantly and had a very short solidification time of ~50 s at the edge (T1 and B1) compared

with more than ~250 s at the panel along the sprue (T2 and B2) shown in Fig. 4-17 on the right. The

cooling rate was significantly different at the edge (T1 and B1) as compared with that in the sprue (T2

and B2) and panel near the sprue (T3 and B3), see Fig. 4-17 on the left. The fast cooling rate at the edge

of the face sheet prevented sufficient feeding of the struts.

A section of the sprue from each casting was examined for changes in microstructure with the

alloying addition. As-cast AZ91 contains a eutectic microstructure with lighter region of primary α-Mg


48

and darker region of β- Mg17Al12 phase [Avedesian and Baker ed., 1999] which is consistent with S1

(contains small amount of porosity) shown in Fig. 4-18. The corresponding cooling and first derivative

curves are shown next to the microstructure to illustrate the consistency in casting conditions between

samples. S2 alloy (3 wt% Al addition) exhibited a similar microstructure as S1. With further addition of

Al (S3), there was a large increase in the amount of β, Mg17Al12 phase compared to reference sample.

Significant spherical porosity was also observed throughout the sample. A possible source of gas porosity

would be severe gas absorption [Davis ed., 1993] with high aluminum content and at the high superheat

(~370 °C) but without proper hydrogen removal procedure. Mg2Si has a dark grey chinese script

morphology found within magnesium alloys [Avedesian and Baker ed., 1999] which is consistent with

S4, S5 and S6 alloys. These alloys also showed small amounts of porosity in all samples.

500

600

700

800

T1 T2 T3 B1 B2 B3

Maxim

um

Tem

pera

ture

(°C

)

Thermocouple Location

6(A) Temperature

0

5

10

15

20

T1 T2 T3 B1 B2 B3

Tim

e to reach the m

axim

um

te

mpera

ture

(s)


6(B) Time

Figure 4-16. Temperature (left) and time (right) profiles: T1-T3 (top of panel) and B1-B3 (bottom of

panel).

0

50

100

150

200

250

300

350

400

450

T1 T2 T3 B1 B2 B3

Solidific

ation T

ime (s

)


7(A) Solidification Time

0

1

2

3

4

T1 T2 T3 B1 B2 B3

Cooling R

ate

(°C

/s)


7(B) Cooling Rate

Figure 4-17. Solidification time (left) and cooling rate (right) profiles: T1-T3 (top of panel) and B1-B3

(bottom of panel).


49

Considering now the cooling curves in Fig. 4-18, each sample contained one peak at ~430 °C

which was consistent to the temperature formation of the β- Mg17Al12 phase in Mg-Al phase diagram

[Avedesian and Baker ed., 1999]. The temperature of formation for Mg2Si is higher (1085 °C) [Avedesian

and Baker ed., 1999] than the temperature of the melt which explained why no peak was detected during

casting. The initial formation of α Mg-Al phase could not either be detected because when the melt

flowed across the panel, the melt temperature was already below the AZ91 liquidus (630 °C) [Avedesian

and Baker ed., 1999].

S4: AZ91D + ~0.6 wt% Si

Mg17Al12

Mg2Si

S2: AZ91D + ~3 wt% Al

S3: AZ91D + ~15 wt% Al

Mg17Al12

S1: AZ91D Reference

Mg17Al12

Mg17Al12

100 µm


50

Figure 4-18. Optical micrographs (at same magnification, see scale bar from sample S4 or S6) and

thermal analysis of the castings.

Fig. 4-19 presents the porosity levels measured by Archimedes method for each alloy. All of the

casting except S3 had ~5 % porosity. The high porosity measured in S3 could be due to hydrogen gas

from the relatively high content of aluminum without the use of aluminum degasser, as explained

previously. S6 has the best fillability based on the least number of unfilled struts and comparable porosity

level with S1. S6 represents the best compromise in term of castability.

Figure 4-19. The measured porosity levels with Archimedes principle.

S5: AZ91D + ~0.8 wt% Si

Mg2Si

Mg17Al12

Mg2Si

Mg17Al12

S6: AZ91D + ~1.4 wt% Si

100 µm


51

Hardness measurements were performed following ASTM E92-82 [Mayer ed., 2006] on the sprue

of each sample. The results are presented in Fig. 4-20. A hardness value of ~77 HRE was found for S1

and was comparable to previously published data (HRE 75) [Avedesian and Baker ed., 1999]. Hardness

was found to increase with the addition of 3 wt% Al (S2). It was impossible to do accurate hardness

measurement on the S3 sample, due to the presence of significant gas porosity. With Si addition (S4, S5

and S6), there was no apparent change in hardness as compared with AZ91D.

Figure 4-20. Hardness of the cast samples.

Aluminum and silicon additions were used in this project to improve fluidity of AZ91D

magnesium alloy. The results suggested that 1.4 wt% silicon addition showed the best compromise in

terms of porosity and fillability. With this composition, it was possible to produce magnesium PCMs with

very thin sections.

4.2.4. Preliminary Compressions Test Results

Preliminary mechanical testing in uniaxial compression was conducted on filled sections of the

gravity cast (S1) from section 4.2.2. since any vacuum assist LFC PCMs had sand adhesion due to metal

penetration through ceramic shell. A typical test coupon is shown in Fig. 4-21 (2×3.5 pyramidal truss core

cells). The compression result is shown in Fig. 4-22, the stress first increased linearly with strain, then


52

reached a maximum (peak strength) and finally decreased to a plateau. The series of decreasing peaks

after the initial maximum corresponded to strut member fractures and delaminations between the struts

and the face sheets, the onset of which was observed during loading. Two distinct types of failure were

observed after the initial elastic region. In the first case, failure of the struts occurred at a peak strength of

~10 MPa (10.2 MPa for the sample shown in Fig. 4-22). In the second case, failure occurred at only

~5 MPa. The peak strength in this case corresponded to early fracture at the node / face sheet interface

(Fig. 4-23); the nodes that had failed were only partially complete, suggesting that this type of failure

mechanism is an artifact of sample preparation. Fillability was a significant issue in the lost foam casting

of magnesium alloy PCMs; the reduced cross section of the struts resulted in rapid solidification.

Figure 4-21. Typical compression test coupon (2×3.5 pyramidal truss core cells).

Figure 4-22. Typical uniaxial stress strain curves showing the mechanical behaviour with and without

sample preparation edge effects.


53

Figure 4-23. Progressive node/face sheet fracture during compression testing (image labels correspond to

the testing points shown on Fig. 4-22).

With pure aluminum additions, S2 was chosen for compression loading because it did not have

the level of porosity found in S3. With aluminum-silicon additions, 1.4 wt% silicon added AZ91 (S6)

showed the best compromise in terms of porosity and fillability. The previous gravity cast PCM without

edge effect (as S1) was used to compare as typical stress-strain curve (see Fig. 4-24) for the AZ91 PCM

with AZ91+ 3 wt% Al and AZ91 + 1.4 wt% Si. There were some variations in the characteristic of the

loading curves among the samples within the same casting. The average of 3 peak strengths in each

conditions were presented as 8.7 MPa ± 2.1 for S1, 9.2 MPa ± 1.8 for S2 and 8.2 MPa ± 1.5 for S6.

Figure 4-24. Comparisons between different AZ91 PCMs in compressions.

Sample-to-sample variability from within each casting was larger than the effect of alloying

addition on the mechanical properties and may be due to casting defects (porosity) or the foam structure

A B

C D


54

defects (uneven strut thickness). In addition, the defects caused by the hot melt glue could cause

delaminations between the core strut during loading. Therefore, the peak stresses between the different

alloy compositions could not be meaningfully compared.

Overall, this part of the study used a light alloy approach to achieve a high specific strength ratio

at a high relative density of 22 %. The ~8.7 MPa peak strength of AZ91 PCMs corresponded to a specific

strength of 23 MPa/(g/cm3), which is comparable to the specific strength of 24 MPa/(g/cm

3) for

investment cast Cu-2%Be (having a relative density ~2.25 %) [Chiras et al. 2002].


55

4.3. Integrated Pyramidal Sandwich

In this study, a new pattern was developed to eliminate the use of hot melt glue between the face

sheets and struts. This design reduced the amount of heat extracted and gas evaporated while filling the

struts to improve the fillability of the pattern. A lower strut thickness could then be used without misrun

issues. This study used 4 different strut thicknesses (1.5 mm, 2.0 mm, 2.5 mm, 3.0 mm) to determine the

degree of fillability and the effect of relative density (11 %, 12 %, 14 % and 16 %) on the compressive

mechanical properties. For each strut thickness design, a special template was machined to produce four

foam patterns that were attached to each sprue (see Fig. 4-25A) for casting in A356 aluminum and AZ91

magnesium alloys (see Fig. 4-25b). The casting parameters were held constant (Table 4-17) and the only

variable in this study was the strut thickness (see Table 4-18) resulting in different relative densities. Four

castings were produced for each alloy, each with a different relative density. Within each casting, four

sandwich panels were sectioned and some were heat treated (solutionized or solutionized and aged) to

determine differences in mechanical properties. Then each sandwich panels were sectioned to 2-by-2

struts for uniaxial compressive loading.

Figure 4-25. Integrated sandwich assembly (A) and AZ91 as-cast sample (B).

Table 4-17. Summary of casting parameters for integrated sandwich.


750 °C Probeads with 0.021 g/cm3 (1.5 pcf) AZ91, A356


56

Table 4-18. The dimensions of integrated sandwich pattern assembly.

Web thickness tc

(mm)

Face sheet

thickness hfs (mm)

Relative Density

ρ~ (%)

Pattern size (mm) Sprue dimension

(mm)

1.5, 2.0, 2.5, 3.0 5 11, 12, 14, 16 ~70 X ~60 350 X 40 X 40

4.3.1. A356 PCMs Compression Test Results

The typical compressive loading curves for A356 PCMs are shown in Fig. 4-26 (left). As the

PCMs compressed and passed their yield strengths, the strut members continued to plastic deform and

their cross-section area increased with no observable crack initiated or propagated (Fig. 4-27B). During

compression, the PCMs were unloaded and loaded again to determine if there was any change in PCMs’

loading modulus as they were plastically deformed. During the entire compression, A356 PCMs showed

continuous and linear increase in stress but the rate of increase in stress reduced to a steady rate (Fig. 4-

26) after yielding. The photographed strut morphology showed the surface texture of the as-cast strut was

largely influenced by surface texture of the foam pattern. The surface of the as-cast strut had great surface

details, the gaps between poorly fused foam bead could be seen.

Figure 4-26. Stress-strain curves (left) and rate of change in stress with strain (right) for typical A356

PCMs in compression with different relative density ρ~ .


57

Figure 4-27. Strut morphology initially (A) and after compression testing (B).

A summary of the different compressive yield strengths of the A356 PCMs as a function of

relative density is shown in Fig. 4-28; the A356 PCMs’ yield strength increased linearly with relative

density due to increased strut cross sections. Heat treatments were also applied to some of the PCMs prior

to compression as an attempt to improve their compressive properties by modifying the A356

microstructure. The compressive stress-strain curves of the solutionized, and the solutionized and aged

samples were similar to as-cast as shown in Fig. 4-28, when the PCM yield strength increased with

relative density linearly. A356 PCMs’ yield strengths were summarized with different relatives and heat

treatments in Fig. 4-29. The solutionized PCMs (Fig.4-29 middle) demonstrated slightly lower yield

strength compared to the as-cast PCMs (Fig. 4-29 left), and the solutionized and aged PCMs (Fig. 4-29

right) showed the highest yield strengths out of all three.

Fig 4-28. Stress-Strain curves for typical solutionized (left) and solutionized and aged (right) A356 PCMs

in compression with different relative density ρ~ .


58

Figure 4-29. A356 PCM’s yield strength (0.2 % offset) with respect to relative density ρ~ : as-cast (left),

solutionized (middle), and solutionized and aged (right).

As-cast, and solutionized and aged A356 PCMs’ strains increased linearly with relative density at

which they yield at 0.2 % offset as shown in Fig. 4-30. In contrast, the solutionized PCMs behaved as

opposite, the strains decreased linearly with relative density. A series of unloading and reloading during

compression were applied to determine PCMs’ unloading modulus as they were deformed. The unloading

modulus increased with strains and deformations but there was no clear difference between heat

treatments found in Fig. 4-31. Fig. 4-32 showed no clear relationships between PCM’s relative density in

each heat treatment conditions. The linear increase in unloading modulus was related to the amount of

deformations, strain or change in strut angle with deformation introduced to PCMs.

Figure 4-30. A356 PCMs’ strains at which they yield at 0.2 % offset with respect to relative density after

different heat treatments.


59

Figure 4-31. A356 PCMs’ modulus with respect to strains after different heat treatments.

Figure 4-32. A356 PCMs’ modulus with respect to strains and different relative densities: as-cast (left),


Three etched microstructures were presented in Fig. 4-33 to show the difference after heat

treatments. Fig. 4-33A had typically dendritic structure found in as-cast A356, the dark regions are silicon

precipitates. Fig. 4-33B had the typical microstructure found after solutionized heat treatment, there was

less amount of silicon precipitates (dark shade regions) found between the dendrite arms (light grey

regions). Fig. 4-33C showed the change of more finely dispersed silicon precipitates around the structures

after solutionized and aged.


60

Figure 4-33. A356 etched microstructures in as-cast (A), solutionized (B), and solutionized and Aged (C).

Different aluminum alloys in investment cast PCMs [Desphande and Fleck, 2001; Zhou et al.,

2004], extruded PCM [Queheillalt et al., 2008] and stretch formed PCMs [Kooistra et al., 2004] were

compared with the present results presented in Fig. 4-34. The yield strengths from aged PCMs [Kooistra

et al., 2004; Queheillalt et al., 2008] demonstrated higher increase in slope with relative density compared

to the rest. The present study had higher PCM density due to thicker strut designs. The trend of increasing

yield strength with relative density was consistent between present study and the rest. Yet the aged PCMs

in present study did not show as much increase as seen in data from Kooistra et al., 2004.

Figure 4-34. A comparison of different aluminum alloy PCMs and the current studied LFC PCMs.


61

4.3.2. AZ91 PCMs Compression Test Results

In contrast, AZ91 PCMs behaved very differently beyond yielding during compression compared

to A356 PCMs. In as cast condition, the compression results from AZ91 PCMs were presented in stress-

strain curves in Fig. 4-35 (left), the cracks initiated (Fig. 4-36B) and propagated within the struts, and

fractured at peak stress on the plane normal to the aligned strut (Fig. 4-36C). After the stress dropped

rapidly, it plateau as the fractured core compressed until face sheets were deformed then stress rose again

(Fig. 4-36D). The rate of change in stress was rapid and unsteady after the peak load, this can be seen in

Fig. 4-35 (right). This might be affected by the surface detail left behind from the poor fused foam beads

acting as stress concentrations for crack initiation. AZ91 PCMs were solutionized or solutionized and

aged as well, the compression results showed similar behaviours as as-cast AZ91 PCMs, shown in Fig. 4-

37.

Figure 4-35. Stress-strain curves (left) and change in strain (right) for typical AZ91 PCMs in compression

with different relative densities ρ~ .

Figure 4-36. Struts morphology initially (A), after peak (B), expanded view of the fractured struts after

peak (C) and after stress plateau (D).


62

Fig 4-37. Strain-strain curves for typical solutionized (left) and solutionized and aged (right) AZ91 PCMs

in compression with different relative densities ρ~ .

AZ91 PCMs’ yield strengths increased with relative density as summarized in Fig. 4-38, the line

of best fits showed a linear relationship similar to A356 as-cast PCMs. Heat treatments were also applied

to AZ91 PCMs as well, the yield strengths are higher with solutionized and aged, followed with as-cast

and solutionized samples similar to A356 PCMs. Solution heat treatment able to improve the maximum

toughness and with following artificial aging, the maximum hardness and yield strength increased but

lowered the ductility [Avedesian and Baker ed., 1999].

Figure 4-38. AZ91 PCM’s yield strength (0.2 % offset) with respect to relative density ρ~ : as-cast (left),


In Fig. 4-39, the amount of strains introduced to AZ91 PCMs when they yield at 0.2 % offset

increased with relative density for as-cast and solutionized and aged samples. Yet solutionized samples

behaved in opposite, the amount of strains decreased with relative density when they yielded. The same


63

behaviours were found in A356 PCMs as well. In contrast, AZ91 PCMs deformed and collapsed after

certain peak stresses which were not found from A356 PCMs. The amount of strains required for peak

stress increased with relative density as seen in Fig. 4-40 for all three heat treatment conditions. The

solutionized PCMs required very different amount of strains for yielding and peak stresses as relative

density increases. In yielding, the amount of strains required decreased but the amount of strains required

increased with peak stress.

Figure 4-39. AZ91 PCMs’ strains at which they yield at 0.2 % offset with respect to relative densities

after different heat treatments.

Figure 4-40. AZ91 PCMs’ strains at peak stress with respect to relative density after different heat

treatments.

Unlike A356 PCMs, AZ91 PCMs’ unloading modulus were scattered across the different strains

and there were no correlations between heat treatments from Fig. 4-41. In Fig. 4-42, each heat treatment


64

also showed no clear relationship between unloading modulus with strains on different relative densities.

This might be affected by the crack initiation and propagated around 0.05~0.1 strain at peak stress. Fig. 4-

43 showed the difference in microstructures with different heat treatments; the as-cast microstructure

showed typical equiaxed dendritic microstructure with lighter region of primary α-Mg matrix and darker

region of β- Mg17Al12 phase [Avedesian and Baker ed., 1999]; the solution treatment dissolved all the β-

Mg17Al12 lamella around grain boundaries but all its spheroidal precipitates remained within the grains

which showed similar microstructure as found literature [Fujii et al., 2007]; the aged treatment caused all

the β- Mg17Al12 to reappeared as larger clusters along grain boundaries also found similar to the literature

[Fujii et al., 2007].

Figure 4-41. AZ91 PCM’s modulus with respect to strains after different heat treatments.

Figure 4-42. AZ91 PCMs’ modulus with respect to strains and different relative densities: as-cast (left),



65

Figure 4-43. AZ91 etched microstructures in as-cast (A), solutionized (B), and solutionized and Aged (C).

Since there was no magnesium PCMs produced and studied to date, a comparison of different

magnesium foam was made with the current study in Fig. 4-44. AZ91 Foams [Yamada et al., 1999;

Yamada et al., 2000] has very low density but the yield strengths are relatively low compared to LFC

PCMs. The mechanical difference between the present studied PCMs and previously studied metal foams

is due to its stretch-dominated mechanisms and bending-dominated mechanisms respectively [Ashby,

2005]. The effectiveness of the foam structure decreases with high relative density foams [Wen et al.,

2004] compared to bulk yield strength of ~90 MPa in AZ91 and ~21 MPa in pure magnesium [Avedesian

and Baker ed., 1999].

Figure 4-44. A comparison on yield strength with respect to density between magnesium foams and

present study of AZ91 LFC PCMs.


66

4.3.3. Lower and Upper Bound Model

A theoretical model was developed to predict the compressive property of PCMs, assume the

PCM’s strut would failure at the minimum cross section regions by yielding (see Fig. 4-45). This lower

bound model was derived by using the material yield strength with oriented compressive force and cross-

section area with the following formula,

φ

σσ

SI'A

'A

PCM

ys

PCMys

0= (Eqn. 4-1)

where σys PCM = compressive yield strength of PCM, σys = compressive yield strength of material, A0 =

minimum cross-section area of the strut, ' = number of struts with each PCM under compression, APCM =

cross-section area of PCM under compression, φ = Strut angle (60º) and tc = strut thickness. And the

minimum cross-section strut area can be found from the following equations as well,

φφφ 2

2

2

2

0

42

2

2

TA'

t

SI'

t

SI'

tA ccc −= (Eqn. 4-2)

Figure 4-45. Lower bound model based on minimum strut cross-section thickness, tc.


67

The material compressive yield strength could be found from literature but with variations

depending on the processing parameters. Therefore, individual square blocks (length to diameter ratio

below 2 to avoid buckling [Faupel and Fisher, 1981]) were machined from the PCM’s casting sprue to

determine the closest compressive property of A356 and AZ91 with different heat treatments. The same

stress-strain curve behaviour can be seen in A356 square compression blocks (Fig. 4-46) with A356

PCMs. The aluminum blocks would yield and deformed continuously with increased stress until the

equipment’s load limit. Within the tested compressive stress, the aluminum blocks showed cracks

initiated and propagated on the block’s surface and the height was shortened with width expanded as seen

in Fig.4-47 and Fig. 4-48.

0

200

400

0 0.2 0.4

ε

σcomp (MPa)

A356 As Cast

A356 Solutionized

A356 Solutionized and Aged

Figure 4-46. Stress-strain curve of A356 square compression blocks with different heat treatment

conditions.

Figure 4-47. A356 sample block after compression (A) and magnified surface feature of the block (B).


68

Figure 4-48. A356 square compression block`s surface with crack propagations.

The results from the stress-strain curves of AZ91 compression blocks (Fig. 4-49) confirmed the

similar deformation behaviour found in AZ91 PCMs, the sample would deformed to a certain strain and

fractured. AZ91 square compression block would failure by shear at 45º to the loading axis. The fracture

surface (Fig. 4-50 and 4-51) demonstrated many parallel plateau and ledges that matched to transgranular

cleavage found in [Hertzberg, 1996]. It was transgranular rather than intergranular cleavage fracture

because the crack did not follow 3-dimensional grain morphologies. This was different from A356 square

compression block, A356 did not fracture and only showed minor cracks initiated and propagated on the

surface (Fig. 4-47 and 4-48).


69

0

200

400

0 0.1 0.2 0.3

ε

σcomp (MPa)

AZ91 As Cast

AZ91 Solutionized

AZ91 Solutionized and Aged

Figure 4-49. Stress-strain curve of AZ91 square compression blocks with different heat treatment

conditions.

Figure 4-50. AZ91 sample block’s fractured cross section after compression (A) and its magnified surface

(B) showing transgranular fractures with parallel plateau and ledges.


70

Figure 4-51. AZ91 square compression block fracture surface after 45° fracture.

The yield strengths from A356 and AZ91 compression blocks with different heat treatments are

shown in Fig. 4-52, the solutionized AZ91 and A356 compression blocks were found to be higher than as-

cast conditions (Fig. 4-52), which was conflicted with their PCMs.


71

Figure 4-52. Compressive yield strength of A356 and AZ91 compression blocks with different heat

treatments.

With the compression block results, the lower bound model was developed with the determined

yield strength with different heat treatments in AZ91 and A356. Fig. 4-53 showed the difference between

experimental data found in A356 PCMs and lower bound model found from Eqn. 4-1 and Eqn. 4-2. Both

line of best fits showed similar slopes which proved that the change in strut thickness was related to yield

strength required for its derived minimum cross-section area. The gaps between the experimental data and

lower bound model can be explained by difference between the theoretical and actual strut thickness. The

gaps were not consistent between the three different heat treatments, the solutionized PCMs tended to be

lower than predicted; their yield strength has reduced significantly compared to their bulk samples.

Similar trend is shown in Fig. 4-54, AZ91 with increasing yield strength with relative density. The gap

between the experimental data and lower model was more consistent throughout different heat treatments

compared to A356. By comparing the yield strength of compression blocks to PCMs in A356 and AZ91,

the compression blocks showed slightly higher yield strengths with A356 relative to AZ91 but PCMs’

yield strengths between the two alloys show very similar values.


72

Figure 4-53. The comparisons between summarized experimental A356 PCMs’ yield strengths with lower

bound model: as-cast (left), solutionized (middle), and solutionized and aged (right).

Figure 4-54. The comparisons between summarized experimental AZ91 PCMs’ yield strengths with

lower bound model: as-cast (left), solutionized (middle), and solutionized and aged (right).

An upper bound model was also developed by predicting PCM yield strengths with solid

compressive yield strength results and average strut thickness. The average strut thickness accounted by

redistributing the whole strut volume with radius of curvature into uniform thickness throughout the

struts. Since the strut thickness is uniform, upper bound model predicts stress required yield for that cross

section at any location on the strut. The light grey trend lines in Fig. 4-53 and Fig. 4-54 show the

difference in compressive yield strength results if the strut morphology becomes more efficient at the

same relative density. The curvature of radius in the strut morphology affects the predicted compressive

yield strength greatly. The experimental PCM compressive yield strength would improve greatly if the

curvature of radius could be minimized.

The mechanical properties of PCMs are affected by bulk material’s properties but not in a linear

fashion. And the difference in strut thickness and strut morphology between the measured and actual can


73

widen the gap between results. The compression PCM results can be more consistent if an automated or

more precise process in fabricating truss pattern in foam for LFC. Overall, integrated pyramidal sandwich

design improved the limit in producing thin strut without fillability issue.

Chapter 5: Conclusions

74

5. Conclusions

Magnesium pyramidal lattice sandwich structures were successfully fabricated for the first time

using the lost foam casting process. A significant challenge was filling the thin struts of the pyramidal

lattice. The following conclusions regarding PCM fabrication by LFC can be made:

i. Vacuum casting caused metal penetration through the ceramic coating due to high

surface area of the PCM architecture.

ii. A vertical sprue with side feeding design showed the best compromise between fillability

and pattern collapse.

iii. Al or Si alloying addition can be used to aid fillability of the truss core.

iv. The integrated pyramidal sandwich design eliminated the use of hot melt glue during

sandwich panel assembly and minimized casting defects.

Mechanical testing of the as-cast and heat treated Al and Mg alloy micro-trusses showed that the

compression yield strength of the PCM architectures increased with relative density and artificial aging.

The strength of aluminum alloy PCMs increased beyond the yield strength, while magnesium alloy PCMs

underwent a peak stress due to strut fracture.

From the results of this thesis, the following future work is suggested: improve dimensional

accuracy from alternative technique in foam pattern fabrication; minimize the casting defects in filling

PCM patterns by modifying foam material; develop periodic cellular structures other than pyramidal

lattice with lost foam casting.

75

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80

Appendix

A.1. Schematic Diagram

A.2 Phase Diagrams

A.3. Grain Fineness =umber

A.4. Compression Test Results

A.5. Previously Studied PCMs and Metal Foams

Appendix

81

A.1 Schematic Diagrams

Figure A-1. Schematic diagram of corrugation sandwich core with dimensions.

Total Area

θθθθθ sintan

2

tantansin

ccccccc tdhdhdtBase +

−=

−+−= and chHeight =

+

−=

θθ sintan

2 cccc

tdhhAreaTotal

Triangular Open Space

θtan

cc dhBase

−= and cc dhHeight −=

( )θtan

2

cc dhAreaTriangular

−=

Relative Density

( )

+

−

−

−=

θθ

θρ

sintan

2

tan1~

2

cccc

cc

tdhh

dh

Appendix

82

A.2. Phase Diagrams

Figure A-2. Aluminum-silicon phase diagram (left) and magnesium-silicon phase diagram (right) [Nayeb-

Hashemi and Clark, 1988].

Figure A-3. Aluminum-magnesium phase diagram [Nayeb-Hashemi and Clark, 1988].

Appendix

83

A.3. Grain Fineness =umber

The refractory sand particle size can affect the LFC decomposition product permeability. The

average refractory sand size was determined using the American Foundry Society’s grain fineness

number (AFS GFN). From a representative sample, different grit-size mesh sieves were used to separate

different sand particles by size after shaking for 15 minutes. The retained sand weight from each sieve

was used to determine the grain distribution. The final AFS GFN number was calculated using Table 10

and the following formula:

∑∑=

i

ii

f

fMGF' (Eqn. A-1)

Where Mi = Multiplying factor and fi = weight fraction of sand left on sieve.

Table A-1. AFS sieve number and the multiplying factor Mi [after Rao, P.N., 1999].

US Series equivalent

No. (ASTM)

Mesh

opening

(mm)

IS Sieve no.

(µm)

Multiplying

factor

6 3.327 3.35 3

12 1.651 1.70 5

20 0.833 850 10

30 0.589 600 20

40 0.414 425 30

50 0.295 300 40

70 0.208 212 50

100 0.147 150 70

140 0.104 106 100

200 0.074 75 145

270 0.053 53 200

Pan - - 300

Appendix

84

A.4. Compression Test Results

Table A-2. Experimental measured yield strength (σ yield PCM) and peak strength (σ peak PCM) of A356 and AZ91 PCMs in compressions.

Relative

Density

ρρρρ (%)

σσσσ yield PCM σσσσ peak PCM

A356 AZ91 AZ91

As-cast Solutionized Aged As-cast Solutionized Aged As-cast Solutionized Aged

10.5

4.0 2.6 4.5 5.2 5.0 5.2 12.0 12.3 9.4

5.4 5.1 6.4 5.7 4.3 3.4 11.2 9.4 5.5

4.8 4.9 5.1 7.7 5.5 8.1 17.8 13.1 14.1

4.0 6.4 5.2 5.9 12.2 8.0

Average 4.7 4.1 5.6 6.2 5.0 5.7 13.7 11.7 9.3

St. Dev 0.7 1.1 1.0 1.3 0.5 1.9 3.6 1.6 3.6

12.3

6.5 5.4 6.7 8.2 6.7 8.0 16.7 16.3 14.1

6.9 5.8 15.6 8.5 5.8 11.4 20.1 15.0 23.9

7.2 6.3 11.3 7.5 6.4 9.5 16.1 15.5 17.7

5.6 8.1 7.2 7.1 15.3 13.8

Average 6.9 5.8 10.4 8.1 6.5 9.0 17.6 15.5 17.4

St. Dev 0.4 0.4 3.9 0.5 0.6 1.9 2.1 0.5 4.7

14.0

6.2 4.7 12.3 9.1 7.5 9.4 23.9 22.6 19.6

8.0 6.2 8.8 8.6 6.5 10.4 19.9 20.9 19.9

6.2 6.3 8.1 8.9 8.1 10.4 25.0 19.6 24.6

6.3 10.2 5.7 11.3 16.0 17.2

Average 6.8 5.9 9.8 8.9 6.9 10.4 23.0 19.8 20.3

St. Dev 1.0 0.8 1.8 0.3 1.1 0.8 2.7 2.8 3.1

15.7

13.6 9.0 12.9 10.9 9.2 10.0 29.3 23.2 27.0

10.7 8.6 13.6 10.0 8.0 11.2 26.1 23.7 20.6

8.8 7.2 10.4 10.6 9.5 11.0 25.1 24.9 23.3

5.7 14.4 7.2 10.0 24.9 16.6

Average 11.0 7.6 12.9 10.5 8.5 10.5 26.8 24.2 21.8

St. Dev 2.4 1.5 1.7 0.5 1.1 0.7 2.2 0.8 4.4

Appendix

85

Table A-3. Experimental measured yield strength (σ yield) and peak strength (σ peak) of A356 and AZ91 square blocks in compressions.

σσσσ yield σσσσ peak

A356 AZ91 AZ91

As-cast Solutionized Aged As-cast Solutionized Aged As-cast Solutionized Aged

109 128 193 95 91 128 307 291 306

113 136 211 97 105 131 316 293 306

119 142 220 92 110 138 316 299 309

Average 114 135 215 95 108 135 313 296 308

St. Dev 4.9 7.1 6.6 2.6 3.4 5.3 5.0 4.2 1.9

Appendix

86

A.5. Previous studied PCMs and Metal Foams

Table A-4. Previous studies of cast PCMs, summarizing the alloy, architecture, bulk metal density (ρs), relative density ( ρ~ ), PCM’s density

(σ peak PCM), PCM’s specific peak strength (σ peak PCM / ρ PCM), PCM’s yield strength (σ yield PCM) and PCM’s specific yield strength (σ yield PCM / ρ PCM).

Authors Alloy Architecture

ρρρρs ρ~ ρρρρ PCM σσσσ peak PCM σσσσ peak PCM

/ ρρρρ PCM σσσσ yield PCM

σσσσ yield PCM

/ ρρρρ PCM

g/cm3 % g/cm

3 MPa

MPa /

(g/cm3)

MPa MPa /

(g/cm3)

Deshpande and Fleck, 2001 Cu-4Si-1.4Zn Tetrahedral 8.8 8 0.70 10.5 14.9 4.2 6.0

Chiras et al., 2002 Cu-2Be Tetrahedral 8.26 2.25 0.19 4.8 25.8 4.0 21.5

Wang et al., 2003 Cu-2Be Kagome 8.26 2.25 0.19 4.8 25.8 3.8 20.4

Li et al., 2008

Ti-6Al-4V

Pyramidal

4.46 5.4 0.24 22.0 91.3 19.0 78.9

4.46 5.5 0.25 31.0 126.4 29.0 118.2

Ti-6Al-2Sn-

4Zr-2Mo 4.54 5.5 0.25 41.0 164.2 35.0 140.2

Appendix

87

Table A-5. Previous studies of aluminum PCMs, summarizing the alloy, architecture, bulk metal density (ρs), relative density ( ρ~ ), PCM’s density

(σ peak PCM), PCM’s specific peak strength (σ peak PCM / ρ PCM), PCM’s yield strength (σ yield PCM) and PCM’s specific yield strength (σ yield PCM / ρ PCM).


ρρρρs ρ~ ρρρρ PCM σσσσ peak PCM σσσσ peak PCM

/ ρρρρ PCM σσσσ yield PCM

σσσσ yield PCM

/ ρρρρ PCM

g/cm3 % g/cm

3 MPa

MPa /

(g/cm3)

MPa MPa /

(g/cm3)

Deshpande and

Fleck, 2001 Al-7Si-0.3Mg Tetrahedral 2.68 8 0.21 7.3 34.0 4.2 19.6

Wallach and

Gibson, 2001

(443)

Al-4Si-0.2Fe Pyramidal 2.69 6.2 0.17 9.1 53.5 N/A N/A

Zhou et al., 2004

(516.1) Al-3Mg-1Si-1Fe-

1Mn

Pyramidal

2.65 5.3 0.14 5.9 42.0 2.9 20.6

(518.0) Al-8Mg-1.8Fe-

0.35Si 2.57 5.3 0.14 5.6 41.1 3.8 27.9

(A356) Al-7Si-0.4Mg 2.68 5.3 0.14 5.8 40.8 3.5 24.6

Kooistra et al.,

2004 AA6061 Tetrahedral 2.7

8.3 0.22 20.7 92.5 16.2 72.5

5.5 0.15 11.4 76.8 9.3 62.4

3.7 0.10 6.1 61.4 5.7 57.4

3 0.08 4.0 49.6 3.0 37.5

2 0.05 2.8 52.2 2.4 44.3

8.3 0.22

N/A N/A

5.2 23.2

5.5 0.15 2.4 16.2

3.7 0.10 1.1 10.8

3 0.08 1.0 11.7

2 0.05 0.8 14.4

Queheillalt et al.,

2008 AA6061 Pyramidal 2.7 6.2 0.17 11.1 66.1 10.1 60.2

Appendix

88

Table A-6. Previous studies of magnesium metal foams, summarizing the alloy, architecture, bulk metal density (ρs), relative density ( ρ~ ), PCM’s

density (σ peak PCM), PCM’s specific peak strength (σ peak PCM / ρ PCM), PCM’s yield strength (σ yield PCM) and PCM’s specific yield strength (σ yield PCM / ρ PCM).


ρρρρs ρ~ ρρρρ PCM σσσσ peak PCM σσσσ peak PCM /

ρρρρ PCM σσσσ yield PCM

σσσσ yield PCM /

ρρρρ PCM

g/cm3 % g/cm

3 MPa

MPa /

(g/cm3)

MPa MPa /

(g/cm3)

Yamada et al.,

1999

AZ91 Foam

1.81

3.0 0.053 0.132 2.480 0.113 2.111

3.1 0.055 0.135 2.428 0.118 2.127

3.0 0.054 0.105 1.962 0.083 1.540

3.0 0.053 0.083 1.545 0.063 1.186

2.8 0.051 0.069 1.354 0.058 1.140

Yamada et al.,

2000

3.7 0.066 0.143 2.153 0.136 2.053

3.4 0.062 0.124 2.018 0.114 1.851

3.4 0.062 0.106 1.714 0.102 1.649

3.4 0.062 0.112 1.817 0.105 1.703

3.5 0.063 0.081 1.292 0.074 1.175

Korner et al., 2004 65.7 1.19 230.6 193.78 19.8 16.6

56.9 1.03 238.9 231.9 14.0 13.6

Wen et al., 2004 Pure Mg 1.738 65 1.13 12.0 10.6 12.0 10.6

45 0.79 12.0 15.3 10.0 12.8

lost foam casting of periodic cellular materials

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