PROCEDURES FOR PREDICTING PRESSURES INSIDE CORES LEONARD WINARDI

Transcription

1 PROCEDURES FOR PREDICTING PRESSURES INSIDE CORES by LEONARD WINARDI ROBIN D. GRIFFIN, COMMITTEE CHAIR J. BARRY ANDREWS GREGG M. JANOWSKI PETER M. WALSH SRINATH VISWANATHAN HARRY E. LITTLETON A DISSERTATION Submitted to the graduate faculty of The University of Alabama at Birmingham, in partial fulfillment of the requirements for the degree of Doctor of Philosophy BIRMINGHAM, ALABAMA 2007

2 Copyright by Leonard Winardi 2007

3 PROCEDURES FOR PREDICTING PRESSURES INSIDE CORES LEONARD WINARDI MATERIALS ENGINEERING ABSTRACT Core gas defects are among the most aggravating defects because they are difficult to control and may not be found until castings are machined. These defects occur when the pressure in the core is higher than the external pressure acting on the core from the metal-head pressure. The prediction of core and mold gas defects requires a determination of the permeability of the cores and the rate and volume of gas evolved from the cores in contact with molten metal. Techniques were developed for measuring core permeability and gas evolution. Gas permeability was measured at pressure levels that are seen inside cores during casting. The volume and rate of gas evolution from cores submerged in molten metal was also measured. Use of these techniques was demonstrated on commercial cores that were submerged in molten aluminum and iron. The effects of common core-making variables and casting temperatures were determined. Permeability depended mainly on compaction level, with increased density associated with reduced permeability. Coatings decreased permeability, while binder, sand type, and additives had no affect. Gas evolution volumes and rates for cores immersed in molten metal were higher in phenolic urethane cold box cores than in epoxy acrylic cores. Higher binder content, additives, coatings, immersion temperatures, core length, and metal contact area all increased evolved gas volumes and rates. iii

4 A method for calculating the core pressure in simple geometries was developed and confirmed experimentally. The data generated from the gas evolution measurements were used to build a physical model on binder decomposition and the resultant gas evolution during casting. This model was used to determine the amount of gas evolved from cores at various geometries and temperatures. The model accurately predicted the volume of gas evolved. However, the composition of the gases, the core temperature profile, and more precise interfacial heat transfer and sand thermal conductivity data are also required to match the experimental rate curves.. iv

5 DEDICATION I would like to dedicate this book to my wife, Nethaniah T. Winardi, my boy, Tobias Toby Winardi, and my parents, Budi K. Winardi and Shinta S. Winardi for their supports and patience throughout my study. v

6 ACKNOWLEDGEMENTS I thank my advisor, Dr. Robin D. Griffin, for her support during the course of this study, especially in the completion of this thesis. She has been kind enough to accept the most tedious job in the world (correcting and editing my thesis) in the last few months before I graduate. I want to express my gratitude to Mr. Harry E. Littleton, who has been my mentor and friend during this study. I am very grateful to John A. Griffin and Dr. Preston Scarber, Jr., who cared enough to criticize and give insight throughout my study. I am indebted to Professor Charles E. Bates, who convinced me to come to the University of Alabama at Birmingham (UAB) from the Wisconsin tundra. I want to thank the members of the core gas consortium for their support and guidance. The members of core gas consortium are Eck Industries, Cummins Engine, Nissan Automotive, Caterpillar, Inc., HA International, Foseco, Inc., Ashland Casting Solutions, Exone Corp., Deere Co., Kohler, Volvo Power-Train, Scania Trucks, Swecast, Citation Corp., GM-Hormel, and UAB. Finally, I want to thank Jesus Christ, my inspiration, for the invaluable insights and wisdom found in His words. vi

7 TABLE OF CONTENTS Page ABSTRACT... iii DEDICATIONS...v ACKNOLWEDGEMENTS... vi LIST OF TABLES... ix LIST OF FIGURES...x INTRODUCTION...1 Current Methodology for Calculating Core Internal Pressure...4 Analysis of Prior Research...7 Present Study...9 EXPERIMENTAL PROCEDURES...16 Core Density...16 Permeability of Uncoated Cores...16 Permeability of Coated Cores...18 Loss on Ignition...19 Gas Evolution in Contact with Molten Metal...19 Gas Evolution During LOI Determinations...22 Core Internal Pressure Measurements During Immersion...22 NEW TECHNIQUE FOR MEASURING PERMEABILITY OF CORES MADE FROM VARIOUS SANDS, BINDERS, ADDITIVES, AND COATINGS...36 EFFECTS OF COATING DRYING METHODS ON LOI, GAS EVOLUTION, AND CORE PERMEABILITY...69 GAS PRESSURES IN SAND CORES EMPIRICAL MODEL OF GAS EVOLUTION RATES AND VOLUMES FROM SAND CORES vii

8 TABLE OF CONTENTS (Continued) Page CONCLUSIONS Development of Gas Evolution and Permeability Measurement Devices Methodology to Predict Core Internal Pressure Effects of Production Variables on Gas Evolution Effects of Production Variables on Core Permeability Proposed Model to Predict Gas Volatilization and Condensation of Cores in Contact With Molten Metals Future Research GENERAL LIST OF REFERENCES APPENDIX: OBSERVATIONS ON THE EFFECTS OF GAS EVOLUTION RATE ON CORE TEMPERATURE AND BINDER DECOMPOSITION RATE viii

9 LIST OF TABLES Table Page NEW TECHNIQUE FOR MEASURING PERMEABILITY OF CORES MADE FROM VARIOUS SANDS, BINDERS, ADDITIVES, AND COATINGS 1 The composition and formulation of each core system A decrease in volumetric flow rates was observed as a result of an increase in core densities The change in permeability of cores due to coating is tabulated below EFFECTS OF COATING DRYING METHODS ON LOI, GAS EVOLUTION, AND CORE PERMEABILITY 1 Core formulations...86 GAS PRESSURES IN SAND CORES 1 Observations on bubbles escaping from 1.6% PUCB cores immersed in Al-356 at 730C at various depths EMPIRICAL MODEL OF GAS EVOLUTION RATES AND VOLUMES FROM SAND CORES 1 Cores used in this study ix

10 LIST OF FIGURES Figure Page INTRODUCTION 1 Schematic illustrating the formation of gas blow defects The schematic of the mold (dimensions in millimeters) and the locations of blow defects in the casting The viscosity of various gases as a function of temperature...15 EXPERIMENTAL PROCEDURES 1 Schematic apparatus used to measure the air flow through a core Effect of pressure drop on gas flow rate through uncoated and coated AFS GFN sand containing 1.75% epoxy acrylic binder Effect of furnace soaking time on measured LOI (%) Schematic illustration tube furnace used to measure LOI and gas produced in oxidizing conditions Schematic of the displacement apparatus used to determine gas evolution volumes and rates during contact with molten metal Volume of gas collected from water evaporation Sensitivity of the displacement apparatus on various flow rates Responsiveness of the displacement apparatus Gas evolution from an AFS GFN core bonded with 1.75% epoxy acrylic immersed in aluminum at 1350 F Experimental set-up for measuring the core internal pressure by observing the bubble escaping from the cores when immersed into melt at certain depth...33 x

11 LIST OF FIGURES (Continued) Figure Page 11 Experimental set-up for measuring real-time pressure Schematic of the mold. All dimensions are in millimeter...35 NEW TECHNIQUE FOR MEASURING PERMEABILITY OF CORES MADE FROM VARIOUS SANDS, BINDERS, ADDITIVES, AND COATINGS 1 Castings of cored plate mold cast with hollow (1.a) and solid (1.b) shell cores Typical Cornell-Katz plot for sand cores Schematic of apparatus used to measure the air flow through sand cores The volumetric flow rates of gas through cores before and after coating application are shown The volumetric flow rates from uncoated cores and from cores coated with three different coatings (WB-AlSiO4#1, WB-AlSiO4#2, and WB-Mixed) are shown The volumetric flow rates of gas through cores produced with phenolic urethane and two different anti-veining additives (AV-VS and AV-V) are illustrated The volumetric flow rates of gas through cores produced with epoxy acrylic binder with and without AV-M additive are illustrated The sand structure of PUCB cores with AV-M (b), AV-VS (c), AV-V (d) additives and without additive (a) are shown The volumetric flow rates from cores made with 1.6% and 1.9% PUCB and 1.6% and 1.8% epoxy acrylic binders are illustrated The volumetric flow rates from cores produced with three different types of sand (silica: lake sands and in-land sand, mixture of lake sand and chromite, and green sand) are illustrated The volumetric flow rates from 1% PUCB cores produced with 51, 63, and 75GFN in-land sands are illustrated...67 xi

12 LIST OF FIGURES (Continued) Figure Page 12 The volumetric flow rates through cores produced with various sands and air flow temperatures are illustrated...68 EFFECTS OF COATING DRYING METHODS ON LOI, GAS EVOLUTION, AND CORE PERMEABILITY 1 Schematic illustration tube furnace for LOI (%) measurements Effect of furnace soaking time on measured LOI (%) The total gas volume evolved during LOI (%) measurements from cores containing various amounts of binders and volatiles (as determined in LOI test) Effect of drying method on LOI (%) Effect of binder type on LOI (%) Effect of coating viscosity on LOI (%) Interaction between coating viscosity (Baume ) and drying method on LOI (%) Effect of binder type on volume of gas evolved during LOI (%) Effect of drying method on gas volume evolved during LOI (%) Interaction effects between coating viscosity (Baume ) and drying method on gas volume evolved during LOI (%) The first rate peaks from cores containing various amounts of binders and volatiles (as determined in LOI test) The total gas volume evolved at 50 seconds from cores containing various amounts of binders and volatiles (as determined in LOI test) Gas evolution volumes and rates from Cores made with 1.75% PUCB resin, 52GFN S-1 sand, and immersed in iron at 2450 o F Gas evolution volumes and rates from cores made with 1.6% EA resin, 52GFN S-2 sand, and immersed in iron at 2450 o F...99 xii

13 LIST OF FIGURES (Continued) Figure Page 15 Effect of binder type on first peak rate Effect of drying method on first peak rate Interaction effects of coating viscosity (Baume ) and drying methods on first peak rate Effect of binder type on second peak rate Effect of coating viscosity (Baume ) on total gas volume evolved after 50 seconds Air flow rates as a function of density for cores produced from uncoated PUCB and EA resins Coating thickness as a function of binder type, coating viscosity (Baume ), and drying methods Air flow rates as function of binder type, coating viscosity (Baume ), and drying methods Comparison between permeability and gas evolution results GAS PRESSURES IN SAND CORES 1 Experimental set-up for measuring real-time pressure Experimental set-up for verifying core internal pressure A schematic illustrating the dimension and gating system of the mold used for verifying the proposed pressure predictive equation (Equation 3) Gas evolution rate and volume from cores immersed in aluminum at 696 o C (1285 o F) Gas evolution rates and volumes evolved from cores made with 1.6% PUCB Comparison between calculated and measured pressures at the center of the core xiii

14 LIST OF FIGURES (Continued) Figure Page 7 Pressure inside the core calculated from the measured gas evolution rate Core internal pressure calculated from the measured gas evolution rate and the metal-head pressure are illustrated Observations of the timing and occurrence of core gas defects Effect of distance to core print on core internal pressure Effect of core permeability on pressure EMPIRICAL MODEL OF GAS EVOLUTION RATES AND VOLUMES FROM SAND CORES 1 Schematic of the displacement apparatus used to determine gas evolution volumes and rates during contact with molten metal Effects of binder type on gas evolution rate and volume from cylindrical cores made with 1.6% PUCB and epoxy acrylic resins in contact with aluminum at 730 o C (1350 o F) Effects of binder content on gas evolution rate and volume from cylindrical cores made with 1.6% and 0.8% PUCB resin and with and without water-based coating in contact with aluminum at 730 o C (1350 o F) Effects of addition of AV-VS and AV-V additives on gas evolution rate and volume from cylindrical cores made with 1.6% PUCB resin in contact with aluminum at 730 o C (1350 o F) Effects of coating on gas evolution rate and volume from cylindrical cores made with 1.6% epoxy acrylic resin and with and without waterbased coating in contact with aluminum at 730 o C (1350 o F) Effects of core age on gas evolution rates and volumes in PUCB cores Effects of sand grain fineness and density on gas evolution rate and volume Effects of melt type and temperature on gas evolution rate and volume Effects of core length on gas evolution rate and volume from cylindrical cores made with 2.6% PUCB resin in contact with aluminum at 730 o C xiv

15 LIST OF FIGURES (Continued) Figure Page 10 Effects of core length on gas evolution rate and volume from cylindrical cores made with 2.6% PUCB resin in contact with aluminum at 730 o C when both curves were divided by unit area wetted by metal Effects of core modulus on gas evolution rates and gas volume from cylindrical cores made with 2.4% PUNB and modulus of 0.6 and 1.1 cm 3 /cm 2 in contact with aluminum at 680 o C Effects of core modulus on gas evolution rates and total gas volume evolved from cylindrical cores made with 2.4% PUNB and modulus of 0.6 and 1.1 cm3/cm 2 in contact with aluminum at 815 o C Effects of core modulus on gas evolution rates and total gas volume from rectangular cores in contact with aluminum at 680 o C Effects of core shape on gas evolution rates and total gas volume from rectangular and cylindrical cores in contact with aluminum at 815 o C The mold burnt profile (a) and binder content (b) as a function of distance from melt are illustrated Schematic describing the physical model of the gas evolution from resin during contact with molten metal Temperature profile at the center of hotbox core immersed in aluminum at 730 o C Gas evolution rate and volume from 2.1% hotbox cores immersed in aluminum at 730 o C Comparison between calculated and measured gas rate and volume from larger-modulus (1.1 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C Comparison between calculated and measured gas rate and volume from smaller-modulus (0.6 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C Comparison between calculated and measured gas rate and volume from larger-modulus (1.1 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C when condensation was accounted xv

16 LIST OF FIGURES (Continued) Figure Page 22 Comparison between calculated and measured gas rate and volume from smaller-modulus (0.6 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C when condensation was accounted Percentage of the total gases evolved that were being condensed calculated for various core modulus and melt temperature as a function of time a. The calculated volatilization and condensation zones developed in radial direction when 0.6 cm 3 /cm 2 modulus cylindrical cores immersed into aluminum at 680 o C b. Schematic illustrating the location where the temperature and binder content curves were calculated a. The calculated volatilization, condensation, and initial zones developed in axial direction when 0.6 cm 3 /cm 2 modulus cylindrical cores were immersed into aluminum at 680 o C b. Schematic illustrating the location where the temperature and binder content curves were calculated Comparison between calculated and measured gas rate and volume from cores immersed in aluminum at 815 o C Comparison between calculated and measured gas rate and volume from cores immersed in aluminum at 1350 o C Calculated volume of low molecular weight gases evolved from cores with 0.6 and 1.1 cm 3 /cm 2 modulus and immersed into 680 and 815 o C aluminum melt xvi

17 1 INTRODUCTION Surface and sub-surface gas porosity are significant problems in the casting industry. These defects are difficult to control and may not be found until the castings are machined [1-2]. The presence of gas porosity also reduces the ductility of the castings. For example, the presence of gas porosity was found to decrease the percent elongation and ductile-to-brittle transition temperature of 0.26 wt% carbon steel [3]. There are two main sources of gas defects in castings: pinhole and blow defects. Pinhole defects are caused by decreasing gas solubility as the melt temperature decreases and the metal solidifies. These gases separate from the melt and nucleate gas bubbles. The presence of alloying elements can further decrease the solubility of the gases and increase the tendency for pinhole defects to form [4-5]. Blow defects are caused by the bubbling of core gases into the molten metal. Gas bubbles can form when the internal core pressure exceeds the metal-head pressure. The internal pressure of the core increases as soon as molten metal comes into contact with the core, because this causes the binder to volatilize. Figure 1 shows a schematic of the pressure inside the core (dashed line) and two possible metal-head pressure situations versus time. For situation 1, the metal-head pressure (P M,1 ) always exceeds the core pressure, so gas bubbles will not form. For situation 2, the internal core pressure (P C ) exceeds the metal pressure (P M,2 ) up until time t 1 and, during this time, volatile gas can bubble into the molten metal. These bubbles can be trapped in the casting when the

18 2 metal skin forms at the mold surface. No bubble will enter the metal after the metal skin envelopes the core or after the metal-head pressure exceeds the internal core pressure. Most bubbles will escape from the casting as they float up to exit through the mold wall. How the bubbles were trapped in the casting can be revealed from the locations of the porosity after the melt solidifies. There are four possible locations where the bubbles can be trapped, as illustrated in Figure 2 [6-8]. In Figure 2a, the casting and its gating system of a plate mold are illustrated. A cylindrical core made with organically bonded sand was inserted in the middle of the mold. A real-time x-ray system was used to view bubbles from the core during mold filling. The bubbles can be trapped close to the mold wall on the cope side and form subsurface pores, as shown in Figure 2.b. These subsurface pores typically are elongated as the melt compresses the bubbles during filling and frequently occurs in the junction where the metal solidifies at different rates. The bubbles can also form blisters that peel off during shot blasting, as shown in Figure 2.c. The blister is formed when the bubbles prevent the welding of metal skin and the rest of the casting. Bubble trails can form in the middle of the casting when bubbles enter a mushy metal. The bubbles are broken up and entangled as they look for a path to float up and out through the mold wall. An example of a bubble trail is illustrated in Figure 2.d. Finally, the bubbles can indent the metal surface adjacent to the core because they can no longer penetrate to the solidifying metal above the core, although no metal skin has formed at the core surface. A series of indentations adjacent to a core is shown in Figure 2.b. The composition of core gases entering the melt was found to influence the severity of the defects. It has been reported that more reducing gases give a better

19 3 surface finish as less sand adheres to the as-cast surface [9-14]. Furthermore, the presence of hydrogen, nitrogen, and oxygen can be harmful if it exceeds the solubility limit of the solidifying metals [15-19]. The body of work by Bates et al. [9-12] identified that the major gaseous species evolved from cores were H 2, CO, CO 2, and CH 4. Nitrogen, oxygen, and aliphatic hydrocarbon gases were also found. This research effort will focus on controlling the blow defects that originated from a core. The pressure necessary to form a gas bubble can be described using the following expression [20-21]: P M = P C + (2σ/r) Equation 1 where, P M = metal-head pressure (density x metal height x gravitational constant), P C = the core and mold internal pressures, σ = surface tension, and r = radius of the gas bubble. For most situations, the surface tension term is negligible because there are plenty of nucleation sites for the bubble on the mold and core surfaces. Therefore, the pressure needed to form a gas bubble is controlled by the metal-head pressure and core internal pressure. The metal-head pressure can be readily determined if one knows the pouring rate, gating system, melt viscosity, and temperature. However, no method to accurately determine the core internal pressure has been developed [22]. The literature review on methodology for determining the core internal pressure is presented in the following section.

20 4 Current Methodology for Calculating Core Internal Pressure Caine and Toepke [23] reported the effects of the core-metal contact area, the core print area (or the core area uncovered by the melt where the gas can exit), and the gas path length before exiting from the core print on the core internal pressure. The internal core pressure increased proportionally with an increased core-metal contact area and gas path length. The pressure decreased proportionally with an increase in core print area. The influences of core density, permeability, and the percentage of binder decomposed were not evaluated. Worman and Nieman [24] proposed an equation to calculate pressure inside the core based on the results of Caine and Toepke. The equation is described below: P = K( ac a a Pe C P ) Vρ C W N C Equation 2 where, P = pressure (psi), a C = core-metal contact area (cm 2 ), a P = core gas escape area or the total area of vents and prints (cm 2 ), V = total gas volume per gram of sand (cm 3 /g of sand), ρ c = core density (g/cm 3 ), C = percent of core decomposed (%), Pe W-N = permeability (cm 4 /g/min), and K = constant; There are two limitations that reduced the accuracy of this method. First, although the procedure was very useful for comparing gas evolution rates and volumes from different binder systems, the procedure did not imitate the casting conditions, where the core was surrounded by molten metal. The heating rate will be higher for cores surrounded by molten metal, so the gas evolution rate predicted using this procedure is low [25-29]. Second, the pressure of the gas is a function of the number of moles of the

21 5 gas formed and the decomposition temperature. Since the heating rate is different in the tube furnace, the type of gas evolved will be different when cores are heated in tube furnace than in the mold during casting. For example, cores in contact with iron and steel were dominated by hydrogen and carbon monoxide. However, the gas volume was dominated by air and carbon dioxide when similar cores were in contact with aluminum [30-32]. These results suggest that the thermal degradation mechanisms depend on the heating rate. The current permeability measurement devices [33-34] determine the permeability by using Darcy s equation: B AΔP Q = 0 Equation 3 μl where, B 0 = Darcy s constant, A = flow area, ΔP = pressure gradient, μ = gas viscosity, and L = distance the fluid flow. For example, the water column-drum type devices described in AFS Procedure s, s, and s measures the time for a fixed volume of air (1000 cm 3 ) to flow through cores with a fixed dimension and head pressure. These devices are suitable for comparing the permeability of cores; however, they operate in a different flow regime than seen in a casting due to the low pressure 0.69 kpa (0.1 psig) used. As a consequence the flow rate data measured by these devices cannot be used to predict the probability of gas entering the metal. The permeability procedure used by Worman and Nieman combined the permeability contribution of the porosity inside the core with the viscosity of the flowing

22 6 gas. The effects of temperature and gas species on permeability were not determined. However, gas viscosity increases with temperature (Figure 3) and, therefore, would raise the pressure inside the core. For example, the viscosity of hydrogen increases from 10-5 to 1.75x10-5 kg/m-s as the temperature changes from 100 to 600 o C [35]. The molding procedure employed by Worman and Nieman produced low density cores. The core used in their study was made of zircon sand with a density of 2.48 g/cm 3 while the typical density of cores made with zircon sand today ranges from 2.73 to 3.08 g/cm 3 [36]. As a result, the effects of print area and distance to print were very low and could be discounted. Cores made with the current molding technology are denser and harder, while vents are more difficult to insert. Hence, the print area and distance to print becomes important and need to be considered in order to predict local pressure. Campbell [37] proposed another equation to calculate the pressure inside the core. This equation accounted for a number of contributions that Worman and Nieman did not incorporate into their model, including print area, distance to print, the thermal properties of the bonded sand, and the gas evolved. The equation is P. V da ρ L T A T Pe C C P 2 = Equation 4 P 1 C where, P = pressure (Pa), V= volume of gas evolved per second from a kilogram of core material (m 3 /kg/s), d = depth of core heated layer, A C = core-metal contact area, ρ c = core density (kg/m 3 ), L P = distance to core print, T 2 = temperature at which the core gas

23 7 evolution was measured, T 1 = temperature in core at the point of generation, A P = core print area, and Pe C = permeability (m 3 -s/kg). Campbell recognized that the distance the gas traveled and the print area affected the core internal pressure. Campbell determined that higher core thermal conductivity resulted in deeper heat penetration and hence increased the pressure inside the core. This analysis was a significant shift when compared to Worman and Nieman s approach, which assumed 100% decomposition of the core binder. Campbell also postulated that an increase in the temperature of the gas volume would increase the core pressure, which shows the importance of gas properties. Increasing gas temperature would increase both the volume and viscosity, which would increase the pressure, according to Equation 3. Campbell also faced some experimental limitations. The gas evolution rate and permeability data were not collected under real casting conditions. The gas evolution rate data was generated during heating in a furnace, and the permeability was measured under very low pressure conditions, which are not suitable for predicting the pressure in cores during casting. Analysis of Prior Research Improvements are needed in five areas to increase the accuracy of predictions about gas evolution and pressures in molds and cores: 1) Most investigators have measured permeability coefficients with procedures that are not applicable in molds and cores in contact with molten metal. The measurements were made with air at 0.1 psig (0.69 kpa) at room temperature. However, the majority of gases generated from decomposing binder are not air, flow at temperature

24 8 much higher than room temperature and at pressures much higher than 0.1 psig (0.69 kpa). Measurements made at low pressures overestimate the ability of gas to flow through the sand and exit at core prints. Furthermore, the temperature and the type of gas need to be considered in calculating the density and viscosity of the flowing gas. 2) Much of the previous research used total gas volumes produced during pyrolysis to predict pressures that might exceed the metal-head pressure. The pressure developed inside the core increases with increasing gas evolution rate, so the total gas volumes produced need to be presented in terms of rate to be useful. However, the gas volumes cannot be converted directly into rates since the cores were tested in a furnace without contact with molten metal. Previous work presented the gas evolution rate data as a function of specimen weight, which must be converted to calculate the rate as a function of core-metal contact area. The calculation to convert the gas evolution rate data evolved per gram of sand to the rate evolved per surface area requires detailed knowledge of sand and binder chemical compositions. Unfortunately, the chemical compositions of sand and binder are not readily available from the companies that produce them. 3) Procedures should be developed to measure the maximum possible gas evolution volume from cores. Binders may not be pyrolyzed completely during immersion because of the short experiment time, the low sand thermal conductivity, the lack of oxygen, and the re-condensation of the binder. However, binders can be completely boiled off when heated in tube furnace. This procedure has been used to measure the binder (or volatile) content of the sand. This procedure is widely known as the loss on ignition test or LOI test [38]. The maximum volume of gas produced can be measured if all gas evolved during LOI can be collected.

25 9 4) New methodology needs to be developed to determine the pressure by using gas evolution rate data obtained with cores in contact with molten metal. Literature surveys reveal that the pressure developed inside a core during casting increased with an increase in core-metal contact area, distance to core print, and gas evolution rate. The pressure, on the other hand, decreased with an increase in core print area and permeability. No data has been published that relate the pressure with the type of gas evolved and the immersion temperature. Real-time x-ray will be used to visually validate the predicted pressures by observing the formation of gas bubble. 5) Finally, a simulation needs to be constructed to predict the pressures for complex geometries. The rate predicted in simulations can be verified by comparing to the gas evolution rate generated during immersion. Present Study Procedures and methodology for measuring permeability and gas evolution rate need to be improved in order to better predict the pressure inside the core. The goals of this research are as follows: 1. Develop procedures for measuring core permeability. 2. Develop procedures for measuring core coating permeability. 3. Develop procedures for measuring gas evolution rate for cores in contact with molten metal. 4. Develop procedures for measuring the total weight of binder. 5. Develop procedures for measuring the total volume of gas evolved from binder during LOI.

26 10 6. Develop methodology for determining the pressure inside the core by using the data generated from the newly developed procedures. 7. Verify the methodology used to determine the pressure inside the core. 8. Construct a model for predicting the amount of gas evolved from cores made with various moduli and immersed in different melt temperature. The procedures were developed to fulfill objectives 1 through 7. A preliminary model for predicting the gas evolution from cores in contact with molten metal was developed. The practical application of this research is also presented. These procedures, the model, and the practical applications of this research are described in the form of four manuscripts. The first paper describes a new permeability measurement technique [39]. The technique allows the permeability to be measured more accurately, as it took into account changes in gas properties with pressure and temperature. From the measurement, specific numbers for the pore structure of sand cores or coatings were assigned that can be used to predict the permeability of similar cores under different casting conditions. Using the new permeability technique, the effects of temperature, sand type, binder type, additives, coating type, coating thickness, coating specific density, and core density on core permeability were determined. A new UAB core-gas evolution measurement technique is described in the second paper. This technique measures the volume and release rate of gases from a core when it was submerged in molten metal. The results were compared to LOI results to compare the effects of heating rate and abundance of oxygen. Measurements were made on cores dried in microwave and conventional ovens.

27 11 The third paper proposes a methodology for predicting core internal pressure, especially peak pressures and their timing. Verifications were conducted by placing a pressure probe inside experimental cores. Real-time observations on bubble formation, both during immersion and in real-time x-ray, were conducted to further validate the proposed methodology and understand defect occurrence. The effects of core geometry, such as distance to core print, core metal contact area, gas flow area, and core permeability were examined. The effects of gas viscosity were investigated by immersing cores made with two different binder types. The fourth paper describes the preliminary model developed for predicting the gas evolution from cores in contact with molten metal. A model describing the gas evolution of the binder in the core was needed to expand the usefulness of the data generated using the UAB experimental set-up. The UAB set-up measured the gas evolution rate and volume from cores in contact with molten metal made with a specific geometry and immersed at a specific temperature. However, the geometry of the core and contact temperature changes during casting. Furthermore, the experimental results showed that the volume and rate of gas evolved changed depending on the shape, core modulus, and contact temperature. An empirical modeling approach was taken instead of conducting an in-depth study on the degradation mechanism to extrapolate the data generated from the current experimental set-up. This model represented the gas volume data with constants that separate the thermal effects (modulus and melt temperature) and chemical effects (volatilization reactions) for the particular system. With this approach, once the gas evolution volume data is available for a particular core system (binder, core age, drying

28 12 methods, additives), the modeler can predict the gas evolution rate and volume for any core geometry and melt temperature from this core. This data can be combined with the predicted gas composition to determine the type of gas evolved and then to calculate the local core pressure.

29 Metal Head (P M ) or Gas (P C ) Pressure Metal Head Pressure t 1 P M,1 P M,2 Core Gas Pressure P C Time Figure 1. Schematic illustrating the formation of gas blow defects. 13

30 14 Subsurface Porosity Core Indent 2.a [1] 2.b Blister Blister Bubble Trail 2.c 2.d Figure 2. The schematic of the mold (dimensions in millimeters) and the locations of blow defects in the casting. The gas bubbles and the associated bubble trail illustrated in Figure 4.d were found in the middle of the cope side of the casting.

31 Viscosity of Gas Evolved from Cores in Contact with Iron (kg/m/s) 1.0E E E E E E E E E E E E E E E E E E E E E+00 CRC Handbook of Chemistry and Physics, 1984 N Temperature (C) O 2 CH 4 H 2 CO 2 CO Figure 3. The viscosity of various gases as a function of temperature [35]. 15

32 16 EXPERIMENTAL PROCEDURES Core Density Core samples that were 4.45 cm (1.75 inch) in length were cut and placed in a V- block. The ends of each core were ground plane and parallel to produce a specimen with a length of 3.8 cm (1.5 inch). The length and diameter were measured at three locations using a digital caliper with an accuracy of 0.01 cm. These dimensions were used to calculate specimen volumes. Specimen weight was determined using an electronic balance with an accuracy of g. The balance was calibrated against standards each time measurements were made. Density was calculated from the specimen weight divided by its volume. Permeability of Uncoated Cores Gas flow coefficients were measured by placing a 4.45-cm (1.75-inch) long specimen in a shrink-wrap tube, heating the tube to collapse it onto the specimen surface, and then flowing air through the specimen as a function of pressure drop. The pressure drop was measured using electronic differential pressure transducers attached to the ends of the shrink wrap tubing. The flow rate through a specimen was measured using an air flow meter. A schematic of the apparatus is illustrated in Figure 1. The pressure drop and gas flow rates were inserted in Forscheimer s equation [36],

33 17 MPm AΔP RTμLQρ stp = 1 + B 0 A 0 Qρ stp Aμ Equation 3 where, M = molecular weight of air (29), P m = average pressure inside the test sample, A = cross-section area of the test sample, ΔP = the change in pressure through the sand (P u P d, Pa), R = universal gas constant (8310 J/kg.K), T = absolute temperature of air (295 K), Q = volumetric flow rate (cm 3 /cm 2 /s), μ = dynamic viscosity of air ( dyne.s/cm 2 ), L = thickness of the test sample (cm), ρ stp = air density in standard condition ( g/cm 3 ), B 0 = Darcy s constant (cm 2 ), and A 0 = inertial flow coefficient (1/cm). MPm AΔP RTμLQρ stp Qρ stp Aμ A graph of vs. data was made and a linear regression analysis conducted on the linear portion of the plot. The y-intercept of the regression is the reciprocal of Darcy s constant (1/B 0 ), and the slope of the regression is Forscheimer s inertial flow coefficient (A o ). These coefficients were used to calculate gas flow at any pressure, specimen size (length and area), and temperature using Equation 4: A0 MΔP + B 0 B + 0 μ 2RTμL Q = Equation 4 2A0 ρ μ

34 18 Permeability of Coated Cores The same procedures and equipment were used to measure air flow coefficients in sand samples and coated cores. The coating thickness was determined by measuring the change in the core length as the coating was gradually removed from the core end. The coating was considered to have been removed when the pressure drop through the core did not change as coating was removed from the end of the core. This procedure for measuring thickness reflected both the thickness of the wash above the sand surface and the wash that penetrated between the sand grains into the core. The flow coefficients were calculated by determining the air flow rate through a core section that had a coated end, grinding the core wash off, and determining the flow coefficients through the base sand. The local additions were calculated for the sand-pluscoating combination using the procedures described previously. Sample data obtained from a core prepared from AFS GFN sand and bonded with an epoxy acrylic binder is illustrated in Figure 2. Experimental data is plotted with solid symbols, and the line fitted through the data points represents calculations made after the Forscheimer coefficients were determined. This figure illustrates experimental and calculated air flow data through the core as function of pressure drop. The advantage of using Forscheimer coefficients is that they permit flow rates to be calculated for other gases at other pressures and temperatures, so long as the density and viscosity of the gas of interest is known. A coating approximately 0.3-mm-thick reduced the gas flow at all pressures by well over 50%. The pressure range used was higher than is used in the AFS permeability measurements to provide data under pressures experienced in sand molds.

35 19 Loss on Ignition Core samples from each sand mixture were cut to produce disks with a diameter of cm (1.125 inch) and a thickness of approximately cm. This sample had a high surface area to volume ratio to increase the heating rate in a tube furnace. A disc was placed in an Incoel boat, weighed, and placed in the cold end of the tube furnace, and the furnace was sealed. The specimen was pushed into the hot zone at 982 o C (1800 o F) and held for 30 minutes. After 30 minutes, the boat was removed from the furnace, placed in a desiccator, cooled to room temperature, and reweighed. The LOI was calculated as the percent weight loss of the sand. This LOI procedure is similar to the American Foundry Society (AFS) standard S [40], except that a soaking time of 30 minutes instead of two hours was used. The effect of time in the tube furnace on the measured LOI of bonded sand is illustrated in Figure 3. The LOI value did not change significantly with soaking times between 10 and 120 minutes. Because of the consistency in results, a standard soaking time of 30 minutes was used for organically bonded sands throughout this investigation. This procedure does not reproduce the heating rates experienced in a mold, and the tube furnace provides air to react with binder constituents. Binder constituent reactions with air probably produce a greater gas volume than produced in a mold during contact with molten metal. Gas Evolution in Contact with Molten Metal The core immersion device is illustrated in Figure 5. Sand samples were removed from cores of interest, mounted in an insulated holder, and immersed into metal at a

36 20 specified temperature. The immersion temperature was selected depending on the type of metal and core applications. Hot gases formed during binder pyrolysis flowed through a preheated line connecting the specimen holder to a hot oil tank. The line and oil tank were kept hot to minimize the condensation of volatile compounds. As the gas flowed into the oil tank, the gas displaced oil, which flowed from a vent tube into a container located on a precision electronic balance. The weight of displaced oil was measured as a function of time, and the oil weight and density was used to calculate the volume of gas that caused the oil to be displaced. The volume vs. time curve was then differentiated to determine the rate of gas evolution. Temperatures were monitored at three locations in the line connecting the sample to the oil chamber to ensure that all parts of the system were sufficiently hot to prevent moisture condensation. The condensation from one gram of water in the system would reduce the measured gas volume by over one liter. A leak test was performed at the beginning of each experiment by injecting a known volume into the system at room temperature, except for the oil tank. The oil removed from the tank and heated to 115 o C (240 o F) was converted to volume. The volume measured from oil displacement matched the volume from injection to qualify the system to be leak free. A volume curve from evaporated water measured using the displacement apparatus is shown in Figure 6. A tube containing a known weight of water was immersed in liquid oil at 150 o C (312 o F) to check the sensitivity of the device. The measured volume was on average 15% lower than the ideal volume calculated using the

37 21 ideal gas law. The sensitivity of the apparatus would have been better if the tube was heated faster to increase the evaporation rate and prevent re-condensation. The sensitivity of the displacement apparatus in responding to various flow rates is presented in Figure 7. The sensitivity of the device in detecting the increase in volume increased with an increase in flow rate. The sensitivity increased from 90% to 100% as the flow rates increased from 10 to 29 cm 3 /s. The responsiveness of the displacement apparatus to small volumes of air is illustrated in Figure 8. The apparatus required 3 seconds to detect 60 cm 3 of air. The air was injected into a tube furnace used for measuring LOI. The tube furnace was heated to the LOI test temperature of 982 o C (1800 o F). Gas evolution curves from cores made with 1.75% epoxy acrylic resin and immersed in molten aluminum at 732 o C (1350 F) are illustrated in Figure 9. There are two sets of curves on the graph. One set represents the gas volume produced as a function of time, and these curves show a continual rise with time. The volume of gas evolved was shown on the right side of the graph. This core mixture produced a gas volume of about 200 cm 3 within 50 seconds after immersion. The second set of curves illustrates the gas evolution rates. These curves were obtained by differentiating the volume curves with respect to time. The rate curves illustrate the gas produced when the core surface was in contact with the molten metal. These particular curves show an initial gas evolution peak of 26 and 16 cm 3 /s from coated and uncoated epoxy acrylic cores about 3 seconds after immersion. Two additional peaks were observed as the heat penetrated the core and drove out solvents and other volatiles. The first peak usually occurred from 3 to 5 seconds after the sample

38 22 immersion, and the second and third peaks were found about 15 and 35 seconds after immersion with this binder system. Gas Evolution During LOI Determinations The volume of gas evolved during the LOI measurement was determined by connecting the tube furnace to an oil displacement device described in the previous section. The experimental setup excluding the oil displacement device is illustrated in Figure 4. The total gas volume produced is reported in cm 3 /g of specimen and a cm 3 per percent LOI. Core Internal Pressure Measurements During Immersion Three set of experiments were used to verify the accuracy of the proposed methods to predict the core internal pressure. The first set of experiments involved the observation of bubble formation during the immersion of cores at various depths, as shown in Figure 10. Bubbles float from cores when the head pressure is insufficient to suppress the pressure developed in the core during immersion. The results were compared with the head pressure calculated using the proposed equation, and they validated the equation. The second set of experiments was meant to test whether the proposed equation can accurately predict the peak pressures and their timing. In this set of experiments, pressure probes were made from stainless steel tubes having an O.D. of 0.25 cm (0.096 inch). These probes were inserted into cores, connected to a pressure transducer, and the cores immersed in molten metal. Each core had a diameter of 2.86 cm (1.125 inch) and a length of 5.08 cm (2 inch). The end of the pressure probe was placed in the axial center

39 23 of the core and 1.9 cm (0.75 inch) from the wetted core end. The cores were printed into a steel holder that provided a gas escape path and allowed pressure to be measured in real-time using the apparatus schematically illustrated in Figure 10. The cores were immersed 15 cm (6 inch) deep in molten metal. The same apparatus was also connected to a gas evolution measurement device to measure the rate and total volume of gas produced during immersion. The pressure probe tube could also be connected to a vacutainer for collecting gas sample. Finally, real-time x-ray was used to verify the accuracy of the pressure calculated. The real-time x-ray system consisted of a 320 kv x-ray source with a spot size of 0.08 x 0.08-cm and a 22.9-cm (9-inch) tri-field image intensifier. The operation and image collection were computer controlled. Castings with dimensions described in Figure 12.a were poured in the real-time x-ray system.

40 Over-pressure Regulator Input Pressure Transducer Heat-shrink tubing Output Pressure Gauge Air Out Air Inlet Variable Regulator Shut-off Valve Input Connector Specimen Output Connector Flow Meter Figure 1. Schematic apparatus used to measure the air flow through a core. 24

41 500 Air Flow Rate (cc/cm2/s) Measured Air Flow Through Uncoated Core L = 3.8 cm. A = 6.4 cm 2 Measured Air Flow Through Coated Core L = 0.3 mm. A = 6.4 cm 2 Calculated Air Flow Through Uncoated Core L = 3.8 cm. A = 6.4 cm 2 Calculated Air Flow Through Coated Core L = 0.3 mm. A = 6.4 cm Pressure Drop (PSI) Figure 2. Effect of pressure drop on gas flow rate through uncoated and coated AFS GFN sand containing 1.75% epoxy acrylic binder. 25

42 4.0% 3.5% 3.0% 2.5% Sample: , 1.5% Binder Sample Weight: 40 g (about 10 slices) Firing Temperature: 1800F (982C) LOI (%) 2.0% 1.5% 1.0% 1.579% 1.580% 1.582% y = x R 2 = % 0.5% 0.0% Time (minutes) Figure 3. Effect of furnace soaking time on measured LOI (%). 26

43 To GEM system Inconel Tube Heating element Push rod End cap with a seal ring Boat Thermocouple Figure 4. Schematic illustration tube furnace used to measure LOI (%) and gas produced in oxidizing conditions. 27

44 Resistance Heaters Refill Funnel H Beaker Precision Balance Furnace Data Acquisition System Figure 5. Schematic of the displacement apparatus used to determine gas evolution volumes and rates during contact with molten metal. [1] 28

45 Vol Generated (cc) Temp Vol: 1.04gr of water Water Weight (gr) exp T (F) ideal % difference % % % Temperature inside Cu Tube (C) Time (s) Figure 6. Volume of gas collected from water evaporation. 29

46 Air In 40ft Oil Tank Flow Rates (cc/s) Regulator Flow Meter avg flow rates based avg flow rates on Flow meter chart measured from load (cc/s) cell (cc/s) Flow Rates (cc/s) Figure 7. Sensitivity of the displacement apparatus on various flow rates. 30

47 100 60cc Injected into Heated Tube at 1800F via hyperdermic Needle. Vol Displaced (cc) Assuming there was no leak, the air was heated from room temperature up to 215F. Average time to reach 85% of total volume is 3 seconds Time (s) First Injection Second Injection Third Injection Figure 8. Responsiveness of the displacement apparatus. 31

48 Gas Evolution Rate (cm 3 /s) Rate: Epoxy Acrylic Cores Coated WB#3 Fe at 1275C Rate: Epoxy Acrylic Cores Uncoated Fe at 1275C Vol: Epoxy Acrylic Cores Coated WB#3 Fe at 1275C Vol: Epoxy Acrylic Cores Uncoated Fe at 1275C Gas Volume (cm 3 ) Time (s) 0 Figure 9. Gas evolution from an AFS GFN core bonded with 1.75% epoxy acrylic immersed in aluminum at 1275 o C (1350 F). The sample weight was 40 grams. 32

49 Gas Out Depth Al-356 Melt Tubing Core Figure 10. Experimental set-up for measuring the core internal pressure by observing the bubble escaping from the cores when immersed into melt at a certain depth. 33

50 Pressure Probe from Surface 0.75 from Core Al Melt Level: Gas Out Iron Tube D: Steel Holder D: 1.5 L: 2 Core D: L: 1.5 Figure 11. Experimental set-up for measuring real-time pressure. 34

51 Figure 12. Schematic of the mold. All dimensions are in millimeters. [1] 35

52 36 NEW TECHNIQUE FOR MEASURING PERMEABILITY OF CORES MADE FROM VARIOUS SANDS, BINDERS, ADDITIVES, AND COATINGS by LEONARD WINARDI, HARRY E. LITTLETON, AND CHARLES E. BATES Transactions of American Foundry Society, vol. 113, pp (2005) Copyright 2005 by American Foundry Society Used by permission Format adapted and errata corrected for dissertation

53 37 ABSTRACT Gas blow defects remain as one of the major defects in castings despite the extensive research that has been conducted in the past. Although the theoretical concept is well understood, the lack of accurate data for the rate of gas evolution and permeability of sands makes predicting the exact position of the higher pressure locations in the cores and molds difficult. Efforts to develop accurate gas evolution rates have been widely reported in the literature. However, controversy still exists in determining accurate permeability data for core sands. A new technique is presented in this paper for measuring the permeability of core sands. This technique enables foundries to predict the gas flow rates of various gasses through sands for a range of pressures, temperatures, and compaction levels. Moreover, the data can be presented in the form of flow rates as a function of pressure, which is useful in determining locations prone to gas blows if the gas evolution rates, local head pressures, and core density are known. Permeability data of cores produced from various sands, binders, additives, and coatings is presented. Permeability decreased as the core permeability increased. Permeability was reduced significantly by the application of coatings, but none of the coatings tested produced zero permeability. Sands and additives were also found to have a significant affect on permeability. Regular foundry green sand was found to have significantly lower permeability than chemically bonded sand.

54 38 INTRODUCTION During the casting process, heat from the melt causes the volatiles (binders, moisture, or other chemicals) from the molds and cores to evaporate. This gas evolution can produce a localized gas pressure depending upon the permeability of the gas escape paths. If the gas pressure exceeds the local metal pressure, gas blow defects will occur [1]. These defects are shown in x-ray photographs (Figure 1a and 1b) of cored-plate castings. The gating system and the dimensions for these castings are shown in Figure 1c. Some of the uncoated cores used in this study are shown in Figure 1d. The hollow-core-plate casting is free from gas defects, which is attributed to the minimal resistance experienced by the gas flowing out of the core. On the other hand, a solid-core-plate casting contained many gas blow defects. The existence of these defects is attributed to the high resistance (low permeability) of the escape path for the gas evolved. Internal pressure was built up rapidly in the core and was followed by a burst of the gas into the melt after the internal pressure exceeded the local metal pressure. Permeability (Pe) is defined as the ability of gases to flow through a porous media driven by a pressure gradient. The pressure gradient is generated by the metal-head pressure and gas evolution from cores during metal pouring. The Darcy equation [2] has been used in the past to determine the permeability for laminar flow. Darcy defined the flow rate (Q) of a gas with viscosity (μ) through a core material of area (A), length (L), driven by a pressure gradient of ΔP, as described in Equation 1, Q B AΔP = 0 Equation 1 μl

55 39 B 0 is a coefficient describing the shearing phenomenon between adjacent layers during laminar flow and is unique for each porous structure. Forcheimer s equation was found to best predict gas flow rates through typical foundry molds, cores, and coatings [2]. Forcheimer found that the flow rates of gas through a porous media were lower than those predicted by Darcy s equation due to momentum loss from gas particles colliding with the pore walls. In Forcheimer s flow regime, the gas velocities are large enough that they are taken into account. Forcheimer described this flow regime by adding an inertial coefficient (A 0 ) term to Darcy s equation, which represents the non-linear turbulent flow effects. More detailed descriptions on Darcy s and Forcheimer s flow can be found in a Cornell and Katz paper, as stated in the reference section. Forcheimer s equation can be written as MΔP 2RTμLm& = 1 B 0 + A 0 m μ Equation 2 where, M = molecular weight of air, ΔP = the change in pressure over the medium (P u P d, Pa), R = universal gas constant (8310 J/kg.K), T = absolute temperature of the gas (K),. m = mass flow rate of air per unit cross-sectional area (kg/s.m 2 ), μ = dynamic viscosity of air (kg/m.s), L = thickness of the medium sample (m), B 0 = Darcy s constant (m 2 ), and A 0 = inertial flow coefficient (1/m). The current permeability measurement devices [3,4] determine permeability by using Darcy s equation. For example, the water column-drum-type devices described in AFS Procedure s, s, and s specifically measure the time of a fixed

56 40 volume of air (1000 cm 3 ) to flow through cores with a fixed dimension and head pressure. These devices are suitable for comparing the permeability of cores; however, these devices operate in a different flow regime due to the low pressure 0.69 kpa (0.1 psig). As a consequence the flow rate data measured by these devices cannot be used to predict the probability of gas entering the metal. A new permeability measurement technique is proposed in this paper. This technique measures the permeability more accurately, since it takes into account both the properties of a gas flowing at a specific pressure and temperature. The measurement allows one to assign specific numbers for the pore structure of sand cores or coatings and predict the permeability of similar cores at different head pressures, and temperatures, and with different gases or fluids. The measurement results for different cores are presented to evaluate the effects of sands, binders, additives, and coatings on permeability. CALCULATIONS OF SAND PERMEABILITY Since normal gas flow rates through sand and coatings are best described by Forcheimer s equation, the procedures used to determine Darcy s constant and the inertial constant will make use of this expression. For very low permeability sands, the inertial coefficient will not influence the calculated flow rates; however, for high permeability sands, the inertial coefficient will significantly reduce the calculated flow rates. Forcheimer s equation (Equation 2) can be arranged into a linear form to facilitate the determination of Darcy s constant and the inertial coefficient. A plot of

57 41 MΔP m vs. is known as a Cornell-Katz plot. A typical Cornell-Katz plot obtained 2RTμLm μ when measuring the permeability of sand specimens is shown in Figure 2 [5,6]. The plot shows the two regions, which are distinguished by their slopes. At low pressures and low flow rates the slope is steeper. At higher pressures (more than 0.1 psig or 0.69 kpa) and flow rates the slope is less steep and linear. (The pressure developed in typical foundry core is higher than 0.1 psig or 0.69 kpa.) Since the equation was arranged in a linear form, the nonlinear portion of the curve indicates a different flow regime. The linear portion of the curve indicates agreement with the arranged equation and can be used to determine the flow coefficients. The y-intercept of the linear portion is the reciprocal of Darcy s constant (1/B 0 ), and the slope is the inertial flow coefficient (A 0 ). Once these two constants are determined, the air flow through a core as a function of various pressures, temperatures, and flow area and length can be predicted. To solve for the flow rate, equation 1 can be written as, 2 A 1 4A0 MΔP + A B 0 B + 0 μ 2RTμL Q = Equation 3 2A0 ρ μ 1 2 where, Q = volumetric flow rate of air (cm 3 /s), ρ = density of air at mean pressure (g/cm 3 ), and A = cross-sectional area of the coating sample (cm 2 ). Construction of a Cornell-Katz plot is required to determine the flow coefficients, which involves measurements of the flow rates through sand specimens at several

58 42 pressure differentials. The following is a description of the test apparatus and procedures used to obtain these constants, as well as the specimens used in this study. EXPERIMENTAL PROCEDURES Twelve types of core were studied. This section presents the technique for measuring permeability. The specimen preparation and the core composition are also discussed. Permeability Measurement Several techniques were proposed for measuring the gas permeability of sand. A steady state method, which measured the pressure drop across the specimen as a function of volumetric flow rate of air, was adopted. Figure 3 shows the schematic of the experimental set-up. The pressure drop was determined by electronic differential pressure transducers attached to both ends of the specimen. The flow rate (cm 3 /s) was recorded from the flow meter. Care was taken in preparation of the sand specimens in order to produce accurate permeability data. The preparations started with locating the position in the cores that should be tested. A section of the core about 4.5 cm (1.75 inch) was removed and ground to obtain a 3.81-cm (1.5-inch) specimen, flat at both ends. Dimensions and weight (accurate to gram) of the specimens were measured. The specimen was placed between the connectors as shown in Figure 3 and was covered with a heat shrink tube, which is lined inside with polyolefin adhesive. The heat shrink tube was then heated to seal the specimen and connectors. The seal was tested by

59 43 immersing the specimen connector assembly in a water bath and flowing air at 30 cm 3 /s. This air flow helped to detect any leak (creation of bubbles), while preventing water from entering the specimen. The whole system was later wiped and dried to avoid any absorption of water. The pressures were electronically measured by a voltmeter connected to pressure transducers on both the input and output sides of the sand specimen. For each sand specimen, the input and output pressures were recorded for air flow rates ranging from 80 to 1000 cm 3 /s. Specimen Preparation All cores were hand-rammed into a cylindrical-shaped mold having a length of 20 cm (8 inch) and a diameter of 2.9 cm (1-1/8 inch). The cores were hand rammed so that a variation in density existed along the length to evaluate the permeability as a function of density. The permeability was measured at three different locations in the core from each batch. Two different commercial silica sands, lake sand (S-1) and in-land sand (S-2), and a mixture of 70:30 silica (S-1):chromite sands with AFS Grain Fineness Number from 45 to 55 were used in either phenolic urethane cold box (PUCB#1) or epoxy acrylic (EA) bound cores (lake and in-land sands). The AFS Grain Fineness number (AFS GFN) estimates the average sieve size of the sand. The higher the GFN number, the finer the average sand grain size or vice versa [7]. Lake sand was investigated more thoroughly in this project because of its lower expansion rate, which makes it less susceptible to veining defects. Two types of PUCB binder and lake sands were used. Cores made with

60 44 PUCB binder # 2 were made with lake sand #2 with AFS Grain Fineness Number from 50 to 75. Tri-Ethyl-Amine (TEA) and SO 2 gases were used to cure PUCB and epoxy acrylic (EA) cores, respectively. The amount of binder used was determined by the foundry to achieve the specific core mechanical properties. The amount of binder added ranged from 1.5 to 1.9 % of the sand weight. The binder consisted of two parts and the ratio between the two parts varied with binder type. The percentage of part 1 and part 2 by total binder weight in PUCB and EA cores was 55% and 45% and 65% and 35%, respectively. Three types of anti-veining additive: anti-veining (AV) AV-M, AV-VS, and AV- V were investigated. Cores were dipped into water-based (WB) AlSiO 4 #1 and AlSiO 4 #2, and mixed refractory coatings. Regular foundry grade green sand cores were also tested. Composition and formulation for each core batch are tabulated in Table 1. RESULTS AND DISCUSSION The core permeability is compared by the allowable air flow rates through a core at a set pressure gradient. Cores allowing higher flow rates are more permeable. Comparison between calculated and measured air flow rate for cores compacted into various density is shown in Figure 4. The calculated air flow rates match the measured rates very well. The two constants, B 0 and A 0, can therefore be used to calculate the air flow rates in the pressure range applicable to the foundry. The effects of core density, coatings, and additives, type of sands, binders, and binder content are presented in this section. Application of coatings decreases the

61 45 permeability the most by allowing the least amount of air to flow through the coated surface of the core. Core density can be related to the compactibility of the core. Denser cores are compacted better and have smaller pores in between the sand grain. Effects of Core Density on Permeability Permeability is formally defined as the ability of gasses to flow through a porous body. Theoretically, one can also regard permeability as a specific physical property of a core, which describes its pore structure. It has been reported by Adams [4] that permeability varies with the size of the grain, sand grade, amount and nature of the binding material, amount of moisture, and compaction. The relationship between sand grain and grade to permeability is difficult to determine. However, one might expect that the coarser the sand the greater its permeability because of a more open pore structure. The effect of moisture content on core permeability has not been well understood, as in the case of green sand. In green sand, usually the permeability of sand increases with moisture content until a maximum permeability value is reached beyond which it decreases. Other variables being equal, the permeability depends mainly on compaction. The degree of compaction along the core length can be measured easily by the variation in density. Increasing density or compaction of cores produced from the same batch of sand containing the same amount of binder and additive will decrease the permeability as shown in Table 2. For example, in an uncoated core made of lake Sand containing 1.63% phenolic urethane binder, the air flow rates through the cores decreased from 83.8 to 70.7 cm 3 /cm 2 /s as the core density increased from 1.44 to 1.59 g/cm 3 (89.9 to 99 lb/ft 3 ).

62 46 Similarly, uncoated cores produced from a combination of 70% lake sand and 30% chromite sands permitted approximately 30% less air flow rate, from 91.5 to 61.9 cm 3 /cm 2 /s, as the density increased from 1.62 to 2.06 g/cm 3 (101.4 to lb/ft 3 ). The flow rates tabulated in Table 2 were measured at 13.8 kpa (2 psig) to allow comparisons on equal terms. The inertial flow coefficient and Darcy s constant, A 0 and B 0 respectively, were determined by measuring room temperature flow rates of air through cores over several pressure differentials. Effect of Coatings on Permeability Coatings are applied to molds for two main reasons: 1) to eliminate metal penetration and burned-on sand, and 2) to improve the as-cast surface finish. Table 3 shows that coatings can also be used to prevent gas evolved from cores from entering into the liquid metals, since the permeability decreases significantly after the coatings are applied on cores. (However, one should realize that coatings also produce significant gas during casting and hence can make the castings more prone to gas defects.) The normalized air flow rates through cores, produced from 70/30 lake sand/chromite sand and 1.75% epoxy acrylic binder, were significantly reduced after being coated with water-based AlSiO 4 coating. The normalized air flow rates were reduced by more than 70%, from 76.5 to 14.6 cm 3 /cm 2 /s, by the application of 0.19 cm of coating. Figure 4 illustrates the air flow rate reduction at various pressure differentials for cores produced from 70/30 lake sand/chromite sands and 1.75% epoxy acrylic binder before and after they were coated with water-based AlSiO 4 coating. The difference in air flow rate between the uncoated and coated cores becomes larger at a higher pressure

63 47 gradient. The air flow rate through an uncoated and coated core at 13.8 kpa (2 psig) is 91 and 14 cm 3 /cm 2 /s, respectively. However, at 41.4 kpa (6 psig), the air flow rate through uncoated and coated core is 202 and 36 cm 3 /cm 2 /s, respectively. Similar to cores, coating permeability varies depending on its thickness (or geometrical dimensions), packing (or density), and pore structure. All of the variables mentioned are tied to the composition (type and amount of refractory, water- or alcoholbased solution, binders, suspension, and preservative agents) and to the method of coating application (including the drying procedure). These effects of the type and the amount of coating refractories are illustrated in Figure 5. The data plotted in Figure 5 was calculated not only from 30/70 chromite/lake sand cores but also from other cores by using the experimentally determined A 0 and B 0. Cores coated with AlSiO 4 refractory reduced the air flow rates more significantly than those coated with a mixed refractory (Table 3). This is probably due to different pore structures as a result of a variation in the refractory size and shape. Furthermore, water-based AlSiO 4 coatings with 34 o Baume reduced the air flow rates more than those with 32 o Baume due to the increase in the refractory concentration. Effects of Additives on Permeability The presence of a significant amount of sand additive can affect the core permeability. Addition of 5% (based on sand weight, BOS) AV (anti veining)-v additive to cores made with made with 1.6% PUCB and lake sand (S-1) affected the core permeability, as illustrated in Figure 6. Without AV-V additive, the measured air flow rates decreased from 82 to 70 cm 3 /cm 2 /s as the core density increased from 1.44 to 1.59

64 48 g/cm 3 (or 92 to 100 lb/ft 3 ). With AV-V additive, the measured air flow rates decreased from 85 to 54 cm 3 /cm 2 /s as the density increased from 1.51 to 1.6 g/cm 3 (or 95 to 99 lb/ft 3 ). On the other hand, the addition of a small amount (0.05% BOS) of AV-M additive to cores made with 1.6% epoxy acrylic and lake sand (S-1) did not affect the core permeability, as illustrated in Figure 7. For cores made with and without additive, the measured air flow rates decreased from 84 to 60 cm 3 /cm 2 /s as the core density increased from 1.46 to 1.66 g/cm 3 (or 90 to 103 lb/ft 3 ). Beside the amount of additive, the core permeability was also affected by the type of additive. Cores containing AV-V additive had lower permeability than those containing AV-VS additive. Since the composition of these additives was not known, one can only guess the composition of the sand structures shown in Figure 8. Addition of AV-V and AV-VS additives changed the sand color to red and a more metallic color, respectively. The red color in AV-V containing cores might be due to the presence of red iron oxide, which typically had larger grain size and broader grain size distribution than the typical foundry sand [8]. The addition of AV-V additive will alter the compaction process of the core and can potentially decrease the permeability. The black metallic color in AV-VS containing cores might indicate the presence of black iron oxide (Figure 8). Black iron oxide has a grain size very similar to that of foundry sand and may not affect the compaction process of the core [8].

65 49 Effects of Binder Type and Content on Permeability The type of binders used in this investigation did not affect the core permeability when the small amount of binder was mixed into the core. Figure 9 shows that the permeability of cores made with 1.6% PUCB resin was similar to those made with 1.6% epoxy acrylic resin. The measured air flow rates decreased from 85 to 62 cm 3 /cm 2 /s as the core density increased from 1.44 to 1.66 g/cm 3 (or 90 to 102 lb/ft 3 ). The increase in the binder content from 1.6 to 1.9% in PUCB cores did not affect the permeability. However, increasing the binder content from 1.6 to 1.8% in EA cores increased the permeability. The measured air flow rates decreased from 90 to 85 cm 3 /cm 2 /s as the core density increased from 1.44 to 1.46 g/cm 3 (or 95 to 96 lb/ft 3 ). Effects of Sand Type and GFN on Permeability The effects of sand types on permeability are shown in Figure 10. The permeability of cores produced with lake and in-land sands was similar. The measured air flow rates from lake sand cores decreased from 82 to 60 cm 3 /cm 2 /s as the density increased from 1.5 to 1.66 g/cm 3 (or 95 to 103 lb/ft 3 ). The measured air flow rates from in-land sand cores decreased from 88 to 62 cm 3 /cm 2 /s as the density increased from 1.47 to 1.58 g/cm 3 (or 90 to 98 lb/ft 3 ). Although cores produced from a mixture of lake sand and chromite sands are denser than those produced from just silica sands, the air flow rates through both cores fall into the same flow regime (60 to 90 cm 3 /cm 2 /s). The addition of chromite sand did not change the permeability significantly, but it increased the bulk density of the core due to the higher density of chromite sand.

66 50 Regular green sands have the lowest permeability among the sand tested in this study. The measured air flow rates through green sand cores decreased from 30 to 15 cm 3 /cm 2 /s as the density increased from 1.4 to 1.57 g/cm 3 (90 to 97 lb/ft 3 ). The effects of sand grain fineness are illustrated in Figure 11. Cores made with 75GFN lake sand or with the finest sand could be compacted more and therefore had the lowest air flow rates or permeability. On the other hand, cores made with 51GFN lake sand or with the coarsest sand had a lower degree of compactibility and hence, the highest permeability. Effects of Air Flow Temperature The effects of temperatures on air flow rates or permeability of cores are illustrated in Figure 12. The calculated air flow rates decreased from 96 to 83 cm 3 /cm 2 /s as the temperature of the air increased from 25 to 300 o C for cores made with lake sands. For cores made with a combination of lake and chromite sands, the calculated air flow rates decreased from 88 to 69 cm 3 /cm 2 /s as the temperature of the air increased from 25 to 300 o C. For green sand, the calculated air flow rates decreased from 18 to 15 cm 3 /cm 2 /s as the temperature of the air increased from 25 to 300 o C. The decrease in air flow rates or core permeability was mainly due to the increase in air viscosity. The viscosity of air increases from to Pa.s as the temperature increases from 25 to 300 o C [9].

67 51 CONCLUSIONS AND PRACTICAL IMPLICATIONS The technique described in the experimental procedure was used to determine air flow rates through cores produced from different sands, binders, additives, and coatings at room temperature. For each core produced from the same sand composition, there existed a strong relationship between the degree of compaction, indicated by the density and permeability. Higher densities were associated with lower core permeability. When coatings were used, the permeability of cores was significantly reduced (more than 50%). Therefore, coatings can be used to navigate the gas evolved from cores and molds out to the vents. However, one needs to be cautious, since the coating layer is brittle and can be easily chipped off due to mishandling or backpressure created from gas evolved inside the core. Furthermore, the moistures or volatiles left in coatings due to improper drying practices can be introduced into the melt during the casting process. No significant changes were observed in the permeability of cores with the change in binder type and the addition of additives except for the addition of AV-V additives. The addition of 5% (BOS) AV-V additive was found to lower the permeability for denser cores. The addition of 30 wt% of chromite to lake sand (silica) sand changed the relationship between the density and permeability dramatically due to the higher density and different shape of chromite sand. Furthermore, the permeability of cores made from bonded silica sand was higher than that of green sand for the same density range. The permeability was reported as air flow rates for illustrative purposes. However, for research purposes, inertial flow coefficients, Darcy s constants and gas properties are required to estimate the air flow rates at any specific pressures and

68 52 temperatures. These constants describing the pore structure of sand cores should remain unchanged at elevated temperatures since the pore structure will not change due to the low volume fraction of the volatile materials. Further research needs to be conducted to confirm the assumption that the presence of re-condensation products from moisture and binder at elevated temperature do not alter the permeability significantly. In contrast, Bates et al. [2] have reported that the coating pore structures will be more open at higher temperatures due to the high volume fraction of the volatile materials. As a result, these constants describing the coating pore structures will vary depending on temperatures. Unlike the Darcy s and inertial constants, the gas flow rates will be different at elevated temperatures following the change in the viscosity and density of the gases [9]. This technique takes into account these changes and, therefore, can still accurately predict the flow rates. This paper is part of a larger effort to accurately model gas blow defects generated from the gas evolved in cores and molds during casting. Useful permeability data and accurate gas evolution rates are needed along with the surface tension forces of liquid metals (for bubbles to nucleate, break-off, and grow in the melt) to complete this model. Accurate gas evolution rates will be presented in the next publication. A literature review on surface tension for liquid metals is underway. Once these efforts are completed, the model will be verified using real-time x-ray, and guidelines for minimizing core gas defects in castings can be developed for foundries.

69 53 ACKNOWLEDGEMENTS The authors would like to thank Dr. Wanliang Sun and Dr. Preston Scarber, Jr., for their time in discussing about coating permeability and visual data recording. REFERENCES 1. Campbell, C., Castings, 2nd ed., p. 106, Butterworth-Heinemann., Oxford (2000). 2. Bates, C.E., Littleton, H., Miller, B. and Sheldon, D., Lost Foam Process Control for Precision, American Foundry Society Lost Foam Technology and Applications Conference Proceedings, Akron, OH (September, 1995). 3. Bastic, B.A., Blesch, E.V., Nelson, R.M. and Thomas, P.L., Effect of Head Pressure in measuring Green Sand Permeability, AFS Transactions, vol. 99, pp , (1991). 4. Adams, T.C., Testing Molding Sand to Determine Their Permeability, AFS Transactions, vol. 32, pp , (1925). 5. Cornell, D. and Katz, D.L., Flow of Gases Through Consolidated Porous Media, Industrial and Engineering Chemistry Research, vol. 45, pp , (Oct. 1953). 6. Katz, D.L., Cornell, D., Kobayashi, R., Poettman, F.H., Vary, J.A., Elenbass, J.R. and Weinaug, C.F., Handbook of Natural Gas Engineering, McGraw-Hill Book Co. Inc., New York City, New York (1959). 7. American Foundry Society, Mold and Core Test Handbook, 3 rd ed., pp. 1-11, American Foundry Society, Des Plaines, Illinois (2001). 8. Gilson, D.M., Archibald, J.J. and Vandenbos, S., Fighting Veining Defects with Sand Additives, Modern Casting, vol. 85, no. 5, pp , (1985). 9. Weist, R.C., editor, Chemical Rubber Company Handbook of Chemistry and Physics, 65 th ed., CRC Press Inc., Boca Raton, Florida (1984).

70 Table 1. The composition and formulation of each core system. Resin AFS Resin Resin Sand Catalyst Additive Core Wash Type GFN Ratio % Type In-land None 1.5 None Sand (S-2) WB AlSiO4#1 AV-VS None Lake (8% Based on Sand Weight) Phenolic Urethane (PUCB#1) 1.6 Tri-Ethyl-Amine Sand (S-1) /45 AV-V (TEA) (5% Based on Sand Weight) Lake 1.63 None Sand (S-1) None None WB AlSiO4#1 1.9 Lake Sand (S-1) None None PUCB (#2) /45 TEA 1 Lake Sand (S-3) None None None None Lake WB Mixed Refractory 1.6 Sand (S-1) AV-M (0.05% Based on Sand None Weight) WB Mixed Refractory Epoxy /35 SO2 Acrylic 30% Lake None 1.75 Sand/70% None Chromite WB Mixed Refractory 1.8 Lake Sand (S-1) None None 54

71 Table 2. A decrease in volumetric flow rates was observed as a result of an increase in core densities. The flow rates were calculated from the two constants (A 0 and B 0 ) measured for each core system and normalized for constant pressure differential of 13.8 kpa (2 psig) and length of 3.81 cm (1.5 inch). Specimen Description ρ (g/cm 3 ) ρ (lb/ft3) Q (cm 3 /cm 2 /s) Uncoated Lake Sand % PUCB Uncoated In-Land Sand % PUCB Uncoated Lake Sand % PUCB % AV-VS Uncoated Lake Sand % PUCB % AV-V Uncoated Lake Sand % Epoxy Acrylic Uncoated 70/ Lake/Chromite Sand % Epoxy Acrylic Uncoated Lake Sand % Epoxy Acrylic % AV-M Uncoated Green Sand

72 Table 3. The change in permeability of cores due to coating is tabulated below. The flow rates were calculated from A 0 and B 0 measured for each core system and normalized for constant pressure differential of 2 psi and length of 3.81 cm (1.5 inch). Specimen Description Coating Q (cm 3 /cm 2 /s) Uncoated 63 Lake Sand 1.63% PUCB WBAlSiO 4 #1 Coating Thickness: 0.16mm 24.6 Uncoated 74.9 In-Land Sand 1.5% PUCB WBAlSiO 4 #1 Coating Thickness: 0.2mm :30 Lake:Chromite Sand 1.75% Epoxy Acrylic Lake Sand 1.6% Epoxy Acrylic Lake Sand 1.6% Epoxy Acrylic 0.05% AV-M Additive Uncoated 76.5 WBAlSiO 4 #2 Coating Thickness: 0.19mm 14.8 Uncoated 66.9 WB-Mixed Refractory Coating Thickness: 0.41mm 38.4 Uncoated 69.6 WB-Mixed Refractory Coating Thickness: 0.29mm

73 57 Gas Blow Holes 1.a. Hollow Core 1.b. Solid Core Uncoated Coated 1.c. Solid model depicting casting shown above [1] 1.d. Cores made with various additives, sand types, and coatings. Figure 1. Castings of cored plate mold cast with hollow (1.a) and solid (1.b) shell cores. Figure 1.c shows the schematic of the casting including the gating system. All dimensions shown in Figure 1.c are in millimeters. Figure 1.d shows some of the cores used in this study.

74 Cornell-Katz Μ(Pu 2 -Pd 2 )/2RTμLm (1/cm 2 ) 4.E+06 3.E+06 3.E+06 2.E+06 2.E+06 1.E+06 5.E+05 0.E+00 Intercept indicates B 0 Slope indicates A m/μ (1/cm) Figure2. Typical Cornell-Katz plot for sand cores. 58

75 Over-pressure Regulator Input Pressure Transducer Heat-shrink tubing Output Pressure Gauge Air Out Air Inlet Input Connector Specimen Output Connector Variable Regulator Shut-off Valve Flow Meter Figure 3. Schematic of apparatus use to measure the air flow through sand cores. 59

76 250 Pressure Drop (kpa) Air Flow Rate (cm 3 /cm 2 /s) Measured ρ: 1.62g/cm 3 ρ: 1.75g/cm 3 Thickness: 0.17cm Calculated Pressure Drop (PSI) Figure 4. The volumetric flow rates of gas through cores before and after coating application are shown. The flow rates decrease by more than six times at pressure drop of 2 PSI after the core was dipped into WB-AlSiO 4 #2 coating. Values of air flow rates were calculated using Forscheimer s equation. However, past research has shown that omission of the Inertial constant results in less than 1% error for the velocities in typical foundry mold sands. 60

77 100 Density (g/cm 3 ) Uncoated Cores 80 Air Flow Rate (cm 3 /cm 2 /s) WBAlSiO4#1 (32 o ) WBMixedRefractory WBAlSiO4#2 (34 o ) Density (lb/ft 3 ) Figure 5. The volumetric flow rates from uncoated cores and from cores coated with three different coatings (WB-AlSiO 4 #1, WB-AlSiO 4 #2, and WB-Mixed) are shown. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants determined in the lab. The pressure drop is set to 13.8 kpa (2 psig) and thicknesses are set to 3.81 cm (1.5 inch) and 0.3 cm for the uncoated core and each coating layer respectively. All cores were produced from 70:30 lake sand:chromite sand and 1.75% epoxy acrylic binder. 61

78 100 Density (g/cm 3 ) AV-V Air Flow Rate (cm 3 /cm 2 /s) No Additive 1.6% PUCB AV-VS Density (lb/ft 3 ) Figure 6. The volumetric flow rates of gas through cores produced with phenolic urethane and two different anti-veining additives (AV-VS and AV-V) are illustrated. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants determined in the lab. The pressure drop is set to 13.8 kpa (2 psig) and thickness is set to 3.81 cm (1.5 inch). 62

79 100 Density (g/cm 3 ) No Additive 1.6% Epoxy Acrylic Air Flow Rate (cm 3 /cm 2 /s) % Epoxy Acrylic 0.05% AV-M Density (lb/ft 3 ) Figure 7. The volumetric flow rates of gas through cores produced with epoxy acrylic binder with and without AV-M additive are illustrated. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants determined in the lab. The pressure drop is set to 13.8 kpa (2psig) and thickness is set to 3.81 cm (1.5 inch). 63

80 10.a. Without any additive 10.b. With 0.05% AV-M Additive 10.c. With 8% AV-VS Additive 10.d. With 5% AV-V Additive Figure 8. The sand structure of PUCB cores with AV-M (b), AV-VS (c), AV-V (d) additives and without additive (a) are shown. The presence of high amount of additives affects the pore structure of the sand (indicated by the color change) and can alter the core permeability. 64

81 Density (g/cm 3 ) % PUCB 1.8% Epoxy Acrylic % PUCB Air Flow Rate (cm 3 /cm 2 /s) % Epoxy Acrylic Density (lb/ft 3 ) Figure 9. The volumetric flow rates from cores made with 1.6% and 1.9% PUCB and 1.6% and 1.8% epoxy acrylic binders are illustrated. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants as determined in the lab. The pressure drop is set to 13.8kPa (2psig) and thickness is set to 3.81 cm (1.5 inch). 65

82 100 Density (g/cm3) :30 Silica #1:Chromite 80 Air Flow Rates (cc/cm2-s) In-Land Sand Green Sand Lake Sand Density (lb/ft3) Figure 10. The volumetric flow rates from cores produced with three different types of sand (silica: lake sands and in-land sand, mixture of lake sand and chromite, and green sand) are illustrated. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants as determined in the lab. The pressure drop is set to 13.8 kpa (2 psig) and thickness is set to 3.81 cm (1.5 inch). 66

83 100 Density (g/cm 3 ) % PUCB 51GFN 80 Air Flow Rates (cm 3 /cm 2 /s) % PUCB 75GFN 1% PUCB 63GFN Density (lb/ft 3 ) Figure 11. The volumetric flow rates from 1% PUCB cores produced with 51, 63, and 75GFN in-land sands are illustrated. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants as determined in the lab. The pressure drop is set to 13.8 kpa (2 psig) and thickness is set to 3.81 cm (1.5 inch). 67

84 Temperature (C) Lake Sands ρ = 1.57g/cm 3 (98lb/ft 3 ) Air Flow Rates (cm 3 /cm 2 /s) :30 Lake:Chromite Sands ρ = 2.05g/cm 3 (128lb/ft 3 ) Green Sand ρ = 1.55g/cm 3 (97lb/ft 3 ) Temperature (F) Figure 12. The volumetric flow rates through cores produced with various sands and air flow temperatures are illustrated. The data was calculated from the inertial (A 0 ) and Darcy s (B 0 ) constants as determined in the lab. The pressure drop is set to 13.8 kpa (2psig) and thickness is set to 3.81 cm (1.5 inch). 68

85 69 EFFECTS OF COATING DRYING METHODS ON LOI, GAS EVOLUTION, AND CORE PERMEABILITY by LEONARD WINARDI, ROBIN D. GRIFFIN, AND JOHN A. GRIFFIN In preparation for Transactions of American Foundry Society Format adapted for dissertation

86 70 ABSTRACT The effectiveness of the microwave oven in drying industrial water-based coatings when compared to the more conventional warm-air oven was investigated. The gas evolution when cores were in contact with an iron melt and during combustion in loss on ignition furnace was measured. Loss on ignition (LOI) and sand permeability tests were also measured. The gas evolution and LOI data was compared to analyze the possibility of using LOI data for predicting the gas evolution from cores in contact with molten metal. Ten sand mixtures were evaluated to determine the effects of three variables on gas evolution rate and volume during immersion. The cores were made with either phenolic urethane cold box (PUCB) or epoxy acrylic (EA) binders, coated with thin 31 o Baume or thick, 43 o Baume, water-based coating, and dried either in warm air or microwave ovens. (Coating viscosity is typically reported in degree Baume. Higher Baume number indicates higher viscosity.) The LOI values for uncoated PUCB and EA cores were 1.9% and 1.8%, respectively. Most coated cores had lower LOI values than the uncoated cores because the uncoated cores were not dried as in common foundry practice. However, the LOI for 43 o B coated microwave dried cores had higher LOI values than the uncoated cores. This result indicated that there was interaction effect between the drying method and coating viscosity. The volumes and gas evolution rates were measured using the gas evolution measurement device described in a previous publication [1]. A typical gas evolution curve contained two rate peaks. The first peak occurred within 10 seconds after

87 71 immersion and was usually the highest rate, generally between 0.54 and 0.71 cm 3 /cm 2 /s. The second peak usually occurred between 25 and 35 seconds and had values between 0.37 and 0.5 cm 3 /cm 2 /s. EA cores produced a higher first rate peak than PUCB cores. The effect of coating viscosity (Baume ) depended on the drying method, as predicted from the LOI results. The peak rate remained constant for cores dried in a conventional oven. However, the peak gas evolution rate from cores dried in a microwave oven increased with coating viscosity. The time to reach the peak rates was not affected by binder type, coating viscosity, and drying methods. The second peak rate varied with binder type, with PUCB producing higher values than EA cores. The effects of drying method on second peak rate depended on the binder type and viscosity. The total gas volume evolved at 50 seconds after immersion was higher from cores coated with higher viscosity coating. PUCB-bonded cores produced more gas volume than EA cores, except for cores coated with higher viscosity coating and dried in a microwave oven. Coatings significantly reduced surface permeability. Base sands had a standard air flow of about 100 cm 3 /cm 2 /s. EA cores had higher permeability than PUCB cores because of their lower density. The coated cores produced an air flow rate of less than 5 cm 3 /cm 2 /s at 2 psi (13.8kPa). EA-bonded cores had a lower density and higher permeability compared to PUCB bonded cores. The peak gas evolution rate was higher and the peak was reached in less time in EA cores, even though EA cores contained less binder.

88 72 INTRODUCTION Warm-air oven drying is commonly used throughout the foundry industry for drying water-based coatings applied on cores and molds. The drying process itself requires more than an hour to drive out the moisture from the core. The drying time is significantly longer for complex geometry cores with a high degree of air flow restrictions in the pattern. A microwave oven should ideally dry the core faster and more uniformly, which will result in energy- and time-efficient drying. However, there is a limited amount of study in the literature that can quantitatively prove the superiority of the microwave oven over the conventional warm air oven. The effectiveness of microwave oven drying was studied specifically in PUCB and EA binders coated with water-based coating that was applied in two different coating viscosities (indicated in Baume ). The cores were coated and dried using the equipment and procedures of the commercial foundry where this study was requested. An ANOVA analysis was conducted on a full factorial L8 data matrix to reveal the effect of binder type, coating viscosity, and drying method on LOI and gas evolution. MATERIALS AND EXPEIMENTAL PROCEDURES Materials The data matrix was a full factorial L8 and included second-order interactions. Two types of binder, coating drying methods, and viscosities (Baume ) were investigated. Two types of binder were used, phenolic urethane cold box (PUCB) and epoxy acrylic cold box (EA). PUCB and EA cores were made with silica sand having a grain fineness of 52 (AFS standard). All cores were coated with a water-based coating at two Baume

89 73 values, and the cores were dried in either a conventional gas oven or a microwave oven. Cores were dried in a conventional warm air oven having three stages operating at 384 o F, 374 o F, and 120 o F each for 10 minutes. The rest of the cores were dried in a microwave oven for 6 minutes. The uncoated cores were not dried, as is typical for cold box binders. A complete description of the experimental matrix is tabulated in Table I. Each core was produced as a solid cylinder having a diameter of 1 1/8 inches (2.858 cm) and a length of 8 inches (20 cm). Specimens were cut from these cores to make samples to measure density, permeability, Loss on Ignition (LOI), and gas evolution. Core samples were immersed in a gray iron melt at 2450 o F (1370 o C), and the rates and volume of gas produced in contact with liquid metal were determined. Experimental Procedures The density, LOI, and gas evolution during LOI and when the core is in contact with iron melt at 2450 o F (1340 o C) were measured for all core mixtures listed in Table I. The procedures for measuring core density, gas evolution, and sand permeability have been published previously [1,2]. The procedure for measuring the core weight loss during LOI is described later in this section. A t-test and ANOVA F-test analysis were used to analyze the confidence level in both the mean and standard deviations of the data obtained. The analysis presented in the report was performed using Stat Graphics version 15. Core samples from each sand mixture were cut to produce disks with a diameter of inches (3.175 cm) and a thickness of approximately inches (0.318 cm).

90 74 These discs had a high surface area to volume ratio to maximize the heating rate in the tube furnace. A disc was placed in an Inconel boat, the boat and sample weighed, the boat placed in the cold end of a tube furnace, and the furnace sealed. The specimen was then pushed into the hot zone at 1800 o F (982 o C) and held for 30 minutes. The boat was removed from the furnace, placed in a desiccator, cooled to room temperature, and reweighed. The LOI was calculated as the percent weight loss of the sand sample. The schematic describing the LOI furnace set-up is illustrated in Figure 1. This LOI procedure is similar to the American Foundry Society (AFS) standard S, except that a soaking time of 30 minutes was used instead of the 2 hours recommended in the handbook. The effect of time in the tube furnace on the measured LOI of a sand mixture is illustrated in Figure 2. The LOI value did not change significantly with soaking times between 10 and 120 minutes. Because of the consistency in results, a standard soaking time of 30 minutes was used on organically bonded sands throughout this investigation. The disc for determining the LOI of coated core samples was obtained by slicing the coated end of a core so that approximately 50% of the surface was covered by a coating. The coated surface was placed facing upward with the uncoated surface against the Inconel boat. This procedure does not reproduce the heating rates experienced in a mold, and the tube furnace provides air to react with binder constituents. Binder constituent reactions with air probably produce a greater gas volume than produced in a mold during contact with molten metal.

91 75 RESULTS AND DISCUSSION Weight Loss and Gas Volume Evolved From Binder Combustion During LOI The relationship between weight loss and the total volume of gas evolved measured in the tube furnace described in Figure 1 is illustrated in Figure 3. In general, the total volume of gas evolved increased with increasing weight loss. Cores containing 1.8% binders and volatiles evolved 30 cm 3 /g. Cores containing higher binder content at 2.2% evolved 42 cm 3 /g. All three independent variables (binder type, drying method, and coating viscosity) had a statistically significant effect on LOI. The drying technique had the largest effect (largest sum of squares) on LOI. Coated cores dried in a microwave oven had higher LOI values than those dried in the gas-fired oven, as illustrated in Figure 10 bars indicate a 95% confidence interval). The LOI results suggested that the conventional oven drying had volatilized not only the moisture, but also some of the binder in the core. The high LOI value in the conventional oven was expected since the cores were heated up to 374 o F. The higher LOI results in PUCB cores compared to the EA cores would be expected given the higher binder content of the PUCB cores (Figure 5). The effect of coating viscosity (Baume ) was also found where higher viscosity coating (higher Baume values) produced higher LOI values, as illustrated in Figure 6. The coating layer will be thicker in cores coated with higher viscosity coating. The LOI value in cores coated with higher viscosity coating was possibly due to the presence of thicker coating.

92 76 The interaction between coating viscosity and drying method was also significant. The LOI value was not affected by coating viscosity if dried in the gas-fired oven but did increase with coating viscosity if the core was dried by microwave, as illustrated in Figure 7. The increase in LOI was expected because the microwave oven could only volatilize certain components in the coating, such as the moisture and organic components. Other inorganic components in the coating can not be volatilized using the microwave oven. The increase in coating thickness increased the amount of these inorganic components in the coating and, therefore, increased the LOI values of the core. A conventional oven, given enough time and temperature, can volatilize all coating components. The temperature and the core dwelling time employed in this study were enough to volatilize the coating uniformly, and hence, was adequate to be used for drying cores with coating in this viscosity region. Both binder type and drying method had a statistically significant effect on the LOI gas volume. Binder type had the largest effect, with PUCB producing 41 cm 3 /g compared to 32 cm 3 /g for the EA binder (Figure 8). Gas oven drying reduced the gas volume from 38 to 35 cm 3 /g compared to microwave drying (Figure 9). The higher volumes evolved from PUCB and microwave-dried cores resulted from the higher binder and volatile in these cores, as indicated in their higher LOI values. The interaction between coating viscosity and drying method was significant. LOI gas volume was statistically unaffected by drying method with the thinner coating (Figure 10). With thicker coatings, the gas volume was higher if the coating was microwave dried. The same reasoning for explaining the higher LOI values in cores coated with higher viscosity coating and dried in a microwave oven can be applied here.

93 77 Since a microwave oven can only volatilize the moisture and organic components of the coating, the increase in coating thickness in higher viscosity coating resulted in higher inorganic volatile and produced a higher volume of gas. Gas Evolution from Cores in Contact With Molten Iron There was no relationship found between the LOI and the peak gas evolution rate, as illustrated in Figure 11. Cores containing 1.8% (based on sand weight) binders and volatiles had first rate peaks ranging from 0.4 to 0.7 cm 3 /cm 2 /s. Coated cores still contained more volatiles than uncoated cores even after drying. The amount of volatiles are not going to significantly affect the LOI results because of the slow heating experienced by the core heated in the LOI furnace. The higher volatiles content of the coated cores resulted in a increase in the volume of gas evolved. All coated cores produced higher gas evolution rates, as indicated by the color difference in Figure 11. For the same reason, no relationship was found between the LOI and total gas volume produced, as illustrated in Figure 12. Cores containing 1.8% (based on sand weight) binders and volatiles evolved gas ranging from 512 to 770 cm 3 at 50 seconds. Two gas evolution peaks were observed in all cores immersed in molten metal. The first peaks generally occurred within 5 to 7 seconds after contact with molten iron as solvents and organics pyrolyzed from the binder and coating. The second peaks generally occurred after about 30 seconds as volatile material was driven from the center of the core. The first rate peak, which occurred immediately after immersion, was usually higher than the second rate peak. The multiple rate peaks observed were attributed to the volatilization and condensation of the evolved gases as they traveled out

94 78 from the core. The evolved gases from the core surface were condensed and raised the binder content of the next sand layer that had lower temperature. When the sand layer with higher binder content reached the volatilization temperature of the binder, gas evolution rate peaks occurred. The multiple rate peaks were observed as this sequence was repeated. Typical gas evolution rate and volume curves from PUCB and EA cores are shown in Figure 13 and 14, respectively. The gas evolution rate is shown as a function of core-metal contact area (cm 3 /cm 2 /s). Unlike the LOI, coated cores had a higher gas evolution rate and produced higher total gas during immersion, although they had been previously dried. The higher gas evolution rate and volume evolved from coated cores indicated that the drying methods employed were not able to volatilize all of the volatile in the coating. Binder type and drying method had a statistically significant effect on first peak rate. The binder type had the largest effect on first peak rate, with the EA binder producing peak rates of 0.62 cm 3 /cm 2 /s compared to values of 0.58 cm 3 /cm 2 /s from PUCB binders (Figure 15). The higher gas volume evolved from the EA-bonded cores might be due to their higher binder content. The coated cores dried in microwave oven had a lower first peak rate (0.58 cm 3 /cm 2 /s) values than those dried in the conventional oven (0.62 cm 3 /cm 2 /s), as illustrated in Figure 16. This observation contradicted the LOI results that microwavedried cores had higher LOI values and volume. The lower first rate peak suggested that cores dried in microwave oven, although containing higher total volatiles, had smaller

95 79 amounts of moisture or other low-temperature volatiles than those dried in conventional oven. The effectiveness of the microwave oven was slightly affected by the coating viscosity. The first peak rate was higher for higher viscosity (higher Baume ) coating cores dried in a microwave oven (Figure 17). However, the first peak rate remained statistically indifferent for cores dried in a warm air oven regardless, the coating viscosity. The binder type had a statistically significant effect on second peak rate, with PUCB binder producing a higher rate (0.5 cm 3 /cm 2 /s) compared to EA (0.39 cm 3 /cm 2 /s), as illustrated in Figure 18. The higher second rate peaks might indicate that the gases evolved from PUCB cores were more condensable than those from EA cores. The higher condensation rate resulted in higher binder content in the subsequent sand layer and produced higher second rate peaks. Coating viscosity had a statistically significant effect on total volume, with the thinner coating producing 750 cm 3 of gas, compared to 810 cm 3 from the thicker coating, as illustrated in Figure 19. The increase in total gas volume with increasing coating viscosity was expected, since the coating thickness increased with increasing coating viscosity. Permeability Uncoated cores made with EA cores allowed more air flow or were more permeable compared to uncoated PUCB cores, as illustrated by the data in Figure 20. EA cores provided an average air flow of 105 cm 3 /cm 2 /s air flow at 2 psi, while PUCB cores

96 80 provided an average air flow of 90 cm 3 /cm 2 /s. The higher permeability is attributed to the lower density and more open pore structure of the EA cores. The average density of EA cores was 1.55 g/cm 3 (96.7 lb/ft 3 ) compared to a density of PUCB cores of 1.59 g/cm 3 (99.2 lb/ft 3 ). The 2.5 % lower density of the EA bonded cores increased the permeability by 27%, plus it produced less gas, as seen in both immersion and LOI experiments. The coating thickness increased with increasing coating viscosity for PUCBbonded cores. The coating thicknesses for PUCB oven-dried cores coated with 31 o B and 43 o B were 0.03 and 0.04 cm, as illustrated in Figure 21. The coating thickness in EA cores and dried in conventional oven decreased with increasing coating viscosity. The coating thickness appeared to be affected by drying method, with cores dried in the gas oven being thicker that those dried in the microwave oven. This trend was observed except in EA cores dried in microwave oven. The permeability of coated cores depended on coating thickness, as illustrated in Figure 22. Cores coated at 31 o Baume were more permeable than those coated at 43 o Baume. There were no relationship observed between core permeability and gas evolution, as illustrated in Figure 23. The R-Square values for both coated and uncoated cores were less than 0.5. These results showed that the core permeability and density would not affect the gas evolution measurements.

97 81 SUMMARY This investigation was undertaken to determine sand permeability, gas volume, and rate of gas evolution from cores as a function of (1) binder type, (2) coating viscosity (Baume ), and (3) drying method. The data summary is the following: 1. Two gas evolution rate peaks were observed in all cores immersed in molten metal. The first rate peaks generally occurred within 5 to 7 seconds after contact with molten iron as solvents and organics pyrolyzed from the binder and coating. The second rate peaks generally occurred after about 30 seconds as volatile material was driven from the center of the core. The first rate peak, which occurred immediately after immersion, was usually higher than the second rate peak. 2. The drying technique used to dry the cores significantly affected the sand LOI. Coated cores bonded with either EA or PUCB and dried in a microwave oven had consistently higher LOI values than those dried in a warm air oven. 3. LOI values were also influenced by the coating viscosity and binder type. Increasing the coating viscosity increased the LOI. PUCB cores had higher LOI than EA cores due their higher binder content. 4. The total gas volume from core samples increased with LOI. The total gas volume increased from 1650 to 2300 cm 3 /g of weight loss as the LOI increased from 1.7 to 2.08%. 5. The LOI for coated cores dried in microwave and conventional ovens was higher than the LOI for uncoated cores that were not dried. Most coated cores

98 82 produced less gas than uncoated cores during LOI measurements. This is a result of the fact that core drying removes some solvents from the binder. 6. There was no relation between the LOI and the peak gas evolution rate, due to the fact that the peak gas evolution rate almost always occurred immediately after immersion as solvents from coatings and binders flash from the specimen or mold surface. Similarly, no relationship was found between LOI and the total gas volume evolved from cores during immersion. 7. The gas evolution rate and total gas volume evolved were lower for undried and uncoated cores than for coated cores when cores were in contact with molten metal. 8. The first peak rate of gas evolution was affected by binder type and drying method. Cores made with PUCB binder and dried in conventional gas oven produced higher peak rate than those made with EA binder and dried in microwave oven. 9. The effects of coating viscosity on first peak rate depended on the drying method. Cores coated with higher viscosity coating produced higher first peak rate when dried in microwave oven. However, the first peak rate remained unchanged for cores dried in a conventional oven, regardless of the viscosity. 10. The second peak of gas evolution was affected by binder type. PUCB-bonded cores produced higher peak rate than EA-bonded cores. 11. The total gas volume evolved measured at 50 seconds after immersion was affected by the coating viscosity. Cores with higher viscosity produced higher volume.

99 EA-bonded cores were more permeable than PUCB-bonded cores. The average air flow through a 2-inch-long core under a pressure of 2 Psi (13.8 kpa) of 105 cm 3 /cm 2 /s compared to 90 cm 3 /cm 2 /s through PUCB cores. The average density of EA cores was 96.7 lb/ft 3 (1.55 g/cm 3 ) compared to a 99.2 lb/ft 3 (1.59 g/cm 3 ) for PUCB-bonded cores. The 2.5% lower density increased the permeability by 27%, plus it produced less gas in both immersion and LOI experiments. 13. Coating viscosity increased with coating Baume. The coating thicknesses for PUCB-oven-dried cores coated with 31 o B and 43 o B were 0.03 and 0.04 cm. 14. Gas flow under standard conditions decreased with increasing coating thickness. 15. Coated cores dried in a microwave oven were more permeable than those dried in a conventional oven, and coated cores produced with EA resin were more permeable than those produced with PUCB resin. CONCLUSIONS Epoxy acrylic cores had higher first peak rate than PUCB cores when in contact with iron at 1350 o C (2450 o F). Microwave oven drying was more effective in reducing the first peak rate or the maximum gas evolution rate, which occurred early, within 5 to 7 seconds upon metal contact. The effectiveness of microwave oven drying was found to be somewhat dependent on the coating viscosity. The first peak rate was higher for higher viscosity coating dried in microwave oven. Coating viscosity had a statistically significant effect on total volume, with the thinner coating producing 750 cm 3 of gas compared to 810 cm 3 from the thicker coating.

100 84 No relationship was found between LOI and immersion results. The heating rate experience by the immersed core was significantly greater than those put in an LOI furnace. However, the amount of gas produced per gram of sand was higher in LOI than in the immersion. This was due to the abundance of oxygen in the LOI furnace, which effectively combusted most of the binder in the core. The core permeability decreased with increasing core density, the application of coating, and increasing coating thickness. The epoxy acrylic cores were less dense than the PUCB cores and had higher permeability. The application of coating reduced the allowable flow rate in PUCB cores from 90 to 2 cm 3 /cm 2 /s. Both the coating viscosity and the drying methods did not affect the coating thickness and hence the coating surface permeability. The long-term goal of this research is intended to develop objective data that will allow comparisons and evaluations to be made regarding the effects of different binder systems, resins, resin-curing conditions, coatings, additives, and general foundry practices and environments on gas defect formation. Currently, a database has been put together for comparing the effects of type of binders, additives, sands, and coatings on core permeability and gas evolution rate and volume. This paper shows that common foundry practices, such as drying methods, can affect the amount of gas evolved from cores. The current database needs to be expanded to include the influence of general core making practices, such as strip time, humidity, and sand binder bench life, on gas production from cores and molds.

101 85 ACKNOWLEDGEMENTS The authors would like to thank the members of the core gas consortium for guidance and financial support throughout this study. The authors also thank John A. Griffin for his invaluable insights. REFERENCES 1. Scarber, P. Jr., Understandings and Minimizing Gas Defects in Iron Castings, Foundry Management and Technology, vol. 132, pp , (2004). 2. Winardi, L., Littleton, H.E. and Bates, C.E., New Technique for Measuring Permeability of Cores Made From Various Sands, Binders, Additives, and Coatings, AFS Transactions, vol. 113, pp , (2004).

102 Table I. Core formulations Sand Binder Coating Drying Methods Sand AFS Binder % Coating Coating Drying Time Spec# Type Sand # GFN Type Binder Type Name Thick Baume Method Temp (F) (min) Water Rheotec 1-1 Silica S-1 52 PUCB 1.75 Based 464 Thick 43 Oven 384, 374, Water Rheotec 1-2 Silica S-1 52 PUCB 1.75 Based 464 Thick 43 Microwave N/A 6 Water Rheotec 1-3 Silica S-1 52 PUCB 1.75 Based 464 Thin 31 Oven 384, 374, Water Rheotec 1-4 Silica S-1 52 PUCB 1.75 Based 464 Thin 31 Microwave N/A 6 Epoxy Water Rheotec 1-5 Silica S-2 52 Acrylic 1.6 Based 464 Thick 43 Oven 384, 374, Epoxy Water Rheotec 1-6 Silica S-2 52 Acrylic 1.6 Based 464 Thick 43 Microwave N/A 6 Epoxy Water Rheotec 1-7 Silica S-2 52 Acrylic 1.6 Based 464 Thin 31 Oven 384, 374, Epoxy Water Rheotec 1-8 Silica S-2 52 Acrylic 1.6 Based 464 Thin 31 Microwave N/A Silica S-1 52 PUCB 1.75 N/A N/A N/A N/A N/A N/A N/A 1-10 Silica S-2 Epoxy 52 Acrylic 1.6 N/A N/A N/A N/A N/A N/A N/A 86

103 87 To GEM system Inconel Tube Heating element Push rod End cap with a seal ring Boat Figure 1. Schematic illustration tube furnace for LOI (%) measurements. 4.0% 3.5% 3.0% 2.5% Sample: , 1.5% Binder Sample Weight: 40 g (about 10 slices) Firing Temperature: 1800F (982C) LOI (%) 2.0% 1.5% 1.0% 1.579% 1.580% 1.582% y = x R 2 = % 0.5% 0.0% Time (minutes) Figure 2. Effect of furnace soaking time on measured LOI (%).

104 55 Total Gas Volume Evovled from LOI (cm 3 ) LOI (%) Figure 3. The total gas volume evolved during LOI (%) measurements from cores containing various amounts of binders and volatiles (as determined in LOI test). 88

105 LOI (%) Microwave Oven Drying Method Figure 4. Effect of drying method on LOI (%). The error bars indicate a 95% confidence interval. 89

106 LOI (%) Epoxy Acrylic PUCB Binder Type Figure 5. Effect of binder type on LOI (%). The error bars indicate a 95% confidence interval. 90

107 LOI (%) Baume` Figure 6. Effect of coating viscosity degree on LOI (%). The error bars indicate a 95% confidence interval. 91

108 Drying Method Microwave Oven LOI (%) Baume` Figure 7. Interaction between coating viscosity and drying method on LOI (%). The error bars indicate a 95% confidence interval. 92

109 44 LOI Volume (cc/gm) Epoxy Acrylic PUCB Binder Type Figure 8. Effect of binder type on volume of gas evolved during LOI (%). The error bars indicate a 95% confidence interval. 93

110 44 LOI Volume (cc/gm) Microwave Oven Drying Method Figure 9. Effect of drying method on gas volume evolved during LOI (%). The error bars indicate a 95% confidence interval. 94

111 LOI Volume (cc/gm) Drying Method Microwave Oven Baume` Figure 10. Interaction effects between coating viscosity and drying method on gas volume evolved during LOI (%). The error bars indicate a 95% confidence interval. 95

112 Coated and Dried Cores First Peak Rate (cm 3 /cm 2 /s) Uncoated and Undried Cores LOI (%) Figure 11. The first rate peaks from cores containing various amounts of binders and volatiles (as determined in LOI test). 96

113 Coated and Dried Cores Total Gas Volume at 50 sec. (cm 3 ) Uncoated and Undried Cores LOI (%) Figure 12. The total gas volume evolved at 50 seconds from cores containing various amounts of binders and volatiles (as determined in LOI test). 97

114 Gas Evolution Rate (cm 3 /cm 2 /s) Gas Evolution Rate Temperature = 2470 o F ρ = 97.2 lb/ft Gas Evolution Rate Temperature = 2453 o F ρ = 98.5 lb/ft Gas Evolution Rate Temperature = 2501 o F ρ = 98.9 lb/ft Gas Volume Gas Volume Gas Volume Gas Volumes (cm 3 ) Time (s) Figure 13. Gas evolution volumes and rates from cores made with 1.75% PUCB Resin and 52GFN S-1 sand, and immersed in iron at 2450F. 98

115 Gas Evolution Rate (cm 3 /cm 2 /s) Gas Evolution Rate Temperature = 2478 o F ρ = 96.2 lb/ft Gas Evolution Rate Temperature = 2474 o F ρ = 97.4 lb/ft Gas Evolution Rate Temperature = 2476 o F ρ = 97.4 lb/ft Gas Volume Gas Volume Gas Volumes (cm 3 ) Gas Volume Time (s) Figure 14. Gas evolution volumes and rates from cores made with 1.6% EA resin and 52GFN S-2 sand, and immersed in iron at 2450F. 99

116 First Peak Rate (cc/cm2/s) Epoxy Acrylic PUCB Binder Type Figure 15. Effect of binder type on first peak rate. The error bars indicate a 95% confidence interval. 100

117 First Peak Rate (cc/cm2/s) Microwave Oven Drying Method Figure 16. Effect of drying method on first peak rate. The error bars indicate a 95% confidence interval. 101

118 First Peak Rate (cc/cm2/s) Drying Method Microwave Oven Baume` Figure 17. Interaction effects of coating viscosity (Baume ) and drying methods on first peak rate. The error bars indicate a 95% confidence interval. 102

119 Second Peak Rate (cc/cm2/s) Epoxy Acrylic PUCB Binder Type Figure 18. Effect of binder type on second peak rate. The error bars indicate a 95% confidence interval. 103

120 850 Total Volume (cc) Baume` Figure 19. Effect of coating viscosity (Baume ) on total gas volume evolved after 50 seconds. The error bars indicate a 95% confidence interval. 104

121 150 Density (g/cm 3 ) Air Flow Rates (cm 3 /cm 2 /s) Epoxy Acrylic y = x R 2 = PUCB y = x R 2 = Density (lb/ft3) Figure 20. Air flow rates as a function of density for cores produced from uncoated PUCB and EA resins. 105

122 Thickness (cm) PUCB 31B Oven PUCB 31B Microwave PUCB 43B Oven PUCB 43B Microwave E.A 31B Oven E.A 31B Microwave E.A 43B Oven E.A 43B Microwave Figure 21. Coating thickness as a function of binder type, coating viscosity (Baume ), and drying methods. The error bars indicate the standard deviation. 106

123 Air Flow Rates (cm 3 /cm 2 /s) PUCB 31B Oven PUCB 31B Microwave PUCB 43B Oven PUCB 43B Microwave E.A 31B Oven E.A 31B Microwave E.A 43B Oven E.A 43B Microwave Figure 22. Air flow rates as function of binder type, coating viscosity (Baume ), and drying methods. The error bars indicate the standard deviation. 107

124 1000 Gas Volume at 50 sec. (cm 3 ) Coated Cores R 2 = y = x R 2 = Air Flow Rate - Coated Cores (cm 3 /cm 2 /s) Figure 23. Comparison between permeability and gas evolution results. The air flow rate measured during permeability tests at 13.8 kpa (2 PSI) and the total gas volume evolved when cores immersed into iron melt at 1350 o C (2450 o F) for 50 seconds were presented. 108

125 109 GAS PRESSURES IN SAND CORES by LEONARD WINARDI, HARRY E. LITTLETON, AND CHARLES E. BATES Transactions of American Foundry Society, vol. 115, paper no Copyright 2007 by American Foundry Society Used by permission Format adapted and errata corrected for dissertation

126 110 ABSTRACT A method for calculating pressures inside sand cores is presented based on gas evolution rate data collected from cores in contact with molten metal. The technique uses sand permeability measured at pressures up to four psi and the type and concentrations of the pyrolyzed gases. Calculated peak pressures agreed well with the experimentally measured pressures. Verifications were conducted by putting a pressure probe inside experimental cores and by observing bubbles coming from cores immersed to specific depths below the metal surface. Real-time pressure measurements were made to determine core pressures, and both the time and peak pressures were the same for both the measured and calculated values. Calculations predicted that higher metal-head pressures are needed to minimize blow defects in castings produced BY cores having high evolution rates. Higher core pressure increased with higher gas decomposition rate and viscosity, longer distance to core print, larger core metal contact area, lower core permeability, and smaller gas flow area. The gas evolution rate and viscosity increased with immersion temperature. The pressure in the core decreased with an increase in core print area and core permeability. Future research will be concerned with developing and verifying computer codes that predict pressures in commercial cores.

127 111 INTRODUCTION The rapid development in casting technology allows foundries to cast parts with a high degree of intricacy. More intricate castings generally have larger core surface areas wet by molten metal. Venting practices become increasingly more difficult with increased complexity. Increasing core surface area, coupled with difficulties with venting, can significantly raise the pressure inside cores and increases the tendency to produce gas in castings. The gas can produce blows in the casting or the metal may dissolve hydrogen to produce pinholes. Caine and Toepke [1] reported the effects of the core-metal contact area, core print area, and distance to the core print. The internal core pressure increased proportionally with core-metal contact area and gas path length. The pressure also decreased with an increase in core print area. Worman and Nieman [2] proposed an equation for calculating pressure inside cores based on the results of Caine and Toepke. The equation is as follows; P = K ( a C a a C Pe P ) Vρ C W N C Equation 1 where, P = pressure (psi), a C = core-metal contact area (cm 2 ), a P = core gas escape area or the total area of vents and prints (cm 2 ), V = total gas volume per gram of sand (cm 3 /g of sand), ρ c = core density (g/cm 3 ), C = percent of core decomposed (%), Pe W-N = permeability (cm 4 /g/min), and K = constant; This equation was limited by the available data at the time, in particular gas evolution rates, core permeability, and mold density. Gas volumes were measured by

128 112 pyrolyzing samples of sand in a tube furnace at a controlled temperature, and this procedure is now known to produce higher gas volumes than produced when molten metal contacts a core. Although the procedure was very useful for comparing gas evolution rates from different binder systems, it did not imitate the casting conditions, where the core was surrounded by molten metal. The gas evolution rates are lower in tube furnaces because of the lower heating rates, even though the total volume of gas may be greater due to long exposure to heat. The smaller and thinner sample size also contributes to a more complete pyrolysis of the binder when cores are being heated in the tube furnace. The permeability procedure used by Worman and Nieman combined the permeability of the porosity inside cores with the air viscosity. Consequently, the effects of temperature and gas species on permeability were not considered. The molding procedure employed by Worman and Nieman produced relatively low density zircon cores with values of about 2.48 g/cm 3. The normal density of zircon cores today ranges from 2.73 to 3.08 g/cm 3 [3]. As a result, the effects of print area and distance to print were low. Cores made with current molding techniques provide more dense cores. Hence, the print area and distance to the prints are important and must to be accounted for to predict local pressures. Campbell [4] proposed Equation (2), which incorporates the variables missing from Worman and Nieman s equation. The equation is P. V da ρ L T A T Pe C C P 2 = Equation 2 P 1 C

129 113 where, P = pressure (Pa), V= volume of gas evolved per second from a kilogram of core material (m 3 /kg/s), d = depth of core heated layer, A C = core-metal contact area, ρ c = core density (kg/m 3 ), L P = distance to core print, T 2 = temperature at which the core gas evolution was measured, T 1 = temperature in core at the point of generation, A P = core print Area, Pe C = permeability (m 3 -s/kg). Campbell included a) print area, b) distance to core print, and c) preliminary data on the effects of thermal properties of the sand and gas evolved. Sands with higher thermal conductivity would allow heat to penetrate more quickly into the core and develop a higher gas pressure. This approach differed from Worman and Nieman, who assumed 100% binder decomposition. Campbell also postulated that increasing the gas temperature would increase the pressure, and the nature of the gas would have a significant effect on pressure. However, these investigators did not use gas evolution rates and permeability data under the casting conditions as input to the calculations. The permeability was measured under very low pressure conditions, too low to be useful for describing gas flow in cores. Improvements are needed in four areas to increase prediction accuracy of pressures in molds and cores. Most investigators have measured permeability coefficients at significantly lower pressures than those existing in foundry molds. Measurements made at low pressures overestimate the ability of gas to flow through the sand and exit at core prints. Recent research has illustrated that Forcheimer s equation can accurately describes gas flow at pressures between 1.0 and 6.0 psig (6.9 and 41 kpa), which is the range seen in most commercial molds [5]. Forcheimer s equation can be written as

130 114 MΔP 2RTμLm& = 1 B 0 + A 0 m μ Equation 3 where, M = molecular weight of air, ΔP = the change in pressure over the medium (P u P d, Pa), R = universal gas constant (8310 J/kg.K), T = absolute temperature of the gas (K),. m = mass flow rate of air per unit cross-sectional area (kg/s.m 2 ), μ = dynamic viscosity of air (kg/m.s), L = thickness of the medium sample (m), B 0 = Darcy s constant (m 2 ), and A 0 = inertial flow coefficient (1/m). Procedures have been developed to measure the Forcheimer inertial flow coefficient (A o ) and viscous flow coefficient (B o ) in both laboratory cores and in samples removed from commercial molds and cores [5]. Previous research has also used total gas volumes produced during LOI measurements to predict pressures. However, the gas volumes can not be converted directly into rates of evolution because the cores that are rapidly heated undergo different decomposition reactions. The gas volume evolved during LOI measurements usually involves some combustion with oxygen in the system. The rate of gas evolved from cores in contact with molten metal is higher than in LOI measurements, but the volumes are less, as the combustion with oxygen in the system is lower, and the species formed have not been compared. In more modern LOI measurement devices, where the cores are heated in an inert atmosphere, the oxygen is eliminated, but the heating rate is still lower than when cores are in contact with molten metal. A new methodology was needed to determine the pressure produced during binder decomposition in contact with molten metal. The pressure developed inside cores

131 115 is a function of core-metal contact area, distance to core print, and gas evolution rate. The pressure decreases with an increase in core print area and sand permeability. Simulation programs are also needed to predict the pressure in castings having complex core geometries. These simulations must be developed and verified in order to be useful in the casting industry. EFFECTS OF PROCESS VARIABLES ON GAS PRESSURE: LITERATURE REVIEW This section discusses prior research on the effects of permeability, gas evolution rate, and core diameter on the gas pressures inside courts. The effects of area and distance to core print, the type of gas evolved, and melt type and temperatures are discussed. Permeability The effect of permeability on pressure inside cores has been extensively studied [6-8]. Based on data developed by Locke and Ashbrook [6], gas pressure in molds poured with steel decreased exponentially with increases in core permeability. Locke et al. did not actually measure the pressure inside the core, and the core was not completely surrounded by metal; however, their results predicted a pressure decrease with increasing permeability. Naro et al. [7] investigated the tendency toward blowhole formation with changes in the AFS core sand grain fineness number (AFS-GFN), and found that blows increased with grain fineness number. The AFS GFN estimates the average sieve size of the sand. The higher the GFN, the finer the average sand grain size and vice versa [9]. It was also found that that the permeability increased with a decrease in sand grain fineness.

132 116 Levelink et al. [8] measured the pressure inside cores using a probe inserted into the sand and verified that the internal pressure decreased with an increase in permeability. However, it was found difficult to quantify the effect of permeability because of the lack of data on the rate of gas generation and the properties of gas evolved. A new technique was developed by the authors for measuring the permeability of sand with these factors in mind [5]. The technique measured the ability of gas to flow through the pores inside the core and separated the contributions of the type of gas, temperature, core geometry, and porosity on core permeability. Hence, the effects of permeability could be quantitatively measured if the gas species and temperature were known. Gas Evolution Rate Naro et al. [7] investigated the effect of seven binder systems on blowhole formation. Cores bonded with oil produced higher gas evolution rates compared to phenolic urethane cores in LOI measurements when the tube was filled with inert gas. Cores made with core oil also had a higher propensity to produce blowhole defects in the castings compared to phenolic urethane binders. In general, more blowholes were observed in castings when the gas evolution rates were high. Naro et al. also reported that cores made with 0.75% phenolic urethane had a lower tendency to produce blowhole defects compared to those containing 1% or 1.5% phenolic urethane binder. These results indicated that the gas evolution rates or volumes increased with higher binder concentrations.

133 117 There are several methods available for measuring gas evolution rates. Dietert et al. [9] measured the gas evolved from cores in inert atmosphere at a controlled elevated temperature. Levelink et al. [8] noted that the heat transfer to the core in the casting process was much higher than to core samples in a furnace. Hence, gas evolution rates produced using tube furnace procedures would be expected to be lower than in a mold. The volumes and rates of evolution were affected by the binder composition (type and amount of binder), but tube furnace procedures do not consider the effects of core-metal contact area; core modulus, or surface area to volume ratio; distance to core print, and the type of metal poured. The COGAS technique [10] has been used to measure the gas evolution volumes and rates during core immersion in aluminum. The gas evolved passes through a cold trap to remove water and other condensable materials and then displaces water to determine the volume of gas evolved. The resulting data represents air and low molecular weight gases from the cores, but it is difficult to estimate the total gas volume because of the water and other materials condensed in the cold trap. Another procedure developed by the authors is similar to the COGAS technique but uses heated pipes and hot oil displacement to minimize the condensation of water and intermediate weight hydrocarbons in the system. The technique has been described in the literature [11]. Core Diameter The influence of core-metal contact area on pressure was studied by Levelink et al [8]. The pressure was measured inside 30- and 50-mm-diameter cores having similar

134 118 lengths that were immersed in molten iron, and two higher pressure peaks were found in the smaller diameter cores. It was concluded that modulus, or the ratio of volume to surface area, and the contact area of the core with the metal affects the rate of gas evolution and the pressure within the core. Distance to Print, Area of Core Print, and Venting Levelink et al. [8] also measured the internal pressure in cores as a function of immersion depth. Longer cores were immersed deeper into the metal, and the gas produced had to pass further through the sand to exit the core. Consequently, gas pressures increased linearly with immersion depth. Naro et al. [7] reported that core pressure decreased linearly with print area. Enlarging the core print area reduced the tendency for blowhole defects in castings for all binder systems evaluated. It was also found that vents were more effective in reducing pressure than increasing core print area. Caine and Toepke [1] estimated that the presence of small vents could reduce pressure inside the core by a factor of 10. This implies that venting cores may be more effective in preventing gas holes in castings than in increasing the core print area. Effects of Gas Composition and Viscosity The concentrations of several gases evolved from cores prepared with four resins that were placed in contact with molten aluminum, iron, and steel have been published [12]. The data is useful because it provides gas composition as a function of time for about 15 minutes after pouring. The gas effluent from 1.3% PUCB cores in contact with

135 119 iron and steel contained about 45% hydrogen, 30% carbon monoxide, and 15% nitrogen [12]. The rate and type of gas produced was different when the core was put in contact with aluminum. The maximum rates of gas evolved when 1.3% PUCB cores in contact with iron and aluminum were 0.4 and 0.08 cm 3 /cm 2 /s, respectively [12,13]. The gas effluent from 1.3% PUCB cores in contact with aluminum contained about 1.4% hydrogen, 40% carbon dioxide, and 44% nitrogen. The average gas composition was used in the current study to calculate the gas viscosity and the pressure inside cores [12]. The type of gas also determines its ability to flow through the sand. Hydrogen and carbon monoxide, which are the major binder gas decomposition products, have lower viscosities than air and will flow more readily with less pressure drop through the core [14]. The dynamic viscosities of nitrogen and argon at room temperature are 1.77x10-5 and 2.2x10-5 kg/m-s, respectively [14]. Influence of Metal Type and Temperature The gas evolution rate changes with the type of melt poured, primarily because of temperature. The peak gas evolution rate from epoxy acrylic bonded cores immersed in iron at 1340 o C (2450 o F) was 0.9 cm 3 /cm 2 /s, and increased to 1 cm 3 /cm 2 /s when in contact with steel at 1565 o C (2850 o F) [15]. The peak gas evolution rate from phenolic urethane bonded cores immersed in aluminum at 732 o C (1350 o F) was 0.1 cm 3 /cm 2 /s, and the rate from the same cores immersed in iron at 1340 o C (2450 o F) was 0.3 cm 3 /cm 2 /s [15]. These results suggest that the pressures produced inside cores vary with the metal temperature. Future research will deal with gas evolution when cores are placed in

136 120 contact with magnesium and aluminum at the same temperature to determine if the metal might be affecting the reaction products. EXPERIMENTAL PROCEDURES Pressure probes were made from stainless steel tubes having an O.D. of 0.25 cm (0.096 inch). These probes were inserted into cores, connected to a pressure transducer, and the core immersed in molten metal. Each core had a diameter of 2.86 cm (1.125 inch) and a length of 5.08 cm (2.00 inch). The end of the pressure probe was placed in the axial center of the core and 1.91 cm (0.75 inch) from the core end. The cores were printed into a steel holder that provided a gas escape path and allowed pressure to be measured in real-time using the apparatus schematically illustrated in Figure 1. The cores were immersed six inches (15 cm) deep in molten aluminum to produce a metal-head pressure of about 4 kpa (0.6 psi). The same apparatus was also connected to a gas evolution measurement device to measure the rate and total volume of gas produced during immersion. Furan cores were used in these experiments. A second approach was also used to measure the internal pressure in cores. The procedure consisted of immersing cores in molten aluminum to various depths. The metal-head pressure was established by the height of metal above the core. If the gas pressure exceeded the metal-head pressure, gas would bubble through the molten aluminum. The cores were submerged into the molten metal bath with the core axis parallel to the metal surface to minimize pressure gradient, as schematically illustrated in Figure % PUCB cores with 2.86-cm (1.125-inch) diameter by 5.08-cm (2.00-inch) long cores were used in these experiments.

137 121 Aluminum melt at a temperature of 696 o C (1285 o F) was used in all these experiments. Forcheimer s permeability equation was used to calculate the core permeability coefficient because it takes into account the effects of gas viscosity, temperature, and core density on permeability [5]. A third approach was used for verifying the proposed pressure prediction. The core internal pressure was calculated by using the permeability and gas evolution devices described in previous papers [5,13]. Real-time x-ray was used to see when the bubbles started to come off and whether some of the bubbles were trapped in the casting. A schematic showing the dimension of the mold and its gating system is illustrated in Figure 3. One in-gate was completely blocked, and the other in-gate was partially blocked to fill the casting cavity slower, increase the heat exposure time, and decrease the metal-head rise acting on the core. A 1.6% PUCB-coated core with 2.86-cm (1.125-inch) diameter by 20.3-cm (8-inch) long cores and coated with graphite coating was used in these experiments. EXPERIMENTAL RESULTS Triplicate volume and rate curves from a furan resin are illustrated in Figure 4. This data was obtained using the experimental apparatus illustrated in Figure 1, with all gas evolved going into the gas evolution measurements (GEM) device. The initial peak rate was about 0.1 cm 3 /cm 2 /s and occurred about 10 seconds after contact with the metal. The rate increased to about 0.25 cm 3 /cm 2 /s after 40 seconds contact with the metal. The total volume of gas evolved in 60 seconds was about 300 cm 3.

138 122 Triplicate volume and rate curves from a 1.6% PUCB resin are illustrated in Figure 5. There was a rate peak of about 0.1 cm 3 /cm 2 /s about 5 seconds after contact with the metal, and the rate increased slightly with longer contact times. The total volume of gas involved in 60 seconds was about 120 cm³. METHODOLOGY FOR CALCULATING CORE PRESSURE The pressure rises in a core because binder pyrolysis produces gas that cannot freely flow out. Larger gas volumes produced at higher pyrolysis temperatures generally produce higher internal pressures. Forcheimer s equation was reduced to Darcy s equation by ignoring the inertial constant, since past research has shown that this omission will reduce the predicted flow rates by only 1% in the pressure range of typical foundry molds. Darcy s equation provides a basis for describing flow through porous media, as expressed in Equation 4:. Q AC LPμ ΔP = A B P 0 Equation 4 where, ΔP = pressure gradient inside the flowing media or in this case the core (Pa), Q = gas evolution rate (m 3 /m 2 /s), A C = core-metal contact area (m 2 ), L P = distance from one location to core print (m), μ = gas viscosity (kg/m-s), A P = gas flow area (m 2 ), and B 0 = Darcy s constant of the core porosity (m 2 )..

139 123 The pressure in a core increases with higher gas flow rates, higher core-metal contact area, longer distances to core prints, and with higher gas viscosity values. The pressure decreases with higher print areas and sand permeability. The pressure predicted using Equation 3 uses the gas evolution rate data obtained during contact with molten metal and allows the use of gas composition and core geometry to predict internal pressures. The calculated and measured pressure at the center of a furan resin bonded core in contact with aluminum at 696 o C (1285 o F) is shown in Figure 6. The peak pressures agree quite closely to the predicted pressures, especially considering the fact that accurate gas composition data was not available. The calculations predicted the pressure peak timing accurately. The pressures at the first and second peaks were very similar to the measured values. These results also indicate that the pressure curve follows the shape of the gas evolution rate curve closely. It is anticipated that the model will provide more accurate predictions when gas composition and internal core temperatures are included. A second experimental procedure was used to verify the model by immersing cores to various depths in molten aluminum. The gas evolution curves illustrated in Figure 5 were used to make the pressure predictions in the core. The core pressure calculated from the gas evolution rate data plotted in Figure 5 is illustrated in Figure 7. The maximum pressure in the core was calculated to be about 0.67 kpa (0.1 psi), which corresponds to an aluminum head pressure of about 2.5 cm (1 inch). (Metal-head pressure is equal to the metal density times the gravitational constant and the immersion depth.) If cores are immersed to depths of less than one inch, bubbles should be observed coming through the melt. Immersing to higher depths should prevent gas bubbles.

140 124 The theory was tested by immersing cores made with 1.6% PUCB resin into molten aluminum. Bubbles were observed coming from all cores immersed to depths of up to 1.0 inch (2.5 cm), and no bubbles were observed from cores immersed at depths of 3.75 cm (1.25 inch) or greater, as illustrated by the data in Table 1. The calculated head pressure below which bubbles might be observed was 2.5 cm (1 inch). One core out of three experiments bubbled at this metal depth, and no cores immersed to greater depths bubbled. The third experimental procedure was used to test the accuracy of the theory in predicting the timing when the gas starts to bubble from the core. The core internal pressure and metal-head pressure developed during pouring is illustrated in Figure 8. The gas would enter the metal surrounding the core at time less than 1 second. The metalhead pressure would be higher than the core internal pressure approximately 3.5 seconds after the core-metal contact. The prediction matched the real-time x-ray observation illustrated in Figure 9. Bubbles started to form at less than 1 second after the aluminum melt (732 o C) surrounded the core. The first bubble formed at the axial center of the core or at the longest distance to the core print. The core gas bubbling caused turbulent in otherwise very smooth filling of the cavity. The bubbling process stopped after 3.5 seconds or when the metal-head pressure exceeded the core internal pressure. Core gas defects are observed in the casting at directly above the core and close to the riser. Both Worman and Nieman s and Campbell s equations were used to make pressure calculations, in the core. The Worman and Nieman equation was found to overpredict the pressure inside the core. The maximum pressure inside this core (1.6%

141 125 PUCB) calculated using Equation 1 is 43 inches (109 cm) of aluminum. Campbell s equation under predicted the maximum pressure. The maximum pressure was calculated to be 0.15 inch-al. APPLICABLITY OF PRESSURE PREDICTIONS Darcy s equation can be used to predict the effect of core variables on internal pressure. For example, the effect of distance to the core print is illustrated in Figure 10. The pressure increases linearly with distance, and unless metal is poured rapidly to produce higher head pressures or unless the permeability coefficients of the sand are increased, gas can be expected to bubble through the castings. Six inches (15 cm) of aluminum head pressure are required to suppress gas defects from regions located 10 inches (25 cm) from the core print. However, only 1.2 inches (3 cm) of aluminum head pressure is required for a region 2 inches (5 cm) from the print. Increasing core permeability significantly decreases the pressure in the core, as illustrated in Figure 11. There is an asymptotic decrease in internal gas pressure for the geometry being considered with an increase in permeability. An increase in core permeability from 1 to 2 x m 2 decreases the pressure from 2 (5 cm) to 1 (2.5 cm) inches of aluminum. An increase to 4 to 6 x m 2 decreases the pressure by about another 30%. CONCLUSIONS The methodology described increases our ability to predict gas pressure in cores to minimize blows, hydrogen absorption, and bubble trails in castings. The methodology

142 126 must be extended to more complex commercial shapes, but it has been shown to work in two types of laboratory procedures for measuring pressure in cores. This methodology depends on measuring gas evolution rates from cores in contact with molten metal and determining sand permeability coefficients at higher pressures than historically used. This approach requires the viscosities of specific gases in the pyrolysis effluent to be considered. The equation predicts that the pressure increases as the flow rate, flow distance, and fluid viscosity increases. The pressure decreases as the flow area and the permeability of the media increases. Future research and development will focus on modeling gas evolution inside more complex commercial cores to predict internal pressure and the occurrence of gas defects. The type and concentration of the gases evolved during binder a pyrolysis must be known to accurately calculate the pressure inside the core. ACKNOWLEDGMENTS The authors would like to thank Steven Williams, Timothy LeBeau, Pavan Chintalapati, Preston Scarber, Jr., and John Griffin who invested their thought, time, and energy for the completion of this paper. The authors also appreciate the guidance and support from the UAB Core Gas Consortium members. REFERENCES 1. Caine, J.B. and Toepke, R.E., Gas Pressure and Venting of Cores, AFS Transactions, vol. 74, pp 19-22, (1966). 2. Worman, R.A. and J.R. Nieman, A Mathematical System for Exercising Preventive Control over Core gas Defects in Gray Iron Castings, AFS Transactions, vol. 81, pp , (1973).

143 Miller, B., Sheldon, D., Griffin, J., Littleton, H. and Bates, C., Technical Report for Precision Lost Foam Casting Technology, phase 2 (1995). 4. Campbell, J., Castings, 1 st ed., pp , Butterworth-Heinemann, Oxford, UK (1991). 5. Winardi, L., Littleton, H.E., Bates, C.E., New Technique for Measuring Permeability of Cores Made from Various Sands, Binders, Additives, and Coatings, AFS Transactions, vol. 113, pp , (2005). 6. Locke, Charles, and R.L. Ashbrook, Nature of Mold Cavity Gases, AFS Transactions, vol. 58, pp , (1950). 7. Naro, R.L. and Pelfrey, R.L., Gas Evolution of Synthetic Core Binders: Relationship to Casting Blowhole Defects, AFS Transactions, vol. 91, pp , (1983). 8. Levelink, H.G., Julien, F.P.M.A, and De Man, H.C.J., Gas Evolution in Molds and Cores as the Cause of Casting Defects, AFS International Cast Metals Journal, March, pp , (1981). 9. American Foundry Society, Mold and Core Test Handbook, 3 rd ed., pp. 1-11, American Foundry Society, Des Plaines, Illinois (2001). 10. Dietert, H.W., A.L. Graham and R.M. Praski, Gas Evolution in Foundry Materials Its Source and Measurement, AFS Transactions, vol. 84, pp , (1976). 11. MK-GMBH, Description about COGAS, Bates, C.E. and Scott, W.D., The Decomposition of Resin Binders and The Relationship between Gases Formed and The Casting Surface Quality Part 2 Gray Iron, AFS Transactions, vol. 84, pp , (1976). 13. Scarber, P. Jr., Bates, C.E. and Griffin, J., Effects of Mold and Binder Formulations on Gas Evolution When Pouring Aluminum Castings, AFS Transactions, vol. 114, paper no. 130, (2006). 14. Weist, R.C., editor, Chemical Rubber Company Handbook of Chemistry and Physics, 65 th ed., CRC Press Inc., Boca Raton, Florida (1984). 15. Winardi, L. and Bates, C.E., Gas Evolution from Molds and Cores, Presented in AFS Wisconsin Regional, Feb. 9 (2006). 16. Scarber, P. Jr., and Bates, C.E., Simulation of Core Gas Production During Mold Fill, AFS Transactions, vol. 114, paper no. 138 (2006).

144 Table 1. Observations on bubbles escaping from 1.6% PUCB cores immersed in Al-356 at 730C at various depths. Cores immersed horizontally to maintain equivalent head pressure throughout their length. Head Pressure (Immersion Observations Depth) (cm-al) 1.27 cm (0.5 inch) Bubbles observed in 2 out of 3 Immersions 1.59 cm (0.625 inch) Bubbles observed in 3 out of 3 Immersions 1.91 cm (0.75 cm) Bubbles observed in 3 out of 3 Immersions 2.54 cm (1 inch) Bubbles observed in 1 out of 3 Immersions 3.18 cm (1.25 inch) No bubble observed in all 3 Immersions 3.81 cm (1.5 inch) No bubble observed in all 3 Immersions 128

145 Pressure Probe 1.43 cm from Surface Al Melt Level: 15.2 cm Gas Out Iron Tube D: 1.59 cm L: 15.2 cm Steel Holder D: 3.8 cm L: 5.08 cm Core D: 2.86 cm L: 5.08 cm Figure 1. Experimental set-up for measuring real-time pressure. 129

146 To Gas Evolution Measurement Device Al 356 Head Pressure Metal Surface Core Steel Holder Tubing Figure 2. Experimental set-up for verifying core internal pressure. 130

147 Figure 3. A schematic illustrating the dimension and gating system of the mold used for verifying the proposed pressure predictive equation (Equation 3). The core was made with 1.6% PUCB and coated with graphite coating. Aluminum melt was poured into the mold at 732 o C (1350 o F). [1] 131

148 Gas Evolution Rate (cm 3 /cm 2 /s) Rate Core #2 1350F 102 lb/ft 3 Rate Core #4 1350F 102 lb/ft 3 Rate Core #3 1350F 103 lb/ft 3 Vol. Core #2 Vol. Core #4 Vol. Core # Gas Volume (cm 3 ) Time (s) Figure 4. Gas evolution rate and volume from cores immersed in aluminum at 696 o C (1285 o F)

149 Rate (cm 3 /cm 2 /s) Rate Volume (cm 3 ) Volume Time (s) Figure 5. Gas evolution rates and volumes evolved from cores made with 1.6% PUCB. The cores were immersed in Al-356 at 732 o C (1350 o F)

150 Core-Metal Contact Area (m 2 ) Core Print Area (m 2 ) Distance to Core Print (m) μ (kg/m/s) Permeability Coefficient (m 2 ) 4.1E E E E E Pressure (inch-al) Smoothed Measured Pressure Calculated From Gas Evolution Measurements Time (s) Figure 6. Comparison between calculated and measured pressures at the center of the core. 134

151 Head Pressure (cm-al) Core-Metal Contact Area, Ac (m 2 ) 4.1E-03 Core-Flow Area (m 2 ) 6.4E Distance to Core Print (m) 3.8E-02 Gas Viscosity, μ (kg/m/s) 3.8E-05 Permeability, B 0 (m 2 ) 3.0E Time (s) Head Pressure (kpa) Figure 7. Pressure inside the core calculated from the measured gas evolution rate. 135

152 15 12 graphite b coating 1.6% PUCB Ac (m 2 ) 7.48E-03 Ap (m 2 ) 6.41E-04 Lc (m) 7.00E-04 μ (kg/m/s) 1.50E-05 B 0 (m 2 ) 1.17E-13 Metal Head Pressure Metal Head Pressure (in-al) Calc Internal Pressure at the Center Coated Core Time (s) Figure 8. Core internal pressure calculated from the measured gas evolution rate and the metal-head pressure are illustrated. 136

153 0 sec. 1 sec. 2 sec. 3.5 sec Figure 9. Observations of the timing and occurrence of core gas defects. 137

154 Pressure (inch-al) Distance to Core Print, L P (inch) Figure 10. Effect of distance to core print on core internal pressure. 138

155 Pressure (inch-al) Core Permeability, Pe, (x10 11 m 2 ) Figure 11. Effect of core permeability on pressure. 139

156 140 EMPIRICAL MODEL OF GAS EVOLUTION RATES AND VOLUMES FROM SAND CORES by LEONARD WINARDI, HARRY E. LITTLETON, JOHN A. GRIFFIN, AND ROBIN D. GRIFFIN In preparation for Transactions of American Foundry Society Format adapted for dissertation

157 141 ABSTRACT A method has been developed to accurately measure the volume and rate of gas evolution from cores in contact with molten metal. This paper presents data on the effects of common core-making variables on gas evolution. The core-metal contact area and metal temperature largely control the rate of gas evolution. The binder type, additives, and coatings significantly affect the rates and volumes. The core age and core modulus (volume to surface area ratio) also affected both the gas evolution rate and volume. Older cores produced lower gas rates and volumes, while higher metal temperatures increased the rate and volume of gas produced. Higher modulus cores produced more gas, but lower modulus cores produced more gas evolution peaks in the rate of evolution. Core permeability had no effect on the gas volume or rate of gas formation in the range tested, but it did affect the ability of the gas to pass along the core and exit from the core print. An empirical model for calculating the rate and volume of gas evolved for a specific binder system is presented. This model is based on the proposed physical model of gas evolution. The physical model assumes that some of the gases evolved are condensing as they travel out from the core and that the gases are evolving from at least two volatilization reactions, namely the evaporation of moisture and the volatilization of the hydrocarbons in the binder. The model accurately predicted the total gas volume, but improvements are needed to match the experimental rate curves. With better knowledge of the type of gas evolved and the core temperature profile, the model can be used to calculate the gas evolution in production cores and molds.

158 142 INTRODUCTION Developments in molding and core-making technology in the past 15 years allow foundries to cast parts with much higher intricacy and with better dimensional accuracy than ever before. More intricate castings generally have larger core surface areas wet by molten metal during casting. Venting the cores becomes increasingly challenging with increasing casting complexity, and this difficulty increases the local core pressure and the tendency for gas defect formation. Previous research [1] has shown that the pressure developed inside cores increases with larger core-metal contact area, distance to core print, and gas evolution rate. The pressure decreases with an increase in core print area and sand permeability. However, the variables influencing the rate of gas generation during core-metal contact have not been documented. This paper will present gas evolution data from cores having varying lengths, contact areas, and shapes to illustrate the effects of surface area and modulus. The effects of metal temperature, core permeability, and core age on gas evolution rate and volume are also presented. An empirical model is also proposed for predicting the gas evolution rate and volume from cores in contact with various metals. This model will take into account the effects of melt temperature, type of melts, and core modulus. The model is useful for expanding the gas evolution results to various geometries and pouring temperature encountered during casting.

159 143 EXPERIMENTAL PROCEDURES Selected cores made with phenolic urethane cold box (PUCB), phenolic urethane no bake (PUNB), epoxy acrylic (EA), and hotbox resins were made with lengths of 5.1 cm and 8.9 cm. Two types of PUCB resins were tested. Two anti-veining (AV) additives, AV-VS and AV-V, were added to PUCB cores. Some PUCB cores were also coated with water-based coating. Core diameters were 2.9 cm and 6.1 cm. In addition, 3.8-cm-long rectangular cores with edge lengths of 2.5 cm and 4.1 cm were also made. Cores were immersed into aluminum at temperatures of 680, 700, 730, or 815 o F. These cores were also immersed in iron at 1350 o C. Each core was printed into a steel holder to a depth of 1.3 cm, resulting in 3.8 cm or 7.6 cm of the core in contact with molten metal. The core formulations, dimensions, sand grain fineness (GFN) determined by the methods published by American Foundry Society [2], and permeability used in this study are tabulated in Table 1. Each formulation was tested three times. A single representative data set is plotted in each case. The device used to immerse cores and measure the volume of gas produced is illustrated in Figure 1. When the core contacted the metal, the high-temperature pyrolysis gases flowed through a preheated line connecting the specimen holder to a hot oil tank. The line and oil tank were kept hot to minimize condensation of volatile compounds. As the gas flowed into the oil tank, oil was displaced, and the oil overflowed the tank through a vent tube into a container located on a precision electronic balance. The weight of displaced oil was measured as a function of time, and the oil weight and density were used to calculate the gas volume that caused the oil to be displaced.

160 144 Temperatures were monitored at three locations in the line connecting the sample to the oil chamber to assure that all parts of the system were hot enough to prevent moisture condensation. The precipitation of one gram of water in the system would reduce the measured gas volume by over one liter. The temperature profile of the core at various distances from the melt-core interface was measured using an open bead (0.159-cm or 1/16- inch diameter) type K thermocouple. The thermocouples were inserted into holes drilled into the core. The thermocouples were mechanically fitted, and no sodium silicate or core paste was used. The authors recognized that although the temperature profiles recorded might be lagging due to a lack of sand-bead contact, the thermal events, such as condensation and evaporation of volatiles, could still be monitored. The binder burn profile was also investigated to develop a physical model of binder decomposition during contact with molten metal. A mold made with 4% phenolic urethane no-bake was used to cast cylindrically shaped iron with diameter of 5.08 cm (2 inch) and length of 12.7 cm (5 inch). The iron was poured into the mold at 1450 o C (2650 o F). The burn profile was observed and photographed. The binder composition of each layer of the burn profile was then determined by measuring the weight loss of the sand after being heated at 982 o C (1800 o F) for 30 minutes. EXPERIMENTAL RESULTS AND DISCUSSION The effects of core length, metal contact area, modulus (volume to surface area), and core shape are discussed in the following sections. The effects of core age,

161 145 permeability, and metal type and contact temperatures are also presented. These results were used to verify the empirical model presented later in the paper. Effects of Binder Type, Binder Content, Additives, and Coatings The type, amount, and brand of binder affected the gas evolution from cores in contact with aluminum at 730 o C (1350 o F). Cores containing similar amount of PUCB #1 and EA resins produced significantly different gas evolution rate and volume, as illustrated in Figure 2. The maximum gas evolution rate produced from 1.6% PUCB and 1.6% EA cores was 3 and 6 cm 3 /s, respectively. The corresponding total gas volume produced at 60 seconds was 140 and 160 cm 3. Increasing the binder content from 0.8% to 1.6% increased the maximum gas evolution rate from 3 to 5 cm 3 /s in cores made with 1.6% PUCB #2 (Figure 3). Similarly, the total gas volume increased from 120 to 225 cm 3. Comparison on the gas evolution from PUCB # 1 (Figure 2) and # 2 (Figure 3) shows that the variation existed not only in the type and amount of binder added, but also in the brand of the binder (different manufacturer of the binder). Addition of additives and application of coating significantly increased the gas evolution rate and volume from PUCB cores. The gas evolution rate and volume doubled from 3 to 7.5 cm 3 /s and from 140 to 300 cm 3 when AV-VS and AV-V additives were added into 1.6% PUCB cores (Figure 4). Cores washed with a water-based coating had higher first rate peak and produced higher gas volumes. The first rate peak was tripled from 3 to 10.5 cm 3 /s, and the total gas volume was doubled from 160 to 320 cm 3 when the core was coated (Figure 5).

162 146 Effects of Core Age Core age can significantly affect the amount of gas evolved, as illustrated in Figure 6. PUCB cores placed in contact with aluminum immediately after blowing produced higher rates and volumes. The peak rates obtained after one day and after 10 days were 35 and 6 cm 3 /s. The gas volumes were about 300 and 200 cm 3 after the same time lapse. The lower rates of gas evolution associated with aged cores was probably associated with some solvent evaporation and with greater cross-linking of the resin. Although each core was aged in a sealed bag, some volatiles might still have been lost with time. (No statistical data on the weight variation between the un-aged and aged cores was obtained.) Binders with a higher degree of cross linking are also thought to need more energy to pyrolyze and this might reduce the volume of gas produced. More information is needed to minimize speculation, and such an investigation should involve determining the gases evolved as a function of core age. Effects of Sand Grain Fineness and Density (or Permeability) Cores made with finer sand are easier to compact and usually denser. The increase in core density will result in a decrease in core permeability. Therefore, the permeability generally decreases with a decrease in sand grain size [3]. The average sand grain size can be determined using the AFS screen method. In this method, higher grain sand fineness number (GFN) means finer sand [2]. The variations in sand grain fineness and density in the range of this study did not affect the volume or rate of gas evolution, as illustrated in Figure 7. The gas evolution rate and total volume evolved from PUCB cores made with 50 and 75 and GFN were

163 147 similar at 34 cm 3 / s and 290 cm 3, respectively. Although sand grain fineness and density do not affect the amount of gas evolved, it is still an important parameter because it influences the core permeability and hence the pressure developed in the core as the binder decomposes. Lower permeability cores develop higher internal pressures if the other casting parameters are held constant. Effects of Type of Metal The type of metal significantly affected the amount of gas evolved from cores, as illustrated in Figure 8. Furan cores immersed into a gray iron melt at 1620 o C produced about 1400 cm 3 of gas within 60 seconds. The amount of gas evolved within the same period from cores immersed into an aluminum melt at 680 o C was only about 400 cm 3. The maximum gas evolution rate for cores immersed in iron and aluminum was 40 and 12 cm 3 /s. The maximum rate occurred within the first peak for cores immersed in iron. However, the maximum rate occurred at the second peak or after 30 seconds of contact for cores immersed in aluminum. The significant (more than three times) increase in volume can be attributed to the higher heating rate and higher final core temperature. Cores immersed in iron will experience a significantly higher heating rate as the cores are being heated through both conduction and radiation. The final core temperature when in contact with iron melt would be much higher than the temperature required to completely pyrolyze the binder. Cores immersed in aluminum will be heated mostly through conduction. Furthermore, the final core temperature will be lower than the temperature required to completely pyrolyze the binder.

164 148 The gas evolution rate and volume also increased when cores were immersed into an iron melt at higher temperature, as illustrated in Figure 8. The maximum gas evolution rate from Furan cores in contact with iron melt at 1510 and 1620 o C was 36 and 40 cm 3 /s. The total volume of gas was 1000 and 1380 cm 3 after 60 seconds of contact, respectively. The maximum gas evolution rate was reached 5 seconds earlier in cores immersed at lower melt temperature. The heat penetrated faster and deeper into the sand, when the cores were immersed into higher temperature melts. More binder was volatilized, which shifted the gas evolution rate peaks to a later time. Effects of Core Length and Core Metal Contact Area Cores made with phenolic urethane cold box (PUCB) binder and with a diameter of 2.86 cm and lengths of 5.1 cm and 8.9 cm were immersed in aluminum at 730 o C. Core length affected the total volume and the rate of gas evolved from sands made with 2.6% PUCB, as illustrated in Figure 9. Two rate peaks were observed at about 15 and 45 seconds after metal contact. The maximum rate peaks from 5.1-cm cores were 2.75 and 2.6 cm 3 /s, and the total volumes after 60 seconds of contact were 110 cm 3. Higher rate peaks and volumes were observed from the 8.9-cm length cores. The maximum rate peaks were 3.75 and 5.2 cm 3 /s, and the total volume was 200 cm 3. The increase in the volume of gas evolved was due to the increase in the core metal contact area. When the volume curves were divided by the metal contact area, the curves fell on top of each other, as illustrated in Figure 10. The maximum rates of gas evolved from 5.1- and 8.9-cm-length cores were almost the same at 0.07 cm 3 /cm 2 /s about

165 seconds after immersion. The volume of gas evolution evolved at 60 seconds was about 2.8 cm 3 of gas per cm 2 of surface area. Effects of Modulus (Specimen Diameter) and Metal Temperature Cylindrical cores made with 2.4% PUNB and having diameters of 2.9 cm and 6.1 cm were placed in contact with aluminum at 680 o C and 815 o C. The length of the core was held constant at 5.08 cm, with 1.3 cm of the core printed in a holder. The modulus (volume to surface area ratio) of cores having a diameter of 2.9 cm was 0.6 cm 3 /cm 2 and the value for the 6.1 cm diameter core was 1.1 cm 3 /cm 2. The number of peaks was not affected by the modulus of the core. The first rate peaks for cores with a modulus of 0.6 and 1.1 cm 3 /cm 2 produced gas evolution rates of 6 and 14.5 cm 3 /s, respectively, as illustrated in Figure 11 (680 o C). Three additional rate peaks were observed within 60 seconds from cores having a modulus of 0.6 and 1.1 cm 3 /cm 2. The volume of gas produced was higher for the larger modulus cores, with a value of about 420 cm 3, compared to a value of about 240 cm 3 from the lower modulus cores (680 o C). Cores in contact with higher temperature metal produced higher gas evolution rates and volumes, as illustrated in Figure 12. Cores with modulus values of 0.6 and 1.1 cm 3 /cm 2 in contact with 815 o C aluminum produced gas volumes of 400 and 640 cm 3, respectively, compared to volumes of 240 and 420 cm³ at 680 C. The peak rates increased from 5.5 to 10.5 cm 3 /s with longer metal contact time for the lower modulus cores. The number of rate peaks was not affected by the contact temperature. There were four rate peaks observed within 60 seconds of immersion. The height of the rate

166 150 peaks from cores immersed into aluminum at 815 o C with a modulus of 0.6 cm 3 /cm 2 continued to increase, and the highest rate peak was observed after about 55 seconds. The continuous increase in gas evolution rate peaks suggests that some volatile materials from the surface had condensed toward the center of the core and were re-evaporated as the center of the low modulus core was driven to higher temperatures. This increase in rate was not observed on the larger modulus cores because of the smaller contact area and thus the lower heating rate. Effects of Core Shape (Rectangular and Cylindrical) Rectangular cores made with 2.4% PUNB having a cross section of 2.6 x 2.6 cm and 4.1 x 4.1 cm were placed in contact with aluminum at 680 o C, with the results illustrated in Figure 13. The modulus values for the 2.6- and 4.1-cm square cores were 0.6 and 0.8 cm 3 /cm 2, respectively. Again, more gas was evolved from the higher modulus cores. The peak rate of gas evolution in the larger cores was 9 cm 3 /s after about five seconds, and the peak rate in the lower modulus cores was approximately 6.5 cm 3 /s at the same time when the cores were in contact with aluminum at 680 o C. The total volumes of gas evolved in 60 seconds were about 200 and 280 cm 3 from the cores with modulus values of 0.6 and 0.8 cm 3 /cm 2, respectively. Gas evolution rate and volume curves from same modulus (0.6 cm 3 /cm 2 ) cylindrical and rectangular cores in contact with aluminum held at 815 o C are illustrated in Figure 14. The total gas volume evolved from the cylindrical and rectangular cores

167 151 was 400cm 3, respectively. The first rate peak for cylindrical and rectangular cores was about 5 and 11 cm 3 /s. The first gas evolution rate was significantly higher for the rectangular cores because of the larger contact area and the edge effect. Further investigations are required to quantify this effect. The number of rate peaks was not affected by the core shape. Five rate peaks were observed within 100 seconds of metal contact. PROPOSED MODEL OF GAS EVOLUTION FROM CORES IN CONTACT WITH MOLTEN METAL A model describing the gas evolution of the binder in the core needs to be constructed for expanding the usefulness of the data generated using the gas evolution measurement device. The current set-up measured the gas evolution rate and volume from cores in contact with molten metal made with different geometries and immersed at various temperatures. However, the geometry of the core and contact temperature changes during casting. The experimental results show that the volume and rate of gas evolved changed depending on the shape, core modulus, and contact temperature. The shape, core modulus, and contact area and temperature affected the thermal history of the core and the volume and rate of gas evolution. The model should be able to extrapolate the experimental results to different core shape, modulus, and contact temperature. Extrapolating the experimental results to account for the effects of chemical reactions, such as core age, binder type, or additives, is not practical, as the financial resources required to analyze the very complex chemical bonding are enormous. Therefore, the authors are pursuing an empirical modeling approach. This model will represent the gas volume data with constants that separate the thermal effects (modulus

168 152 and melt temperature) and chemical effects (volatilization reactions) for the particular system. With this approach, once the gas evolution volume data is available for a particular core system (binder, core age, drying methods, and additives), the modeler can predict the gas evolution rate and volume for any core geometry and temperature from this core. The modeler can then combine this data with the predicted gas composition to calculate the local core pressure. Observations on the Burnt Profile and Variation of Binder Content after Casting The burn profile of a mold made with 4% PUNB and 63 GFN sand used to cast 5 cm (2 inches) diameter cylinder is shown in Figure 15.a. The iron was poured into the mold at approximately 1450 o C (2450 o F). The picture was taken the next day after both the melt and the mold had cooled to room temperature. The burn profile shows that the severity of the burnt varied with distance from the melt. The color of the sand closest to the melt changed from gray (or the original color of the mold) to white (the original color of the sand). The sand in this layer lost almost all of its binder and became loose sand. The next sand layer was carbon-black in color, indicating that most of the volatile was evaporated and the hydrocarbon component of the binder had started to burn. The next sand layer was a brown to yellow color. The volatile from the binder in this sand layer had started to evolve. This burn profile predicted that the binder content increased with distance from the melt. Furthermore, if the gas evolved from the binder was condensed, the binder content of the region next to the yellow sand layer would have been higher than the original binder content of the

169 153 mold. The binder content of these sand layers should be measured to confirm the existence of condensation. The binder content of these sand layers was measured by carefully grinding the sand at 3.8 cm (1.5 inch) below the mold top surface. The LOI results (Figure 15.b) indicated that there were slight increases at the region next to the yellow sand layer. The binder content was increased to 5.3%, or about 1% higher than the original binder content of the mold. These results suggested that some of the gas evolved had condensed at the sand layer with temperature less than the volatilization temperature of the binder. The physical model is constructed based on these observations and presented in the following section. Physical Model of Gas Evolution from Cores in Contact with Molten Metal The current model of gas evolution from cores in contact with molten metal is illustrated in Figure 16. This schematic illustrates the outermost volume of core sand in initial contact with the molten metal. Three zones exist within the core. The zone closest to melt is the volatilization zone (VZ). The temperature rises above the volatilization temperature, resulting in the volatilization of the binder. The binder content in the volatilization zone is reduced and can be completely removed depending on the distance from the melt. The gas evolved in the VZ flows toward the core print indicated by the arrow in the schematic. The lower temperature that exists in the adjacent sand layer resulted in the condensation of the volatilized gases. This zone where the binder concentration increases is called the condensation zone (CZ). The gas condensation releases heat into the sand in the CZ and increases the apparent conductivity. The zone

170 154 furthest away from the melt is called initial zone (IZ), where the temperature and the binder content have not been affected. The volatilization and condensation zones become thicker at the expense of the initial zone as the sand heats up with time. A typical temperature profile at the center of the hotbox core with diameter of 2.9 cm is shown in Figure 17. Two inflection points were observed at 13 and 60 seconds after immersion and at temperatures of 100 o C and 275 o C, respectively. The first and second inflection points represented the condensation and evaporation of the binder component or, in this case, the moisture and the hydrocarbons. The moisture in the gas evolved started and finished re-evaporating at 13 and 29 seconds after immersion, respectively. Some hydrocarbons started and finished re-volatilizing at 60 and 64 seconds at the center of the core, respectively. The longer temperature arrest observed at the boiling point of moisture indicated that the binder contained more water than hydrocarbons. The temperature arrest at temperature of 100 o C and 275 o C was 16 and 4 seconds, respectively. The gas evolution rate and volume measured during immersion of this hotbox core are shown in Figure 18. The first peak and the second peak were observed at 20 and 60 seconds after immersion, respectively. The temperature profile at the center or at the coldest point of the core showed that the volume evolved up to the first rate peak was dominated by moisture. The volume evolved at the second rate peak was probably dominated by hydrocarbons. These results highlight two important points: a) the gases evolved as result of at least two volatilization reactions: the moisture and the hydrocarbons in the binder; and b) the gases condensed due to the large temperature gradient in the sand. Hence, the

171 155 development of an accurate model for predicting the gas evolution rate requires accurate predictions of temperature, the volume of gas evolved and condensed, and the volatilization rate of the binder. The following sections describe how the temperature, total gas volume evolved and condensed, and the gas composition were predicted. Finally, the modeling results are discussed to highlight further improvement necessary for this approach. Temperature Predictions Accurate temperature predictions are vitally important for calculating the amount of gas evolved and condensed at any given times and locations in the core. The temperature profile in a cylindrically shaped core as a function of time and distance from the core-metal interface was calculated using Equation 1 [4]: T ( r, x, t) T T T i = A C e 2 C 2 λ αt r λ Cr J o A r0 W e 2 W 2 λ αt r λw x Cos L Equation 1 [14] where, Jo = first-order Bessel functions, A C, λ c = coefficients used in the one-term approximate solution of transient one-dimensional heat conduction in cylinders, A w. λ W = coefficients used in the one-term approximate solution of transient one-dimensional heat conduction in plane wall, H = interfacial heat transfer (W/m 2 /K), and α = thermal diffusivity (m 2 /s). The thermal diffusivity was given by k α = Equation 2 ρ. Cp

172 156 where, k = thermal conductivity (W/m/K), ρ = density (kg/m 3 ), and Cp = heat capacity (J/kg/K). The accuracy of the temperature prediction using this calculation depends on the accuracy of our thermal conductivity and heat transfer input data. For the preliminary model, the heat transfer coefficient and thermal conductivity were assumed to be constant regardless of the modulus and melt temperature. The interfacial heat transfer coefficient was set to be 750 W/m 2 /K for all cores immersed into an aluminum melt at 680 and 730 o C. The bulk thermal conductivity increased with greater gas volume evolution and condensation. The variation in thermal conductivity values due to these two events should be magnified at a location closer to the mold-metal interface and at the beginning of immersion when the highest gas volume and condensation rates occur. The importance of accurate thermal conductivity data is highlighted in the following sections by the inaccuracy in the calculations for predicting the occurrence of rate peaks. Empirical Model for Calculating Gas Evolution Rates and Volumes The empirical model for predicting gas evolution rates and volumes must address 1) the volatilization of binder, 2) the condensation of evolved gas, and 3) the revolatilization of the condensed gas. The volatilizations of binder were calculated using the Kobayashi method, and the condensation of evolved gas were estimated using nucleation theory. Both processes will be discussed in the following sections.

173 157 Volatilization of Binder Bates et al. reported that the gases evolved from PUCB cores immersed in iron and steel were dominated by hydrogen and carbon monoxide. However, the gas volume was dominated by air and carbon dioxide when similar cores were in contact with aluminum [5]. These results suggest that at least two different thermal degradation mechanisms occurred in these cores. Kobayashi [6] proposed an empirical two-step competitive reaction model for predicting the volatile yield during coal combustion. This model postulated that at least two reactions occurred and that the reactions contributing the higher volatile yields had higher activation energies. This model assumed that both gas (volatile) and solid (condensates) are produced during the primary (lower energy) decomposition reactions. The condensates formed from the primary reactions and the remaining solid would require higher activation energy to decompose. The schematic of the reactions is shown below: Binder k 1 k 2 Volatile 1 (V 1 ) + Residue 1 (R 1 ) Volatile 2 (V 2 ) + Residue 2 (R 2 ) In the schematic, k 1 and k 2 are the rate of binder decomposition for reaction 1 and 2, respectively. V 1 and V 2 are the volume of gas evolved as a function of time. R 1 and R 2 are the residue left as a function of time. The reaction rate is given by

174 158 k 1 = Z 1 Exp (-E 1 /R.T) Equation 3 where, Z 1, E 1, R, and T are the frequency factors, activation energy for reaction 1, ideal gas constant, and temperature, respectively. The frequency factor can be estimated by using transition state theory [7] that Z kt = Equation 4 h 1 K 1 where, k is the Boltzmann constant, T is the temperature in a particular location, h is the plank constant, and K is the equilibrium constant for the formation of the gas. To predict the amount of gas evolved as a function of time, one needs to calculate the rate of gas evolution. The rate of gas evolution or dv/dt is equal to the reaction rate times the volume yet to be released. dv dt ( k + k )( V V ) = 2 1 Equation 5 where, V is the volatile yet to be released from the binder. The total volume of gas evolved at a particular time or V is then equal to ( t) V ( 1 Exp( ( k + k ) t)) V 1 2 = Equation 6

175 159 where, t represents the time the core in contact with molten metal. The maximum volatile yield or V is equal to ( k V + k V ) 1 1,max ( k + k ) ,max V = Equation 7 where, V 1,max and V 2,max is the maximum gas evolved from reaction 1 and 2. Condensation of Evolved Gases The high molecular weight gases evolved initially or at lower temperature are condensable, while the lower molecular weight gases or the elemental gases such as hydrogen, nitrogen, and oxygen are not condensing in the core. The condensation process is modeled as a heterogeneous nucleation phenomenon. The nucleation theory postulates that the nucleation rate is equal to the energy released divided by the temperature gradient squared [8] as described in Equation 8. EC ( T ) kc = Z C ( T, L) Exp Equation 8 2 RΔT where, Z C is the frequency factor, and E C is the energy released to the system due to condensation. The frequency factor, Z C can be estimated by kt = Equation 9 h Z C K 1L P

176 160 where, L P (dimensionless) is the weighted length the gasses have to travel to exit the core from the core print. The energy released, E C, can be calculated by E C 8πγ T = Equation 10 3L 3 2 C 2 V where, γ, T m, and L V are the gas/solid interfacial energy of the polymer, critical temperature, and latent heat of vaporization of polymer, respectively. This empirical model assigns six constants for each core mixture. The six constants are V 1,max, V 2,max, K 1, K 2, E 1, and E 2. These six constants are calculated from the gas evolution data of cores immersed at various temperatures and geometries. Once these constants are known, one can predict the gas evolution rate and volume accurately for any metal temperature and core geometry. This model can be combined with the published pressure equation [1] for predicting gas defects in casting. Re-Volatilization of Condensed Gases This model assumes that only gases generated from reaction 1 are condensable. The gases decomposed at a lower temperature or energy range are typically comprised of high molecular gases, which are readily condensable. These gases will re-volatilize at a higher temperature to produce lower molecular weight gases. This model assumes that all condensed gases from reaction 1 will participate in reaction 2. Consequently, the maximum volume evolved from reaction 2 depends on both on the amount of binder and condensed gases. The gases evolved at higher energy and temperature reactions (or

177 161 reaction 2) have lower molecular weight and therefore are non-condensable. Typical gases evolved from reaction 2 are hydrogen and nitrogen. Modeling Results The discussions of the modeling results are aimed at illuminating the strengths and weaknesses of the empirical approach and to highlight the importance of condensation and the sand temperature profile. The empirical model can predict the maximum gas evolution rate and volume accurately for cores with modulus of 0.6 and 1.1 cm 3 /cm 2 and temperature of 680 o C. However, the empirical model can only predict the total volume accurately and cannot predict the maximum rate peaks from cores immersed at 815 o C. It is suspected that the temperature profile existing in the sand creates pockets of high condensation regions and increases the condensation rate significantly. This increase in condensation rate further increases the sand thermal conductivity. Modeling Results: Without Condensation The model accurately predicted the total gas volume evolved for both 1.1 and 0.6 cm 3 /cm 2 modulus, as illustrated in Figures 19 and 20. However, without accounting for the condensation of evolved gases, the model did not predict the maximum gas evolution rates in cores with a modulus of 1.1 (Figure 19). The predicted and measured maximum gas evolution rates from cores with modulus 1.1 were 29 and 14.5 cm 3 /s, respectively. The model could predict the maximum gas evolution rate for the smaller diameter cores more accurately even without accounting for gas condensation. The predicted and

178 162 measured maximum gas evolution rates from cores with modulus 0.6 cm 3 /cm 2 were 6.4 and 6.1 cm 3 /s, respectively. One should also notice that the predicted timings of the rate peaks were off. This could be due to the 3- to 5-second delay of the measuring device and to the assumption that no condensation had occurred. The peak rates would shift to a later time because the binder content would increase with further distance from the core-metal interface. Modeling Results: With Condensation The calculated gas evolution rate curve matched the experimental results better when the condensation of gases was accounted in the calculations for cores with modulus of 1.1 cm 3 /cm 2 (Figure 21). The calculated and experimental gas evolution rates reached maximums at 17 and 14.5 cm 3 /s, respectively. The calculated maximum gas evolution rate was 29 cm 3 /s from cores with modulus of 1.1 cm 3 /cm 2 when no gas condensation was assumed (Figure 19). Although the calculations were able to predict the maximum gas evolution rate well, the calculations were still not adequate to predict the timing and the number of rate peaks. The inclusion of condensation calculations were not affecting the accuracy of the model in predicting the total volume of gas evolved and the gas evolution rate from low modulus cores (0.6 cm 3 /cm 2 ). The predicted and measured total volumes of gas evolved from cores with modulus of 1.1 cm 3 /cm 2 at 60 seconds were 400 and 420 cm 3, respectively. The predicted and measured maximum gas evolution rates for cores with modulus of 0.6 cm 3 /cm 2 were 5.9 and 6.2 cm 3 /s, respectively (Figure 22). The measured

179 163 and calculated volume were similar. The measured and calculated gas volumes at 60 seconds from cores with modulus 0.6 cm 3 /cm 2 were 240 and 250 cm 3, respectively. The percentage of total gas condensed that was calculated for various core modulus and melt temperature is illustrated in Figure 23. The amount of condensation decreased with time but increased with core modulus. Up to 50% of the gases evolved from cores with modulus 1.1 cm 3 /cm 2 and immersed into aluminum melt at 680 o C were predicted to condense in the beginning of immersion. After 100 seconds, only about 8% of the total gases evolved were condensed in these cores. The amount of gases condensed was initially higher when cores were immersed into higher temperature melt. Up to 72% and 35% of the total gas evolved was condensed when cores were immersed into 815 o C and 680 o C aluminum melts, respectively. However, the amount of condensed gases was similar at about 7% after 100 seconds of immersion. Figures 24 and 25 illustrate the development of volatilization, condensation, and initial zones at radial and axial sections of a core, respectively after 10 seconds of immersion. In a radial section (Figure 24) 0.3 cm from the core end, the temperature decreased from 680 o C to 313 o C at about 1 cm away from the interface and then increased to 352 o C at the radial center of the core. Volatilization is apparent when the binder concentration drops below the initial value of 2.4%. All of the binder was volatilized at the core-metal interface. The binder concentration remained below 2.4% up to a radial distance of 0.05 cm from the core-metal interface. Condensation is indicated when the binder concentration exceeds the initial binder concentration. The gas started to

180 164 condense at 0.05 cm from the interface. No initial zone was observed since the binder content was higher than the initial concentration up to the center of the core. Along the length and at the radial center of the core (or 1.43 cm from core metal interface), the temperature decreased from 680 o C to 25 o C at 3.81 cm from the core end (Figure 25). The binder was volatilized until 0.4 cm from the core end. At 0.4 cm from the core end, the gas evolved started to condense and increased the binder content deeper into the core. The binder content reached maximum at 0.76 cm from the core end. The higher condensation in the radial direction in this case highlights the importance of temperature gradient. Although the calculated temperatures were higher than 300 o C at any point along the radial axis, condensation still existed because of the large temperature gradient. The condensation at this particular location was particularly high because it was also close to the core end. No condensation was observed at the center of the core or 1.9 cm away from the core end. Limitations of the Model This model is less accurate in predicting the timing of gas evolution rate peaks, the volume, and the gas evolution rate peaks within 10 seconds of immersion. The number of gas evolution rate peaks is also lower than that observed from the experiment. The experimental and calculated gas evolution rate and volume curves for cores immersed in aluminum at 815 o C are illustrated in Figure 26. The experimental results show that there were four rate peaks observed at 5, 18, 35, and 52 seconds of immersion. The last peak occurred after 52 seconds at 10.4c m 3 /s. The calculation shows that the maximum gas evolution rate peak would be reached after just 19 seconds and at 9.2

181 165 cm 3 /s. Furthermore, only two gas evolution rate peaks were observed in the calculated rate curve. There are two possible causes that could lead to the inaccuracy of the model in predicting the timing of the gas evolution rate peaks. The first cause might arise from the inaccuracy in calculating core temperature. Better temperature prediction can be achieved by calculating core temperature using commercial software. The second possible cause is that the condensed gas might evolve at a different rate and/or at a higher energy than the reaction to produce the low molecular weight gases. The model then needs to include three reactions instead of two. The addition of the third reaction into the model should increase the maximum number of peaks to five and shift the maximum peak to letter time. The difference in both volume and rate in the beginning of immersion is greater when cores were immersed in iron at 1350 o C, as shown in Figure 27. The volume of gas evolved within 0.25 seconds was calculated to be 104 cm 3, which was much higher than the volume measured at 0.2 cm 3. The significant differences observed might be due to the change in the re-volatilization reactions of the binder. The binder might de-volatilize at higher energies when exposed to an iron melt or significantly higher temperature. Therefore, it is advisable to assign those six constants not only for each binder system but also for each melt type. Predictions of the Amount of Low Molecular Weight Gas Evolved This model can also be used for predicting the amount of low and high molecular weight gases evolved. The calculations predicted that the amount of low molecular

182 166 weight gas, such as hydrogen, was very low for all cores immersed at 680 o C (Figure 28). The volume of low molecular weight gases was about 75 cm 3 or less than one-third of the total gas volume evolved. However, almost 75% of the total gas evolved from cores immersed into 815 o C aluminum bath was predicted to contain low molecular weight gases. These results suggested a higher rate of re-evaporation of the high molecular weight gases when cores were immersed into a higher temperature melt. These results were consistent with the data reported by Bates et al. [3] that almost no hydrogen was present in the gases evolved from a large mold in contact with aluminum. However, the hydrogen gas dominated the gas composition when iron or steel was poured into the same size molds. CONCLUSIONS The gas evolution rates from cores were measured and found to be affected by metal contact area, core age, modulus, and metal temperature. The core-metal contact area was the most significant variable affecting heat transfer from the metal into the core. Cores in contact with higher temperature metal also produced more gas. The gas evolution rates were not affected by the core length and permeability. This data was used to develop a model to describe gas evolution from cores as a function of geometry and contact area with metal. The goal is to predict the gas pressure inside cores. If the gas pressure exceeds the metal-head pressure, gas will be blown into the metal, resulting in hydrogen absorption, blow holes, pinholes, and oxide trails in aluminum castings.

183 167 A model based on chemical reactions for producing the gas evolved was proposed. The model can accurately predict the evolved gas volume. However, more understanding and data are needed to match the experimental rate curves. The condensation and re-evaporation of high molecular weight gases needs to be better understood. In addition, more data is needed, especially on the composition of the evolved gas as a function of time and more precise interfacial heat transfer and sand thermal conductivity data are also required. ACKNOWLEDGEMENTS The authors would like to thank the members of core gas consortium for guidance and financial support throughout this study. The authors also thank Dr. Preston Scarber, Jr., for his invaluable insights. REFERENCES 1. Winardi. L., Littleton, H.E. and Bates, C.E., Pressures in Sand Cores, to be published in AFS Transactions (2007.) 2. American Foundry Society, AFS S: Grain Fineness Number, AFS GFN, Calculation, Mold and Core Test Handbook, 3 rd ed., pp. 1-11, American Foundry Society, Des Plaines, Illinois (2001). 3. Winardi, L., Littleton, H.E. and Bates, C.E., New Technique for Measuring Permeability of Cores Made from Various Sands, Binders, Additives, and Coatings, AFS Transactions, vol. 113, pp (2005). 4. Cengel, Y.A., Heat Transfer: A Practical Approach, 2 nd ed., chapter 4, McGraw-Hill (2002). 5. Bates. C.E. and Monroe, R.W., Mold Binder Decomposition and Its Relation to Gas Defects in Castings, AFS Transactions, vol. 83, pp , (1981).

184 Kobayashi, H., Howard, J.B. and Sarofim, A.F., Coal Devolatilization at High Temperatures, Sixteenth Symposium (International) on Combustion, pp , (1976). 7. Frost, A.A. and Pearson, R.G., Kinetics and Mechanism, 1 st ed., pp. 188, John Wiley and Sons (1952). 8. Porter, D.A. and Easterling, K.E., Phase Transformations in Metals and Alloys, 2 nd ed., pp. 188, Chapman-Hall (1992).

185 Length (cm) Table 1. Cores used in this study. Diameter Grain or Body Binder Shape % Binder Size Width Type (AFS GFN) (cm) Modulus (cm 3 /cm 2 ) Age (days) Melt Temp. ( o C) Permeability Air Flow Rate (cm 3 /cm 2 /s) Cylinder 2.9 PUCB # Al Cylinder 2.9 PUCB #2 Epoxy Acrylic (E.A) Al Cylinder 2.9 PUCB # Al PUCB #2 3.8 Cylinder 2.9 AV-VS PUCB # Al AV-V PUCB 3.8 Cylinder 2.9 Water Based Coating Al Cylinder PUNB Al Cylinder 2.9 PUNB Al Rectangular Cylinder 2.9 PUCB Al Al 3.8 Cylinder 2.9 PUNB Al Fe Cylinder 2.9 PUCB Al Cylinder 2.9 Hotbox Al

186 Resistance Heaters Refill Funnel H Beake Precision Balance Furnace Data Acquisition System Figure 1. Schematic of the displacement apparatus used to determine gas evolution volumes and rates during contact with molten metal [1]. 170

187 Gas Evolution Rate (cm 3 /s) % Epoxy Acrylic (E.A) T = 730 o C Modulus = 0.6 cm 3 /cm 2 1.6% PUCB T = 730 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume E.A Gas Volume PUCB Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 2. Effects of binder type on gas evolution rate and volume from cylindrical cores made with 1.6% PUCB and epoxy acrylic resins in contact with aluminum at 730 o C (1350 o F). 171

188 Gas Evolution Rate (cm 3 /s) % PUCB T = 730 o C Modulus = 0.6 cm 3 /cm 2 0.8% PUCB T = 730 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume 1.6% PUCB Gas Volume 0.8% PUCB Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 3. Effects of binder content on gas evolution rate and volume from cylindrical cores made with 1.6% and 0.8% PUCB resin and with and without water-based coating in contact with aluminum at 730 o C (1350 o F). 172

189 Gas Evolution Rate (cm 3 /s) % PUCB (No Additive) T = 730 o C Modulus = 0.6 cm 3 /cm 2 1.6% PUCB (AV-V) T = 730 o C Modulus = 0.6 cm 3 /cm 2 1.6% PUCB (AV-VS) T = 730 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume No Additive Gas Volume AV-VS Gas Volume AV-V Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 4. Effects of addition of AV-VS and AV-V additives on gas evolution rate and volume from cylindrical cores made with 1.6% PUCB resin in contact with aluminum at 730 o C (1350 o F). 173

190 Epoxy Acrylic Coated T = 730 o C Modulus = 0.6 cm 3 /cm Gas Evolution Rate (cm 3 /s) Uncoated T = 730 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume Coated Gas Volume Uncoated Epoxy Acrylic Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 5. Effects of coating on gas evolution rate and volume from cylindrical cores made with 1.6% epoxy acrylic resin and with and without water-based coating in contact with aluminum at 730 o C (1350 o F). 174

191 1 0.9 Small Cylindrical T = 1325 o F 75 GFN Gas Evolution Rate (cm 3 /cm 2 /s) Small Cylindrical T = 1325 o F ` 75 GFN 10 Days Later Gas Volume 10 Days Later Gas Volume Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 6. Effects of core age on gas evolution rates and volumes in PUCB cores. 175

192 Gas Evolution Rate (cm 3 /s) Small Cylindrical ρ = 1.78g/cm 3 T = 718 o C 75 GFN Small Cylindrical ρ = 1.60g/cm 3 T = 718 o C 50 GFN Gas Volume 75 GFN Gas Volume Evolved (cm 3 ) 10 5 Gas Volume 50 GFN Time (s) 0 Figure 7. Effects of sand grain fineness and density on gas evolution rate and volume. Cores made with higher finer sand or higher GFN value had higher density and therefore lower permeability. 176

193 50 45 Rate: Iron 2% Furan 1621 o C Rate: Iron 2% Furan Vol: Iron 1621 o C 1600 Gas Evolution Rate (cm 3 /s) o C Vol: Iron 1510 o C Vol: Al 696 o C Gas Volume (cm 3 ) 10 Rate: Al 2% Furan o C Time (s) 0 Figure 8. Effects of melt type and temperature on gas evolution rate and volume. 177

194 Gas Evolution Rate (cm 3 /s) Rate T = 730 o C L = 8.9cm Rate: T = 732 o C L = 5.08cm Vol: T = 732 o C L = 5.08cm Vol: T = 730 o C L = 8.9cm Gas Volume Evolved (cm 3 ) Time (s) Figure 9. Effects of core length on gas evolution rate and volume from cylindrical cores made with 2.6% PUCB resin in contact with aluminum at 730 o C

195 Gas Evolution Rate (cm 3 /cm 2 /s) Rate: T = 730 o C L = 8.9cm Rate: T = 732 o C L = 5.08cm Vol: T = 730 o C L = 8.9cm Vol: T = 732 o C L = 5.08cm 9 6 Gas Volume Evolved (cm 3 /cm 2 ) Time (s) Figure 10. Effects of core length on gas evolution rate and volume from cylindrical cores made with 2.6% PUCB resin in contact with aluminum at 730 o C when both curves were divided by unit area wetted by metal

196 Large Cylinder T = 688 o C Modulus = 1.1 cm 3 /cm Gas Evolution Rate (cm 3 /s) Small Cylinder T = 673 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume Large Cylinder Gas Volume Small Cylinder Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 11. Effects of core modulus on gas evolution rates and gas volume from cylindrical cores made with 2.4% PUNB and modulus of 0.6 and 1.1 cm 3 /cm 2 in contact with aluminum at 680 o C. 180

197 Gas Evolution Rate (cm 3 /s) Large Cylinder T = 800 o C Modulus = 1.1 cm 3 /cm 2 Small Cylinder T = 814 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume Large Cylinder Gas Volume Small Cylinder Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 12. Effects of core modulus on gas evolution rates and total gas volume evolved from cylindrical cores made with 2.4% PUNB and modulus of 0.6 and 1.1 cm 3 /cm 2 in contact with aluminum at 815 o C. 181

198 Gas Evolution Rate (cm 3 /s) Small Rectangular T = 667 o C Modulus = 0.6 cm 3 /cm 2 Large Rectangular T = 670 o C Modulus = 0.8 cm 3 /cm 2 Gas Volume Large Rectangular Gas Volume Small Rectangular Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 13. Effects of core modulus on gas evolution rates and total gas volume from rectangular cores in contact with aluminum at 680 o C. 182

199 Gas Evolution Rate (cm 3 /s) Small Rectangular T = 814 o C Modulus = 0.6 cm 3 /cm 2 Small Cylinder T = 816 o C Modulus = 0.6 cm 3 /cm 2 Gas Volume Small Rectangular Gas Volume Small Cylinder Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 14. Effects of core shape on gas evolution rates and total gas volume from rectangular and cylindrical cores in contact with aluminum at 815 o C. 183

200 cm 15.a 6% 5% Possible Condensation Locations Binder Content (LOI, %) 4% 3% 2% 1% Original Binder Content: 4.3%. The error bars represent 95% Confidence Interval. Binder Content at the outer radius of the mold was unaffected. Mold Radius: 13.3 cm 0% Distance from melt (cm) 15.b Figure 15. The mold burnt profile (a) and binder content (b) as a function of distance from melt are illustrated. The mold was made with 4.3% PUNB and 63 GFN silica sand. The color of the mold changed depending on the severity of the burnt. The binder content, as measured from LOI tests, increased and reached up to 5.3% at about 4.3 cm from the melt. The binder content at the outer radius of the mold remained at 4.3% or unaffected by the heat.

201 2. Condensation 1. Volatilization 3. Initial Metal Sand Temp./Comp. Initial Binder Comp. Initial Temp Binder Comp. Temp. Gas Flow X (Distance Away from Metal) Figure 16. Schematic describing the physical model of the gas evolution from resin during contact with molten metal. The direction of gas flow is indicated by the arrow pointing away from the metal/core sand interface. Three zones exist as soon as a core is in contact with molten metal. The zone closest to the mold metal interface is called the volatilization zone (VZ). The sand temperature in the VZ rises above the binder volatilization temperature. The amount of binder is reduced and is completely volatilized at the volume very close to the surface. The gases condense in Zone 2 due to the lower temperature. The condensation of evolved gas increases the binder concentration in Zone 2. The temperature and the binder concentration in Zone 3 have not changed and still reflect the initial conditions. 185

202 Binder Volatilization Temperature ( o C) Moisture Evaporation Time (s) Figure 17. Temperature profile at the center of 2.1% hotbox core immersed in aluminum at 730 o C. The core length and diameter were 3.8 cm (1.5 inch) and 2.9 cm (1.125 inch), respectively. 186

203 Rate (cm 3 /s) Volume (cm 3 ) Time (s) 0 Figure 18. Gas evolution rate and volume from 2.1% hotbox cores immersed in aluminum at 730 o C. 187

204 25 2.4% PUNB cores in contact with Al at 680 o C. Cores diameter and length was 6.1cm and 3.8cm, respectively Gas Evolution Rate (cm 3 /s) Calc. Rate Exp. Rate Calc. Vol Gas Volume Evolved (cm 3 ) Exp. Vol Time (s) 50 0 Figure 19. Comparison between calculated and measured gas rate and volume from larger-modulus (1.1 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C. The calculations did not account for the condensation of the gases evolved. The core was made with PUCB binder and had diameter and length of 6.1 and 3.8 cm, respectively. 188

205 25 2.4% PUNB cores in contact with Al at 680C. Cores diameter and length was 2.9cm and 3.8cm, respectively Gas Evolution Rate (cm 3 /s) Exp. Rate Calc. Rate Exp. Vol Gas Volume Evolved (cm 3 ) Calc. Vol Time (s) Figure 20. Comparison between calculated and measured gas rate and volume from smaller-modulus (0.6 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C. The calculations did not account for the condensation of the gases evolved. The core was made with PUCB binder and had diameter and length of 2.9 and 3.8 cm, respectively

206 25 2.4% PUNB cores in contact with Al at 680 o C. Cores diameter and length was 6.1cm and 3.8cm, respectively Gas Evolution Rate (cm 3 /s) Calc. Rate Exp. Rate Calc. Vol. Exp. Vol Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 21. Comparison between calculated and measured gas rate and volume from larger-modulus (1.1 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C when condensation was accounted. The core was made with PUCB binder and had diameter and length of 6.1 and 3.8 cm, respectively. 190

207 25 2.4% PUNB cores in contact with Al at 680C. Cores diameter and length was 2.9cm and 3.8cm, respectively Gas Evolution Rate (cm 3 /s) Exp. Rate Calc. Rate Exp. Vol Gas Volume Evolved (cm 3 ) Calc. Vol Time (s) 0 Figure 22. Comparison between calculated and measured gas rate and volume from smaller-modulus (0.6 cm 3 /cm 2 ) cores immersed in aluminum at 680 o C when condensation was accounted. The core was made with PUCB binder and had diameter and length of 2.9 and 3.8 cm, respectively. 191

208 100 Percentage of Gas Condensed at Particular Time (%) Mod. = 0.6 Temp. = 815 o C Mod. = 0.6 Temp. = 680 o C Mod. = 1.1 Temp. = 680 o C Time (s) Figure 23. Percentage of the total gases evolved that were being condensed calculated for various core modulus and melt temperature as a function of time. The high molecular weight gases were condensed as they interacted with the colder sand on the way out to the core print. 192

209 3 2.5 % Binder at 0.3cm from Core End % Binder Content at 10 sec Condensation Zone Volatilization Zone Initial Binder Content Temperature Original Binder Content was 2.4% Melt Temperature = 680C Core Radius = 1.48cm Core Modulus = 0.6cm 3 /cm Temperature after 10sec. (C) 0.3cm R=1.43cm L = 3.81cm Radial Distance from Core-Metal Interface (cm) 24.a. The calculated volatilization and condensation zones developed in radial direction when 0.6 cm 3 /cm 2 modulus cylindrical cores immersed into aluminum at 680 o C. The temperature and binder content curves were calculated for various radial distances from core-metal interface at 0.3 cm from the core end and at 10 seconds after immersion. 24.b. Schematic illustrating the location where the temperature and binder content curves were calculated. The total length of the core was 3.81 cm. The radius of the core was 1.43 cm. 193

210 3 2.5 Condensation Zone Initial Zone R = 1.43cm % Binder Content at 10 sec Volatilization Zone Initial Binder Content % Binder at the radial center and along the axial axis of Core. Temperature Original Binder Content was 2.4% Melt Temperature = 680C Core Length = 3.81cm Core Modulus = 0.6cm 3 /cm Temperature after 10sec. (C) L =3.81cm Distance from Core End (cm) 25.a. The calculated volatilization, condensation, and initial zones developed in axial direction when 0.6 cm 3 /cm 2 modulus cylindrical cores were immersed into aluminum at 680 o C. The temperature and binder content curves were calculated for various axial distances from the core-metal interface at the radial center and at 10 seconds after immersion. 25.b. Schematic illustrating the location where the temperature and binder content curves were calculated. The total length of the core was 3.81 cm. The radius of the core was 1.43 cm. 194

211 % PUNB cores in contact with Al at 815 o F. Cores diameter and length was 2.9cm and 3.8cm, respectively Gas Evolution Rate (cm 3 /s) Calc. Rate Exp. Rate Calc. Vol Gas Volume Evolved (cm 3 ) Exp. Vol Time (s) 0 Figure 26. Comparison between calculated and measured gas rate and volume from cores immersed in aluminum at 815 o C. The core was made with PUCB binder and had diameter and length of 2.9 and 3.8 cm, respectively. The core modulus was 0.6 cm 3 /cm

212 % PUNB cores in contact with Al at 1350 o C. Cores diameter and length was 6.1cm and 3.8cm, respectively Gas Evolution Rate (cm 3 /s) Exp Rate Calc Vol. Calc Rate Exp. Vol Gas Volume Evolved (cm 3 ) Time (s) 0 Figure 27. Comparison between calculated and measured gas rate and volume from cores immersed in iron at 1350 o C. The core was made with 2.4% PUCB binder and had diameter and length of 2.9 and 3.8 cm, respectively. The core modulus was 0.6 cm 3 /cm

213 Volume of Low Molecular Weight Gases (cm 3 ) Mod. = 0.6 Temp. = 680 o C Mod. = 0.6 Temp. = 815 o C Mod. = 1.1 Temp. = 680 o C Time (s) Figure 28. Calculated volume of low molecular weight gases evolved from cores with 0.6 and 1.1 cm 3 /cm 2 modulus and immersed into 680 and 815 o C aluminum melt. 197

214 199 CONCLUSIONS A recent survey of foundry engineers and personnel indicated that eliminating porosity defects should be the top research priority (41). The current research focused on developing methods for predicting the occurrence of blow defects, a major source of porosity in castings. Four areas were investigated including 1) development of improved methods for measuring the permeability and gas evolution from cores, 2) development of methods to calculate the internal pressure in cores, 3) measurement of the effects of production variables on core permeability, gas evolution, and internal pressure, and 4) development of a model to predict gas volatilization and condensation of binder when cores were in contact with molten metal. The conclusions from each area are listed in the following sections. Development of Gas Evolution and Permeability Measurement Devices The foundry industry currently uses the LOI test to predict the tendency for a core to cause a blow defect. This test does not predict the gas evolution from a core immersed in molten metal, because the cores see an abundance of oxygen that is not present during the casting process. A device was developed for measuring the gas evolution for a core immersed into molten metal. Condensation from the measurement procedure was minimized by heating the device above the binder pyrolyzation temperature. The current standard for measuring permeability gives lower permeability values than would be present in the conditions encountered in a foundry. This is because the

215 200 current device assumes laminar flow and doesn t account for momentum effects. Furthermore, the device does not take into account gas composition and temperature. Because of these limitations, the data generated using this device can only be used for qualitative comparison and not for predicting the core permeability as a function of temperature and gas composition. The permeability measurement device developed for this project could measure permeability regardless of the temperature and composition of the gas by describing the pore structure in the core. The pore structure was described by two constants, which can be used in a simulation program to predict the permeability over a wide range of core internal pressure. Instead of reporting the Darcy s constant, the core permeability was compared by determination of the allowable air flow rates through a core at a set pressure gradient. Cores that permit higher flow rates were more permeable. The accuracy of these two devices was validated. The gas evolution measurement device accurately measured the gas flow up to 60 cm 3 /sec. The permeability measurement device can be used to calculate the air flow rates through cores up to 1000 cm 3 /sec and 41 kpa (6 psi) or 57 cm (22.4 inch) of iron. Methodology to Predict Core Internal Pressure Prior to this research, there are two methodologies for calculating core internal pressure. Both Worman and Nieman (Equation 2) and Campbell (Equation 3) proposed methods that did not accurately predict the pressure. The Worman and Nieman method over predicted the pressure, while Campbell s under-predicted the pressure. Both methods were limited by the techniques available at the time for measuring permeability

216 201 and the gas evolution rate. These limitations forced them to rely on the LOI test results and introduce constants that could vary under different casting conditions. The methodology proposed in this study benefits from the advanced in techniques for measuring the gas evolution rate and core permeability. The new methodology predicted that the core internal pressure increases with increasing gas evolution rate, core-metal contact area, gas viscosity, and distance to the core print. The pressure decreased with increasing core permeability and gas flow area (or print area). This methodology was verified three times to validate the accuracy of this method. The experimental results showed that the pressure curves followed the gas evolution rate curves, which reveals the importance of gas evolution rate measurements for predicting pressure. Effects of Production Variables on Gas Evolution The type and amount of binder affected the gas evolution. The maximum gas evolution rate produced from 1.6% phenolic urethane cold box (PUCB) and 1.6% epoxy acrylic (EA) cores immersed in aluminum at 730 o C was 3 and 6 cm 3 /s, respectively. The corresponding total gas volume produced at 60 seconds was 140 and 160 cm 3. Increasing the binder content from 0.8% to 1.6% increased the maximum gas evolution rate from 3 to 5 cm 3 /s. Similarly, the total gas volume increased from 120 to 225 cm 3, respectively. The variations in sand grain fineness and density in the range of this study did not affect the volume or rate of gas evolution. The gas evolution rate and total volume evolved from PUCB cores made with 50 and 75 and GFN were similar at 34 cm 3 / s and 290 cm 3, respectively.

217 202 The addition of additives and the application of a coating significantly increased the gas evolution rate and volume from PUCB cores. The gas evolution rate and volume doubled from 3 to 7.5 cm 3 /s and from 140 to 300 cm 3 when anti-veining-vs were added into 1.6% PUCB cores and immersed in aluminum at 730 o C. Cores washed with waterbased coating had a higher first rate peak and produced a higher volume of gas. The first rate peak was tripled from 3 to 10.5 cm 3 /s, and the total gas volume was doubled from 160 to 320 cm 3 when the core was coated and immersed into aluminum melt at 730 o C. Core age significantly reduced the amount of gas evolved. The peak rates from hotbox cores decreased from 35 and 6 cm 3 /s after 10 days of aging in a sealed and controlled environment. The gas volumes were about 300 and 200 cm 3 after the same time lapse. The reduction in the rate and volume of gas evolved might be attributed to the evaporation of the solvent at the core surface and to the continuing polymerization reaction of the binder. Cores in contact with higher temperature metal produced higher gas evolution rates and volumes. The maximum gas evolution rate from cores made with PUCB resin and with 0.6 cm 3 /cm 2 modulus (volume to surface area ratio) was doubled from 5.5 to 10.5 cm 3 /s when the aluminum melt temperature was raised from 680 to 815 o C. However, the first peak rate that occurred within the 5 seconds after immersion was the same when these cores immersed in these two temperatures. The same increase in temperature also increased the total volume of gas evolved from 240 to 400 cm 3 after 60 seconds. The gas evolution rate and volume produced from cores increased along with increases in core modulus, length, and core-metal contact area. The effect of core length

218 203 was found to be related to the increase in the core-metal contact area. Both the gas evolution rate and volume were not affected by the core length when these data were divided with the corresponding core-metal contact area. Rectangular cores produced higher gas evolution rate and volume than cylindrical cores when immersed into the melt with the same modulus. Effects of Production Variables on Core Permeability Other variables being equal, the permeability depends mainly on compaction. The degree of compaction along the core length can be measured easily by the variation in density. Increasing density or compaction of cores produced from the same batch of sand containing the same amount of binder and additive will decrease the permeability. All coatings tested during this study decreased the surface permeability of the core. The coating permeability varied depending on its thickness (or geometrical dimensions), packing (or density), and pore structure. All of the variables mentioned were tied to the composition (type and amount of refractory, solution: water or alcohol based, binders, suspension, and preservative agents) and to the method of coating application (including the drying procedure). In general, the type and amount of binder, the sand type, and additives did not affect the core permeability. However, one additive was found to slightly affect the permeability. This inconsistency prompts us to not disregard the influence of additive on permeability.

219 204 Proposed Model to Predict Gas Volatilization and Condensation of Cores in Contact With Molten Metals A model was proposed to describe the volatilization and condensation of binder for cores in contact with molten metal. This model proposed that three zones existed when the core was in contact with molten metal: volatilization, condensation, and initial zones. The volatilization zone was the zone closest to the melt. The temperature in this zone rose rapidly above the volatilization temperature and resulted in the volatilization of the binder. The binder content in the volatilization zone was reduced and can be completely removed depending on the distance from the melt. The gas evolved in the VZ flowed toward the core print or to the lower temperature zone where the gases condensation were condensed. This zone where the binder concentration increased is called the condensation zone (CZ). The gas condensation released heat into the sand in the CZ and increased the apparent conductivity. The zone furthest away from the melt was called the initial zone (IZ), where the temperature and the binder content were not affected. The amount of gas evolved in the VZ was calculated empirically by assigning six constants for each core system. The amount of gas condensed was estimated using the nucleation theory by assuming that the gas condensed heterogeneously in the sand. This approach allowed the modeler to predict the gas evolution rate and volume for any core geometry and temperature and then combine this data with the predicted gas composition to calculate the local core pressure. This model was less accurate in predicting the timing of gas evolution rate peaks, the volume, and the gas evolution rate peaks within 10 seconds of immersion, especially

220 205 when cores were immersed in different melts (iron and aluminum). The number of gas evolution rate peaks was also lower than that observed from the experiment. There are two possible causes of the inaccuracy of the model in predicting the timing of gas evolution rate peaks. The first might be inaccuracy in calculating core temperature. Better temperature prediction can be achieved by calculating core temperature using commercial software. The second possible cause is that the condensed gas evolved at a different rate and/or at a higher energy than the reaction to produce the low molecular weight gases. The model then needs to include three reactions instead of two. The addition of the third reaction into the model will surely increase the maximum number of peaks to five and shift the maximum peak to a later time. Future Research The future research will focus on extending the current steady-state methodology, which is reasonably accurate in predicting pressures for simple shapes, to a transient methodology for more complex geometries. This effort will provide a tool to accurately predict the probability of gas defect formation for complex geometry cores and provide improved mold/core design to the foundry before the first casting is poured. Several tasks must be accomplished to complete the transient model. 1. Accurate prediction of the core temperatures is vitally important. The effects of binder condensation and volatilization or mass transfer on sand apparent conductivity must be accounted for, because we know that the bulk of the core is heated mainly through the mass transfer and condensation rather than heat conduction (Appendix A). An inability to account for the effects of mass transfer

221 206 will result in under-prediction of the rate of binder decomposition (Appendix A) and the gas evolution rate. 2. The validity of the proposed pressure equation needs to be investigated in green sand molds. Green sand produces up to ten times more gas volume than regular chemically bonded sand. The gas produced is more readily condensable due to the high water concentration. The significant increase in the amount of water in the sand will clog the sand pore and reduce its permeability. The local pressure in the mold will then be affected by both the change in permeability and the distance the gas can travel without condensing. The pressure will increase with decreasing permeability. However, the pressure will increase with the reduction in the gas travel distance. 3. The composition of the gas evolved needs to be documented for accurate pressure predictions, especially for calculating the gas viscosity and compressibility. The gas viscosity data is very important for calculating the local pressure. The molecular weight of the gas needs to be determined in order to calculate the compressibility factor and extend the steady-state equation into a dynamic predictive model. 4. The transient model must be validated for complex geometries at typical foundry pouring temperatures. The effects of core age and coating drying methods have not been extensively evaluated. Similarly, the gas evolution rates and volumes immersed in magnesium must be recorded. Future binders also need to be included into the database.

222 207 GENERAL LIST OF REFERENCES 1. Scarber, P. Jr., Bates, C.E. and Griffin, J., Effects of Mold and Binder Formulations on Gas Evolution When Pouring Aluminum Castings, AFS Transactions, vol. 114, paper no. 130 (2006). 2. Scarber, P. Jr., Gas Defects in Aluminum Castings from Cores and Mold Washes, Foundry Management and Technology, vol. 132, no. 10, pp , (2004). 3. Wieser, P.F., Gases in Cast Steel, Cast Metal Institute: Advanced Seminar of Gases in Cast Metals, March (1976). 4. Svoboda, J.M., Fundamentals of Physical Chemistry Relating to Gases in Cast Metals, Cast Metal Institute: Advanced Seminar of Gases in Cast Metals, March (1976). 5. Trojan, P.K. and Flinn, R.A., Gases in Non-Ferrous Castings, Cast Metal Institute: Advanced Seminar of Gases in Cast Metals, March (1976). 6. Levelink, H.G., Julien, F.P.M.A. and De Man, H.C.J., Gas Evolution in Molds and Cores as the Cause of Casting Defects, AFS International Cast Metals Journal, March, pp , (1981). 7. American Foundry Society, Analysis of Casting Defects, 2 rd ed., pp , American Foundry Society, Des Plaines, Illinois (1966). 8. Rowley, M.T., International Atlas of Casting Defects, 1 st ed., pp , American Foundry Society, Des Plaines, Illinois (1974). 9. McSwain, R.H., Bates, C.E. and Scott, W.D., Iron-Graphite Surface Phenomena and Their Effects on Iron Solidification, AFS Cast Metal Research Journal, pp , (Dec 1974). 10. Scott, W.D. and Bates, C.E., Decomposition of Resin Binders and the Relationship Between Gases Formed and the Casting Surface Quality, AFS Transactions, vol. 83, pp , (1975). 11. Scott, W.D. and Bates, C.E., Decomposition of Resin Binders and the Relationship Between Gases Formed and the Casting Surface Quality, Part II, Gray Iron, AFS Transactions, vol. 84, pp , (1976).

223 Scott, W.D. and Bates, C.E., Decomposition of Resin Binders and the Relationship Between Gases Formed and the Casting Surface Quality, Part III, AFS Transactions, vol. 85, pp , (1977). 13. Scott, W.D., Goodman, P.A. and Monroe, R.W., Gas Generation at the Mold-Metal Interface, AFS Transactions, vol. 86, pp , (1978). 14. Sringagesh, Burn Out Sand for Systems with Organic No-Bake Binders, AFS Transactions, vol. 84, pp , (1976). 15. Hughes, I.C.H., The Role of Gases in the Structure of Cast Iron, AFS Transactions, vol. 76, pp , (1968). 16. Chen, F. and Keverian, J., Effect of Nitrogen on Subsurface Pinholes in Steel Castings, Modern Castings, July, pp , (1966). 17. Dawson, J.V., Kilshew, J.A. and Morgan, D.A., The Nature and Origin of Gas Holes in Iron Castings, AFS Transactions, vol. 73, pp , (1965). 18. Pehlke, R.D. and Elliott, J.F., Solubility of Nitrogen in Liquid Iron Alloys, Trans. AIME, vol. 218, pp , (1950). 19. Sims, C.E. and Zapffe, C.A., The Mechanism of Pinhole Formation, Trans. Of AFA, vol. 58, pp , (1950). 20. Portevin, A.M., Gases and Naturally Occuring Blowholes in Foundry Practice, AFS Transactions, vol. 60, pp , (1952). 21. Hernandez, B. and Wallace, J.F., Mechanism of Pinhole Formation in Gray Iron, AFS Transactions, vol. 87, pp , (1979). 22. Winardi, L., Littleton, H.E. and Bates, C.E., Gas Pressures in Sand Cores, AFS Transactions, vol. 115, paper no , (2007). 23. Caine, J.B. and Toepke, R.E., Gas Pressure and Venting of Cores, AFS Transactions, vol. 74, pp , (1966). 24. Worman, R.A. and Nieman, J.R., A Mathematical System for Exercising Preventive Control over Core gas Defects in Gray Iron Castings, AFS Transactions, vol. 81, pp , (1973). 25. Dietert, H.W., Graham, A.L. and Praski, R.M., Gas Evolution in Foundry Materials Its Source and Measurement, AFS Transactions, vol. 56, pp , (1976). 26. Dietert, H.W., Doelman, R.L. and Bennett, R.W., Mold Surface Gas Pressure, AFS Transactions, vol. 52, pp , (1944).

224 Dietert, H.W., Fairfiled, H.H. and Brewster, F.S., Surface Gas Pressure of Molding Sands and Cores, AFS Transactions, vol. 56, pp , (1948). 28. Marek, C.R. and Ward, C.B., Gas Pressures in Green Sand Molds, Modern Castings, vol. 34, pp , (July 1958). 29. Naro, R.L. and Pelfrey, R.L., Gas Evolution of Synthetic Core Binders: Relationship to Casting Blowhole Defects, AFS Transactions, vol. 107, pp , (1999). 30. Bates. C.E. and Monroe, R.W., Mold Binder Decomposition and Its Relation to Gas Defects in Castings, AFS Transactions, vol. 83, pp , (1981). 31. Penko, T., Measurement of Emission Associated with Application of Flammable Solvent-Based Core and Mold Coatings, AFS Transactions, vol. 108, pp , (2000). 32. Locke. C. and Ashbrook, R.L., Nature of Mold Cavity Gases, AFS Transactions, vol. 58, pp , (1950). 33. Bastic, B.A., Blesch, E.V., Nelson, R.M. and Thomas, P.L., Effect of Head Pressure in measuring Green Sand Permeability, AFS Transactions, vol. 99, p 43, (1991). 34. Adams, T.C., Testing Molding Sand to Determine Their Permeability, AFS Transactions, vol. 32, pp , (1925). 35. Weist, R.C., editor, Chemical Rubber Company Handbook of Chemistry and Physics, 65 th ed., CRC Press Inc., Boca Raton, Florida (1984). 36. Miller, B., Sheldon, D., Griffin, J., Littleton, H. and Bates, C., Technical Report for Precision Lost Foam Casting Technology, phase 2 (1995). 37. Campbell, J., Castings, 1 st ed., pp , Butterworth-Heinemann, Oxford, UK (1991). 38. American Foundry Society, Mold and Core Test Handbook: Lost on Ignition Test, 3 rd ed., pp S, American Foundry Society, Des Plaines, Illinois (2001). 39. Winardi, L., Littleton, H.E., Bates, C.E., New Technique for Measuring Permeability of Cores Made from Various Sands, Binders, Additives, and Coatings, AFS Transactions, vol. 113, pp , (2005). 40. American Foundry Society, Mold and Core Test Handbook: Permeability Test for Mold and Core, 3 rd ed., pp S, American Foundry Society, Des Plaines, Illinois (2001). 41. Alchemcast, Major Quality Issues,

225 210 APPENDIX OBSERVATIONS ON THE EFFECTS OF GAS EVOLUTION RATE ON CORE TEMPERATURE AND BINDER DECOMPOSITION RATE

226 211 Gas Evolution Rate and Temperature Profile The effects of the flow of gas evolved in chemically bonded cores can be observed by examining the temperature profile at different locations and binder contents in the sand specimens. In this case, the temperature profiles of cores made with hotbox and sodium silicate resins are compared. These cores had a cylindrical shape with a diameter of 2.9 cm (1.125 inch) and a length 3.81 cm (1.5 inch length) and immersed into aluminum held at 730 o C (1350 o F). The gas volume and rate from these cores were measured using the technique developed in this study. The gas volumes produced from cores made with 3.5% sodium silicate, 0.7%, and 2.1% hotbox binders were 35, 260, and 460 cm 3, respectively, after 50 seconds metal contact, as illustrated in Figure A-1. The baked sodium silicate cores produced very little gas during contact with molten aluminum. The amount of gas produced was 35 cm 3, which reflected the total amount of air occupying the pores between the sand grains in the core. The gas evolution rate from sodium silicate cores was significantly lower than that from cores containing hydrated water and binder. The maximum gas evolution rates were 1, 7, and 18 cm 3 /s from baked sodium silicate, 0.7%, and 2.1% hotbox cores. The maximum rate occurred about 5 seconds after metal contact or at the first peak for sodium silicate cores. The maximum rates occurred at 29 and 22 seconds after contact for cores containing 0.7% and 2.1% binder, respectively. The temperature profiles at the center of the core cores made with each resin in contact with aluminum at 730 o C (1350 o F) are illustrated in Figure A-2. There were no temperature arrests observed in dried sodium silicate cores because the water content was

227 212 very low and the sand contained no organic materials. There was a temperature arrest at 100 C, however, in the sands bonded with 0.7 and 2.1% hotbox resin, and there was an exothermic reaction near 300 C. This temperature arrests indicated the presence of vapor transport zones in hotbox cores. The size of the vapor transport zone at the center of the core depended on the binder content. The center temperature for cores with 0.7% and 2.1% binder stayed at 100 o C for 9.5 and 16.2 seconds, respectively. The effects of the gas evolution rate on increases in the sand heating rate can be observed before the boiling temperature of water. The temperature at the center of the core containing sodium silicate and 0.7% binder reached 100 o C in 47 and 24.6 seconds, respectively. Cores made with 2.1% binder, which produced more volatile, reached 100 o C only in 13 seconds after metal contact (Figure A-2). Gas Evolution Rate and Binder Decomposition Rate The binder decomposition time and rate were affected by the sand heating rate and, therefore, by the gas evolution rate (Table A-1). The binder decomposed at an earlier time in cores containing higher binder. The binder started to decompose at 71 and 63 seconds in cores containing 0.7% and 2.1% binder at the center of the core as indicated in Table A-1. The binder decomposed faster in cores producing a higher gas evolution rate. The binder decomposition rate at the center of cores containing 0.7% and 2.1% binder was 9 and 23 o C/s/g of binder, respectively. The effect of the gas evolution rate on binder decomposition rate can be observed in the binder burnt profile, as illustrated in Figure A-3. Gray iron melt at 1340 o C (2450 o F) was poured into chill molds (ASTM A W3) made with 1.5% and 3.5%

228 213 phenolic urethane no bake (PUNB) resin. The burnt profile is shown to be more extensive in the molds containing 3.5% PUNB, although the 1.5% mold required less energy to break down the binder. This result highlights the importance of quantifying the effects of gas evolution rate in raising the sand temperature and therefore the binder decomposition rate. Further research in this field should tackle this problem to accurately predict the gas evolution rate from binder decomposition in cores and molds during casting.

229 Rate 2.1% Vol 2.1% Rate (cm 3 /s) Rate NaSiO 3 Rate 0.7% Vol NaSiO 3 Vol 0.7% Volume (cm 3 ) Time (s) 0 Figure A-1. Gas evolution rates and volumes from cores made with 0.7% and 2.1% hotbox binder and baked sodium silicate binder in contact with aluminum at 730 o C (1350 o F). 214

230 % Binder Center Measured 400 Temperature ( o C) % Binder Center Measured Baked Sodium Silicate Center Measured Time (s) Figure A-2. Temperature profile at the center of cores made with dried sodium silicate and 0.7% and 2.1% hotbox binders. The core length and diameter were 1.5 inches (3.81 cm) and inches (2.86 cm), respectively. 215

231 Table A-1. Temperature profile data at the center of cores containing 0.7% and 2.1% hotbox resin. % Binder & Location Time to reach water boilng (s) Time end water boiling (s) Change in time due to water boiling (s) Temp Time binder binder start to start boil (F) boil (s) Temp. binder boil end (F) Change in Time Temp. binder due to boil end binder (s) boil (F) Change in time due to binder boil (s) Time from water stop boil to binder start boil (s) Change in Temp. from water stop boil to binder start boil (F) dt/dt after water boil to binder start boil (F/s) dt/dt After binder boils (F/s) 0.7%Cen %Cen %Cen Avg stdev CV %Cent %Cent %Cent Avg stdev CV

232 Figure A-3. Effects of binder content on binder decomposition rate. 217