{"id": "bd4b589933ef4d7e79c207f4", "input": {"query": "What is the weight of the primary packaging (aseptic bag) for the 10 kg aseptic tomato paste product? Express this as grams of packaging material.", "source_url": "https://www.prontofinefoods.com.au/documents/productpdf/TP381.pdf", "document_text": "GRECI\nINDUSTRIA\nALIMENTARE S.p.a.\nCode 381\nSPECIFICATION\nPOLPACHEF\nQuality\nAssurance\nService\nPRODUCT DESCRIPTION: Prepared from selected varieties of ripe and sound tomatoes\nwhich are roughly diced. It's almost completely removed of seeds and skins by means\nof an extrusion operation. The crushed tomatoes are prepared using only tomatoes\nthat have been cultivated in accordance with the provisions of the Integrated Pest\nManagement specifications of the Emilia-Romagna regions, thanks to which the\nproduct has been certified as CQ \"Controlled Quality\" and \"100% Italian tomato\".\nUSE: Ideal for all preparation which include tomato (pizzas, bruschetta, sauces, and\ntypical Mediterranean recipies).\nINGREDIENTS: 100% Tomato pulp\nORGANOLEPTIC CHARACTERISTICS\nConsistency: creamy\nColour: red, characteristic\nOdour: typical of tomato, without aftertaste\nFlavour: sweet, typical of tomato\nMICROBIOLOGICAL CHARACTERISTICS\nTotal Mesophilic microbic load\n<100 ufc/g\nLactic bacteria\n<1 ufc/g\nYeast\n<1 ufc/g\nMoulds\n<1 ufc/g\nSalmonella\n<1 ufc/g\nListeria monocytogenes\n<1 ufc/g\nS. Aureus\n<1 ufc/g\nBacillus Cereus\n<10 ufc/g\nBacilli ssp\n<100 ufc/g\nE. Coli\n<1 ufc/g\nTotal coliforms\n<1 ufc/g\nGMO: This product does not contain genetically modified organisms.\nDate: 05/05/2014\nVerified: QA\nApproved: QAM\nGRECI\nINDUSTRIA\nSPECIFICATION\nCode 381\nPOLPACHEF\nALIMENTARE S.p.a.\nQuality\nAssurance\nService\nPROCESSING: Following careful sorting, the tomato is washed, blanched, pressed,\ndrained and its seeds and skins are removed by means of an extrusion process. It is\nthen transformed into a puree which, after heating, undergoes thermal treatment\nwhich guarantees the retention of its organoleptic characteristics while at the same\ntime guaranteeing health and safety standards. It is then immediately filled in aseptic\nbag in box in an aseptic ambient.\nPRODUCTS DATA\nPackaging: bag 10/1\nNet weight: 10 kg\nPack size: 1 bag\nSTORAGE DATA\nShelf life: 24 months\nSealed product: keep in a cool and dry place\nOpen product: keep refrigerated\nOnce opened, the product must be consumed within 2-4 days.\nTransport temperature: room temperature\nPRIMARY PACKAGING DESCRIPTION\nPrimary packaging description: multilayer aseptic bag\nHeight: 505.0 mm\nDiameter: 400.0 mm\nPackaging weight: 89.00 gr\nSECONDARY PACKAGING DESCRIPTION\nSecondary packaging description: carton\nLength: 186.0 mm\nWidth: 182.00 mm\nHeight: 317.0 mm\nSecond packaging dimensions. mm: 186*182*317\nWeight of secondary packaging: 250.00 gr\nDate: 05/05/2014\nVerified: QA\nApproved: QAM\nGRECI\nINDUSTRIA\nALIMENTARE S.p.a.\nPALLET COMPOSITION\nCode\n381\nSPECIFICATION\nQuality\nAssurance\nPOLPACHEF\nService\nStandard pallet: 80 X 120\nBoxes number per layer: 24\nLayer numbers per pallet: 3\nBoxes number per pallet: 72\nEAN Code:\nEan 13: 8004980003817\nITF14: 08004980103814\nNUTRIONATIONAL VALUE X 100 G:\nEnergetic value: (kcal): 20.60\nEnergetic value: (KJ): 86.00\nProteins: 1.20 gr\nCarbohydrates: 3.00 gr\nFats: 0.50gr\nFiber: 0.90 gr\nALLERGENS:\nAbsence of allergens: yes\nOTHER CHARACTERISTICS\nSuitable for vegetarians: yes\nSuitable for vegans: yes\nSuitable for celiac: yes\nContains alcohol: no\nContains pork: no\nPHYSICO-CHEMICAL DATA\npH:\nRefractometric residue (\u00b0Brix):\nConsistency (Bostwick):\n4,15-4,35\n5.5 - 7.0\n4,0-6,0 cm\nDate: 05/05/2014\nVerified: QA\nApproved: QAM\n"}, "expected_output": {"claims": [{"unit": "grams", "value": 89, "evidence": ["Packaging weight: 89.00 gr"]}]}, "metadata": {"product_category": "Food & beverages", "request_id": "req_788bd7fc45da0da4"}}
{"id": "e72d0ec2-23f9-4af4-930a-08bd863fa684", "input": {"query": "What is the typical mass percentages of K2O in standard float glass composition? Answer in %.", "source_url": "https://www.pnnl.gov/main/publications/external/technical_reports/PNNL-13958.pdf", "document_text": "PNNL-13958\n\n\nPacific Northwest\nNational Laboratory\nOperated by Battelle for the\nU.S. Department of Energy\n\n\nThermochemical Optimization of\nFloat Glass Composition:\nLow-Alumina Glass Development\n\n\nAugust 2002\n\n\nPrepared for the U.S. Department of Energy\nunder Contract DE-AC06-76RL01830\n\n\nDEPARTMENT\n\n\nENERGY\n\n\n1. UNITED\n\n\nUNITED STA\n\n\nOF AMERIC\n\n\nDSTAT\n\n\nDISCLAIMER\n\n\nThis report was prepared as an account of work sponsored by an agency of the\nUnited States Government. Neither the United States Government nor any\nagency thereof, nor Battelle Memorial Institute, nor any of their employees,\nmakes any warranty, express or implied, or assumes any legal liability or\nresponsibility for the accuracy, completeness, or usefulness of any\ninformation, apparatus, product, or process disclosed, or represents that\nits use would not infringe privately owned rights. Reference herein to any\nspecific commercial product, process, or service by trade name, trademark,\nmanufacturer, or otherwise does not necessarily constitute or imply its\nendorsement, recommendation, or favoring by the United States Government\nor any agency thereof, or Battelle Memorial Institute. The views and opinions\nof authors expressed herein do not necessarily state or reflect those of the\nUnited States Government or any agency thereof.\n\n\nPACIFIC NORTHWEST NATIONAL LABORATORY\n\n\noperated by\nBATTELLE\n\n\nfor the\n\n\nUNITED STATES DEPARTMENT OF ENERGY\n\n\nunder Contract DE-ACO6-76RLO1830\n\n\nPrinted in the United States of America\n\n\nAvailable to DOE and DOE contractors from the\nOffice of Scientific and Technical Information,\nP.O. Box 62, Oak Ridge, TN 37831-0062;\nph: (865) 576-8401\nfax: (865) 576-5728\nemail: reports@adonis.osti.gov\n\n\nAvailable to the public from the National Technical Information Service,\nU.S. Department of Commerce, 5285 Port Royal Rd., Springfield, VA 22161\nph: (800) 553-6847\nfax: (703) 605-6900\nemail: orders@ntis.fedworld.gov\nonline ordering: http://www.ntis.gov/ordering.htm\n\n\nThis document was printed on recycled paper.\n(8/00)\n\n\nPNNL-13958\n\n\nThermochemical Optimization of Float Glass\nComposition: Low-Alumina Glass Development\n\n\nP. R. Hrma\n\n\nD. E. Smith\n\n\nJ. D. Yeager\n\n\nO. P. Lam\n\n\nAugust 2002\n\n\nPrepared for the U.S. Department of Energy\nunder Contract DE-AC06-76RL01830\n\n\nPacific Northwest National Laboratory\n\n\nRichland, Washington 99352\n\n\nAbstract\n\n\nThe liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to\n74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to\n14.2 mass% Na2O. Crystalline phases were identified at 900\u00b0C as cristobalite, wollastonite, and devitrite.\nThe primary phases were tridymite and wollastonite. Partial specific Ts were obtained from the data and\ncompared with the literature.\n\n\nE:\n\n\nContents\n\n\nAbstract....\n\n\n111\n\n\nIntroduction...\n\n\nExperimental..\n\n\n5\n\n\nResults\n\n\n7\n\n\n13\n\n\nDiscussion.\n\n\nConclusions..\n\n\n17\n\n\nReferences...\n\n\n19\n\n\nFigures\n\n\nFigure 1. Map of Tridymite Liquidus Surface in Na\u2082O-CaO-MgO-SiO2 System; Shahid and Glasser\n(1972), reproduced in Roth et al. (1981)..\n\n\n1\n\n\nFigure 2. XRD Pattern of Low-Al2O3-Low-Na2O Glass Heat-Treated at 900\u00b0 C for 24 h.\nFigure 3. Crystals of Cristobalite in Low-Al2O3-Low-CaO Glass, 1020\u00b0C...\nFigure 4. Crystals of Cristobalite in Low-Al2O3-Low-MgO Glass, 1000\u00b0C.\nFigure 5. Crystals of Devitrite in Low-Al2O3-High-CaO Glass, 970\u00b0C\nFigure 6. Crystals of Devitrite in Low-Al2O3-High-MgO Glass, 965\u00b0C\nFigure 7. Crystals of Wollastonite in Low-Al2O3-High-CaO Glass, 995\u00b0C.\nFigure 8. Crystals of Wollastonite in High-Al2O3-High-Na\u2082O Glass, 965\u00b0C.\n\n\n7\n\n\n9\n\n\n9\n\n\n10\n\n\n10\n\n\nFigure 9. Crystals of Wollastonite, Nucleated on a Bubble, in Low-Al2O3-High-MgO Glass, 995\u00b0C... 10\n\n\nFigure 10. Crystals of Wollastonite and Probably Devitrite, Nucleated on a Bubble, in High-Al2O3-\nBaseline Glass, 975\u00b0C\n\n\n10\n\n\nFigure 11. Crystals of Devitrite and Wollastonite in High-Al2O3-High-MgO Glass, 950\u00b0C.\n\n\n11\n\n\nFigure 12. The Effect of CaO on the Maximum Temperature at which Tridymite, Devitrite, and\nWollastonite occur in Float Glass at Equilibrium......\n\n\n13\n\n\nFigure 13. The Effect of MgO on the Maximum Temperature at which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium.....\n\n\n13\n\n\nFigure 14. The Effect of Na\u2082O on the Maximum Temperature at Which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium.\n\n\n14\n\n\nFigure 15. T;, Calculated vs. Measured.\n\n\n14\n\n\nFigure 16. Comparison of Experimental T; Values for VGS Glass with Babcock's TL Values for Soda-\nLime Glass..\n\n\n15\n\n\nFigure 17. Comparison of Experimental T Values for VGS Glass with T Values Predicted by Various\nModels\n\n\n15\n\n\nTables\n\n\nTable 1. Component Coefficients TL in \u00b0C\n\n\n2\n\n\nTable 2. Component Coefficients \u03c0 in \u00b0C in the Nonlinear Model by Backman et al. (a)\nTable 3. Composition Regions for VGS Glasses and for Published Models (mass%)(a).\nTable 4. Test Glass Compositions in Mass% of Oxides\n\n\n3\n\n\n5\n\n\nTable 5. Source Chemicals\n\n\nTable 6. Mass Fractions of Crystalline Phases in Glasses Heat-Treated at 900\u00b0 C for 24 h...\n\n\n8\n\n\nTable 7. Maximum Temperatures (\u00b0C) at Which Individual Crystalline Phases Occurred in VGS\nGlasses (a)(b)\n\n\n8\n\n\nTable 8. Component Coefficients (T;;) for the Maximum Temperatures at which Tridymite, Devitrite,\nWollastonite Occur in Float Glasses at Equilibrium (in \u00b0C)(a).\n\n\n13\n\n\nIntroduction\n\n\nSubstantial savings can be achieved if low-alumina float glass is produced. It has been proposed to\ndecrease the Al2O3 content in the float glass produced by the VGS Glass Systems (VGS) from the original\n0.45 mass% to 0.10 mass%. However, the removal of alumina from the current composition must not\nnegatively impact important glass properties, such as viscosity, liquidus temperature (TL), and chemical\ndurability.\n\n\nFloat glass is essentially a four-component mixture within the SiO2-Na2O-CaO-MgO system with\nminor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O). An\nimportant processability requirement is that the glass must avoid the tridymite primary phase field.\nChemical durability and thermal expansion requirements then place the composition to the neighborhood\nof the devitrite-wollastonite boundary\u2014see Figure 1. When Al2O3 and Fe2O3 are present in the mixture,\nthis boundary shifts towards the Na2O corner of the diagram because Al2O3 and Fe2O3 bound Na\u2082O for\ncharge compensation (this allows these oxides to be glass formers).\n\n\n(CaO)\n\n\n(CaO)\n\n\nSiO2\n\n\n61.4\n\n\n(SiO\u2082-Saturated)}\n\n\n1320\n\n\n-1300-\n\n\n(SiO2-Saturated)\n\n\nB\n\n\n1200-\n\n\n>\n\n\nCaSiO3\n\n\n62.5\n\n\nCaSiO3\n\n\n1375%\n\n\n73 73\n\n\n1035\u00b0\n\n\n---\n\n\n-1100-\n\n\nK\n\n\nDiopside\n\n\nDiopside\n\n\nNa2C03S16016\n\n\n-74\nLNa2CO3Si6016\n\n\n74\n\n\n980\u00b0\n\n\n827\u00b0\n\n\n900\n\n\n74 D\n\n\n75\n\n\n780\u00b0\n11:5\nC77271\nA 710\u00b0\n\n\nProtoenstatite\n\n\nProtoenstatite\n\n\n::5 74\n73 C\n75 75\n11:43:8\n\n\n755\u00b0\n\n\n995\u00b0\n\n\n1:5:12\n77\n\n\n800\u00b0\n\n\nM\n\n\n750\n\n\nNo\u2082Mg2SiO15\n\n\n1:5:12\n\n\nNa2Mg2Si6O15\n\n\n76 75 74.\n\n\n18:4\n\n\n3:8\n\n\n78.0\n\n\n77.0\n\n\n74\n\n\n74\n\n\n74\n\n\n1250\u00b0 1018\u00b0\n\n\n780\u00b0\n\n\n733\u00b0\n\n\n(MgO)\n\n\n(No\u2082O) (MgO)\n\n\n(Na\u2082O)\n\n\nWt %\n\n\nWt%\n\n\nFigure 1. Map of Tridymite Liquidus Surface in Na\u2082O-CaO-MgO-SiO2 System;\nShahid and Glasser (1972), reproduced in Roth et al. (1981)\n\n\nOptimizing glass composition with respect glass properties is done most effectively using glass\nproperty-composition relationship in the form of simple models. For small composition regions, these\nmodels have usually a form of linear functions. Thus, the maximum temperatures (T;) at which j-th\ncrystals are at equilibrium with glass can be expressed as a function of glass composition using the\nfollowing relationship:\n\n\nN\n\n\nT = \u03a3\u03a4\n\n\n(1)\n\n\nj\n\n\ni=1\n\n\n1\n\n\nwhere Tji is the i-th component j-th phase coefficient, x; is the i-th component mass fraction, and N is the\nnumber of components. If N is smaller then the actual number of glass components, mass fractions are\nnormalized in such a way that\n\n\nN\n\n\n\u03a3 x = 1\n\n\n(2)\n\n\ni=1\n\n\nSeveral TL models for commercial glasses exist in the literature. Scholze (1990) reports two such\nmodels: \u0160a\u0161ek et al. (1973) and Cuartas (1984). Other models were presented by Babcock (1977) and\nBackman et al. (1997). Component coefficients for these models are summarized in Tables 1 and 2. Table\n1 lists component coefficients for linear models represented by Equation (1). Models by Babcock and\n\u0160a\u0161ek et al. are linear in terms of mass fractions of components (x;). The original model by Cuartas, which\nis linear in x;/xsi02, was converted into the form of Equation (1) by setting Xsio2 = 0.73. Backman et al.'s\n(1997) model is nonlinear. Table 2 shows only coefficients relevant for the SiO2-Na2O-CaO-MgO-Al2O3\nsystem.\n\n\nTable 1. Component Coefficients TL in \u00b0C\nBabcock (a)\n\n\nTridymite Devitrite Wollastonite \u0160a\u0161ek et al. (a) Cuartas (b)\n\n\nBabcock\n\n\nSiO2\n\n\n2324\n\n\n833\n\n\n938\n\n\n1336\n\n\n1384\n\n\n1604\n\n\nAl2O3\n\n\n-7016\n\n\n2093\n\n\n2533\n\n\n1459\n\n\n304\n\n\n3834\n\n\n1229\n\n\nFe2O3\nCaO\n\n\n8936\n\n\n-342\n\n\n3307\n\n\n4349\n\n\n3061\n\n\n2365\n\n\n2060\n\n\n-1225\n\n\nMgO\n\n\n89\n\n\n-765\n\n\nNa\u2082O -4115\n\n\n-2105\n\n\n-2756\n\n\n-1619\n\n\n69\n\n\nK\u2082O\n\n\n3096\n\n\n-2200\n\n\n(a) Babcock's and \u0160a\u0161ek et al.'s models were transformed to the form of Equation (1); the resulting\n\n\ncoefficients are shown.\n\n\n(b)\n\n\nCuartas' coefficients were adjusted to float glass region with 73 mass% SiO2.\n\n\nTable 2. Component Coefficients T in \u00b0C in the Nonlinear Model by Backman et al.(a)\n\n\n(\u00b0C/mass%)\n\n\n8.90E+02\n\n\nIntercept\nCaO\n\n\n3.07E+01\n\n\n-1.83E+00\n\n\nNa\u2082OxNa\u2082O\n\n\nAl2O3 Na\u2082O\n\n\n2.69E+00\n\n\n2.83E-01\n\n\nNa\u2082OxCaOxMgO\n\n\nCaOxCaOxMgO\u00d7MgO\n\n\n-5.30E-02\n\n\nNa\u2082Ox Na\u2082O \u00d7 Na\u2082O \u00d7 Na\u2082O\n\n\n2.10E-03\n\n\n(a)\n\n\nOnly coefficients applicable to the VGS composition region are shown.\n\n\nUnfortunately, the usefulness of T coefficients that cover multiple primary phase fields is rather\ndubious. Because the effect of glass components on T substantially, even dramatically, changes from one\nprimary phase to another, the global models have a significant lack of fit. Global polynomial forms only\npartly remedy the problem. In addition, polynomial models cannot be extrapolated beyond their\nexperimental domain.\n\n\n2\n\n\nGenerally, it is not recommended to apply models beyond the composition region of their validity.\nComposition regions of the literature models are compared with the float-glass composition region in\nTable 3. Babcock used Silverman's (1939) data and a large database available at Owens-Illinois\nlaboratories. Babcock does not provide the composition region of his Owens-Illinois data. Therefore, only\nthe composition region of Silverman's study is shown in Table 3.\n\n\nTable 3. Composition Regions for VGS Glasses and for Published Models (mass%) (a)\n\n\n\u0160a\u0161ek et al.\n\n\nCuartas\n\n\nBackman\n(1997)\nmin max\n\n\nSilverman\n\n\n(1984)\nmin max\n\n\nExperimental\n\n\n(1973)\n\n\n(1939)\n\n\nFloat\n\n\nmin max\n\n\nmin max\n\n\nmin\n\n\nmin max\n\n\nmax\n\n\n67.5 77.2\n\n\n69.4 76\n\n\nSiO2\n\n\n73.1\n\n\n72.7\n\n\n74\n\n\n64.9\n\n\n50.2\n\n\n68.2\n\n\n59 79\n\n\n0.45\n\n\n0.3\n\n\nAl2O3\n\n\n0.4\n\n\n0.5\n\n\n0.1\n\n\n2\n\n\n0.1\n\n\n4\n\n\n0.5\n\n\n3\n\n\n0.2\n\n\n8.2\n\n\nFe2O3\n\n\n1.5\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0.1\n\n\n0.7\n\n\n0.7\n\n\n0\n\n\n0.5\n\n\n0.9\n\n\n12.2\n\n\n9\n\n\n9\n\n\n6.6\n\n\n14.2\n\n\nCaO\n\n\n7\n\n\n8\n\n\n7\n\n\n8.5\n\n\n2.7\n\n\n6\n\n\n18\n\n\nMgO\n\n\n3\n\n\n2.5\n\n\n4.5\n\n\n3\n\n\n0.1 4.9\n\n\n0.5\n\n\n0.2\n\n\n0.2\n\n\n4\n\n\n4\n\n\n5\n\n\n13.1\n\n\n11.8 14.8\n\n\n14.2\n\n\n11\n\n\nNa\u2082O\n\n\n12\n\n\n8.4 22.9\n\n\n15\n\n\n13 16\n\n\n17\n\n\n2 0.03\n\n\nK\u2082O\n\n\n0\n\n\n0 0.5 0.1 2.3 0.5\n\n\n7\n\n\n0.1\n\n\n(a) Values are highlighted yellow/green where maximum/minimum values of\ncomponent ranges were lower/higher than the corresponding maximum/minimum of\nthe experimental range.\n\n\n3\n\n\nExperimental\n\n\nA test matrix of 14 glasses, shown in Table 4, was developed by the VGS project.\n\n\nTable 4. Test Glass Compositions in Mass % of Oxides\n\n\nHigh Al2O3\n\n\nLow CaO High CaO\n\n\n(mass%)\n\n\nStandard\n73.18\n\n\nLow MgO\n\n\nHigh MgO\n\n\nLow Na\u2082O\n\n\nHigh Na\u2082O\n\n\n73.68\n\n\n73.68\n\n\nSiO2\nAl2O3\nTiO2\nFe2O3\n\n\n72.68\n\n\n72.68\n\n\n73.68\n\n\n72.68\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n8.95\n\n\nCaO\n\n\n8.45\n\n\n7.95\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n2.97\n\n\nMgO\nNa\u2082O\nK\u2082O\n\n\n3.97\n\n\n3.47\n\n\n13.63\n\n\n13.63\n\n\n13.13\n\n\n13.63\n\n\n13.63\n\n\n13.63\n\n\n14.13\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n100.00\n\n\n100.00\n\n\nSum\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\nLow Al2O3\n\n\nLow Na\u2082O\n\n\nHigh CaO\n\n\nHigh Na\u2082O\n\n\n(mass%)\n\n\nLow CaO\n\n\nLow MgO\n\n\nHigh MgO\n\n\nStandard\n\n\n73.53\n\n\n73.03\n\n\n74.03\n\n\n74.03\n\n\nSiO2\n\n\n73.03\n\n\n74.03\n\n\n73.03\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\nAl2O3\nTiO2\nFe2O3\nCaO\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n7.95\n\n\n8.95\n\n\n3.97\n\n\nMgO\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n2.97\n\n\n3.47\n\n\n3.47\n\n\n13.70\n\n\n13.70\n\n\n13.70\n\n\n13.20\n\n\nNa\u2082O\nK\u2082O\nSum\n\n\n13.70\n\n\n13.70\n\n\n14.20\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\nThe test matrix consists of two standard glasses one for high-alumina (0.45 mass%) glass and the\nother for low-alumina (0.10 mass%) glass. Both standards are equal in Al2O3+SiO2 content and in\nNa2O+K\u2082O content (the K2O content is 0.10 mass% in the high-alumina standard glass and 0.03 mass% in\nthe low-alumina standard glass. The remaining glasses are derived from the standard glasses by one-at-a-\ntime changes in CaO, MgO, and Na2O. These changes are compensated for by equal changes in the SiO2\nfraction. The composition regions covered by the VGS test matrix is shown as the experimental region in\nTable 3 and represented by a parallelogram in Figure 1.\n\n\nAs evident form Table 3, \u0160a\u0161ek et al.'s (1973) composition region is broader than the experimental\ncomposition region, except for Al2O3. The Cuartas (1984) composition region is wider than \u0160a\u0161ek's\ncomposition region, except for the upper limit for the range of SiO2. Backman et al. (1997) studied\nglasses with the SiO2 content <68.2 mass%. Their glasses also had at least 0.5 mass% of each of the\nfollowing oxides: Al2O3, K\u2082O, B2O3, SrO, and BaO. Composition regions in these studies covered several\nprimary phase fields, including tridymite, devitrite, wollastonite, and other phases, such as diopside.\n\n\n5\n\n\nBecause the composition variations are small (see bold numbers in Table 4), careful preparation of\nglasses and precise measurements of the T were required. Batches were prepared from chemicals listed\nin Table 5.\n\n\nTable 5. Source Chemicals\n\n\nChemical\n\n\nManufacturer Lot Number\n\n\nAl2O3\n\n\n006627\n\n\nFisher\n\n\nCaCO3\n\n\n007112\n\n\nFisher\n\n\nFisher\n\n\n005612\n\n\nFe2O3\nK2CO3\n\n\n031384\n\n\nAesar\n\n\nAldrich\n\n\n08629JF\n\n\nMgO\n\n\nNa2CO3\n\n\nFisher\n\n\n006825\n\n\nSiO2\n\n\nFisher\n\n\n010455\n\n\nTiO2\n\n\nJ.T. Baker\n\n\n525355\n\n\nThe chemicals were weighed to obtain batches for 250-g glass. Batches were blended, first by hand in\na plastic bag (roll and shake), and then milled in an agate disc mill for 2 min. The glass was melted twice,\neach time in a Pt-10%Rh crucible for 1 h at 1450\u00b0C. After the first melt, the glass was hand crushed and\nthen milled in a tungsten carbide disc mill for 2 min.\n\n\nThe TL was measured in a gradient temperature furnace. Pt boats filled with crushed glass (between\n20 and 40 mesh size) were heat treated for 24 h in the temperature gradient of 1.0\u00b0C/mm, spanning the\ntemperature interval from 800\u00b0C to 1050\u00b0C. Rods of heat-treated glass were mounted in epoxy, and the\nportion exposed to T >936\u00b0C was thin-sectioned for microscopic evaluation.\n\n\nTo identify phases, approximately 2.5-g samples of each glass were heat-treated in a Pt crucible for\n24 h at 900\u00b0C. X-ray diffraction (XRD) was performed on these samples (using Scintag DMC-008) to\ndetect the crystalline phases present and to measure their concentrations. CaF2 (5 mass%) was added to\nthe samples as the standard to evaluate fractions of crystalline phases. Jade software was used for the\nanalysis.\n\n\nResults\n\n\nXRD analysis detected three phases in the samples: cristobalite (SiO2), devitrite (Na2O.3CaO 6SiO2),\nand wollastonite (CaO-SiO2) (Figure 2). Table 6 shows the fractions of these phases in the glasses at\n900\u00b0C. The visual appearances of crystals of these phases were identified using both XRD data and the\nliterature (Bartu\u0161ka 2001). Using the optical microscope (Olympus PMG-3), the maximum temperatures\n(T;) at which j-th crystals were observed in the thin sections were determined (Table 7); typical examples\nare shown in Figures 3 to 11. The Ts (the highest T;) are shown in bold in Table 7.\n\n\nIVGS-13.rd] VGS-13\n\n\n5000-\n\n\n4500-\n\n\n4000-\n\n\n3500-\n\n\nIntensity (Counts)\n\n\n3000-\n\n\n2500-\n\n\n2000-\n\n\n1500-\n\n\n1000-\n\n\n500-\n\n\n0\n\n\n77-2245> CaF2 - Calcium Fluoride\n\n\n39-1425> Cristobalite, syn Si02\n\n\n23-0671> Na2Ca3 Si6016 - Sodium Calcium Silicate\n\n\n27-0088> Wollastonite-2M- CaSiO3\n\n\n10\n\n\n30\n\n\n40\n2-Theta()\n\n\n70\n\n\n20\n\n\n50\n\n\n60\n\n\nPNNL\n\n\n[JADE|h3085125]<c:\\program files Jade WVGS glasses> Thursday, Jul 25, 2002 02:39p (MDI/JADE6)\n\n\nFigure 2. XRD Pattern of Low-Al2O3-Low-Na\u2082O Glass Heat-Treated at 900\u00b0 C for 24 h\n\n\n7\n\n\nTable 6. Mass Fractions of Crystalline Phases in Glasses Heat-Treated at 900\u00b0 C for 24 h\n\n\nCristobalite Devitrite Wollastonite\n\n\nHigh Al2O3\n\n\nStandard\n\n\n0.033\n\n\n0.060\n\n\nLow CaO\nHigh CaO\nLow MgO\n\n\n0.041\n\n\n0.061\n\n\n0.043\n\n\n0.045\n\n\n0.047\n\n\n0.061\n\n\n0.061\n\n\nHigh MgO\n\n\nLow Na\u2082O\n\n\n0.072\n\n\n0.076\n\n\n0.076\n\n\nHigh Na\u2082O\n\n\n0.020\n\n\n0.079\n\n\nLow Al2O3\n\n\nStandard\n\n\n0.052\n\n\n0.066\n\n\nLow CaO\nHigh CaO\nLow MgO\n\n\n0.071\n\n\n0.046\n\n\n0.129\n\n\n0.061\n\n\n0.067\n\n\n0.055\n\n\n0.031\n\n\nHigh MgO\n\n\n0.057\n\n\nLow Na\u2082O\n\n\n0.079\n\n\n0.072\n\n\n0.092\n\n\nHigh Na\u2082O 0.027\n\n\n0.055\n\n\nTable 7. Maximum Temperatures (\u00b0C) at Which Individual Crystalline\nPhases Occurred in VGS Glasses (a)(b)\n\n\nTridymite\n\n\nDevitrite Wollastonite\n\n\nTL\n\n\nHigh Al2O3\n\n\nStandard\n\n\n967\n\n\n992\n\n\n992\n\n\nLow CaO\nHigh CaO\nLow MgO\nHigh MgO\n\n\n952\n\n\n973\n\n\n990\n\n\n990\n\n\n973\n\n\n1011 1011\n\n\n962\n\n\n997\n\n\n980\n\n\n997\n\n\n975\n\n\n1001 1001\n\n\nLow Na\u2082O\n\n\n1011\n\n\n971\n\n\n998 1011\n\n\nHigh Na\u2082O\n\n\n989 989\n\n\n965\n\n\nLow Al2O3\n\n\n995\n\n\nStandard\n\n\n965\n\n\n977 995\n\n\n968\n\n\nLow CaO\nHigh CaO\nLow MgO\nHigh MgO\n\n\n1024\n\n\n968\n\n\n1024\n\n\n1000\n\n\n1000\n\n\n965\n\n\n980 1026\n\n\n970\n\n\n1026\n\n\n965\n\n\n994 994\n\n\n1044(c)\n\n\nLow Na\u2082O\n\n\n962\n\n\n990 1044\n\n\nHigh Na\u2082O\n\n\n979\n\n\n970\n\n\n979\n\n\n(a)936\u00b0C was the minimum temperature at which glasses were evaluated.\n(b)Bold numbers indicate T (see also the last column)\n(1044\u00b0C was the maximum temperature at which glasses were evaluated; repeated tests showed that this was also\nthe TL value for the low-Al2O3-low-Na2O glass.\n\n\n8\n\n\nVGS-9, Cristobalite, 1020\u00b0C 0.010 mm\n\n\n0.010 mm\n\n\nVGS-11, Cristobalite, 1000\u00b0C\n\n\nFigure 4. Crystals of Cristobalite in Low-Al2O3-\nLow-MgO Glass, 1000\u00b0C\n\n\nFigure 3. Crystals of Cristobalite in Low-Al2O3-\nLow-CaO Glass, 1020\u00b0C\n\n\nVGS-10, Devitrite, 970\u00b0C\n\n\nVGS-12, Devitrite, 965\u00b0C\n\n\n0.1 mm\n\n\n0.1 mm\n\n\nFigure 5. Crystals of Devitrite in Low-Al2O3-\nHigh-CaO Glass, 970\u00b0C\n\n\nFigure 6. Crystals of Devitrite in Low-Al2O3-\nHigh-MgO Glass, 965\u00b0C\n\n\n9\n\n\nVGS-7, Wollastonite, 965\u00b0C\n\n\nVGS-10, Wollastonite, 995\u00b0C 0.010 mm\n\n\n0.1 mm\n\n\nFigure 7. Crystals of Wollastonite in Low-Al2O3-\nHigh-CaO Glass, 995\u00b0C\n\n\nFigure 8. Crystals of Wollastonite in High-Al2O3-\nHigh-Na\u2082O Glass, 965\u00b0C\n\n\nVGS-1, Devitrite, 975\u00b0C\n\n\n0.1 mm\n\n\nVGS-12, Wollastonite, 995\u00b0C\n\n\n0.010 mm\n\n\nFigure 10. Crystals of Wollastonite and Probably\nDevitrite, Nucleated on a Bubble, in High-Al2O3-\nBaseline Glass, 975\u00b0C\n\n\nFigure 9. Crystals of Wollastonite, Nucleated on a\nBubble, in Low-Al2O3-High-MgO Glass, 995\u00b0C\n\n\n10\n\n\nVGS-5, Devitrite, 950\u00b0C\n\n\n0.1 mm\n\n\nFigure 11. Crystals of Devitrite and Wollastonite\nin High-Al2O3-High-MgO Glass, 950\u00b0C\n\n\n11\n\n\nDiscussion\n\n\nFigures 12 to 14 show that T; is a nearly linear function of composition. To fit Equation (1) to data,\nmass fractions were normalized to the following five components: SiO2, Na2O, CaO, MgO, and Al2O3.\nThe resulting coefficient values are listed in Table 8. The T; values calculated with these coefficients,\nusing Equation (1), are compared with measured T;s in Figure 15.\n\n\nTable 8. Component Coefficients (T;;) for the Maximum Temperatures at which Tridymite,\nDevitrite, Wollastonite Occur in Float Glasses at Equilibrium (in \u00b0C)(a)\n\n\nDevitrite\n863\n\n\nWollastonite\n\n\nTL\n1669\n\n\nTridymite\n\n\nCCCC\n\n\nSiO2\nAl2O3\n\n\n3033\n\n\n763\n\n\n2787\n\n\n862\n\n\n-1936\n\n\n-7560\n\n\n4235\n2499\n\n\n-2525\n\n\nCaO\n\n\n1756\n\n\n1521\n\n\n-3418\n\n\nMgO\nNa\u2082O\n\n\n1260\n\n\n281\n\n\n-2646\n\n\n-230\n\n\n-6593\n\n\n962\n\n\nR\u00b2\n\n\n0.233\n\n\n0.608\n\n\n0.997\n\n\n0.950\n\n\nadj R\u00b2\n\n\n-0.108\n\n\n0.991\n\n\n0.434\n\n\n0.927\n\n\n3.3\n\n\n1.9\n7\n\n\n6.0\n14\n\n\n13.2\n\n\nS\n\n\n14\n\n\n14\n\n\nn\n\n\n(a) The coefficients are valid for the experimental composition region shown in Table 3.\nNote that the effect of Al2O3 is compounded with the effect of K\u2082O replacing Na\u2082O. The\nsymbols on the two bottom lines: s is the standard error, and n is the number of\nobservations.\n\n\n\u25b2 wollastoniteHiAl\n\n\n\u25b2 wollastonite HiAl\n\n\ndevitrite HiAI\ndevitriteLoAl\n\n\ndevitrite HiA\ndevitrite LoAl\n\n\ntridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\n\u25c6tridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\nMgO Effect\n\n\nCaO Effect\n\n\nA wollastoniteLoA\n\n\nA wollastoniteLoAl\n\n\n1030\n\n\n1030\n\n\n1020\n\n\n1020\n\n\n1010\n\n\n1010\n\n\nTemperature \u00b0C\n\n\nTemperature \u00b0C\n\n\n1000\n\n\n1000\n\n\n990\n\n\n990\n\n\n980\n\n\n980\n\n\n\u0394\n\n\n970\n\n\n970\n\n\n960\n\n\n960\n\n\n950\n\n\n940\n\n\n950\n\n\n7.8\n\n\n8\n\n\n8.2\n\n\n8.4\n\n\n8.6\n\n\n8.8\n\n\n9\n\n\n3.2\n\n\n3.6\n\n\n2.8\n\n\nMgO mass%\n\n\nCaO mass%\n\n\nFigure 13. The Effect of MgO on the Maximum\nTemperature at which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium\n\n\nFigure 12. The Effect of CaO on the Maximum\nTemperature at which Tridymite, Devitrite, and\nWollastonite occur in Float Glass at Equilibrium\n\n\n13\n\n\n\u25c6 tridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\ndevitrite HiAI\ndevitriteLoAI\n\n\n\u25b2 wollastoniteHiAl\n\n\n1060\n\n\nNa\u2082O Effect\n\n\nA wollastoniteLoAl\n\n\n1060\n\n\n1040\n\n\n1040\n\n\nTemperature \u00b0C\n\n\nCalculated T; (\u00b0C)\n\n\n1020\n\n\n1020\n\n\n1000\n\n\nA\n\n\n1000\n\n\n980\n\n\n\u25c6 Tridymite\n\n\n980\n\n\n\u25c9\n\n\n960\n\n\nDevitrite\n\n\n940\n\n\n\u25b2 Wollastonite\n\n\n960\n\n\n13.9\n\n\n14.3\n\n\n13.1\n\n\n13.5\n\n\nNa\u2082O mass%\n\n\n940\n\n\nFigure 14. The Effect of Na\u2082O on the Maximum\nTemperature at Which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium\n\n\n940\n\n\n990\n\n\n1040\n\n\nMeasured T; (\u00b0C)\n\n\nFigure 15. T;, Calculated vs. Measured\n\n\nAt 900\u00b0C, the crystalline form of silica as detected by XRD was cristobalite (Figure 2). At\ntemperatures below 1470\u00b0C, the stable form of silica in soda-lime glasses is tridymite; however, as\nreported by Bartuska (2001), cristobalite occurs in industrial glasses at lower temperatures, alone or\ntogether with tridymite. No attempt was made to distinguish between these two forms in this study.\nHowever, we assume that tridymite was the primary phase.\n\n\nAs Table 7 shows, tridymite was the primary phase in all glasses in which the SiO2 content was >75.3\nmass%. In all other glasses, the primary phase was wollastonite. In these glasses cristobalite did not form\nat T >936\u00b0C, but was detected in some glasses that were heat-treated at 900\u00b0C (Table 6). Devitrite and\nwollastonite were found in all glasses, though wollastonite disappeared at lower temperatures and was not\ndetected at 900\u00b0C in most glasses (Table 6). The highest temperature at which devitrite was observed was\nvirtually independent of composition. It varied within a narrow temperature interval of 950\u00b0C to 975\u00b0C,\nbut these variations in T; cannot be explained in terms of composition variation: note the low R\u00b2 value for\nTdevitrite in Table 8.\n\n\nTable 8 shows that the Tcristobalite increases with SiO2 fraction and decreases with the addition of any\nother component, most notably Al2O3 and Na\u2082O. The formation of wollastonite is promoted mainly by\nCaO, but also by MgO and Al2O3, whereas NaO2 suppresses its formation. The component coefficients\nfor tridymite and wollastonite are much different for each component, showing that the slopes of the\ncorresponding liquidus surfaces are substantially different and even opposite in Al2O3, CaO, and MgO\ncomponents. Therefore, the linear model for T (the last column in Table 8) is not a good representation\nof the real behavior. This is well illustrated in Figures 12 to 14.\n\n\nThe low-Al2O3 standard glass is close to the boundary between the tridymite and wollastonite primary\nphases. Its T is only 3\u00b0C higher than that of the high-Al2O3 standard; this difference is within the\nstandard deviation for Tj; of wollastonite (Table 8). It is, however, possible to decrease low Al2O3 glass TL\nby changing the MgO-to-CaO ratio.\n\n\nFigures 16 and 17 compare T; and T\u00cb measured values with those calculated using different models. It\nis not surprising, as seen in Figure 17, that linear or polynomial models are subjected to a large error.\n\n\n14\n\n\n1200\n\n\n1150\n\n\n1150\n\n\n1100\n\n\n1050\n1100\n650\n\n\nCalculated T; (\u00b0C)\n\n\nCalculated T\u2713 (\u00b0C)\n\n\n1050\n\n\n1000\n\n\n\u25c6 Tridymite\nDevitrite\n\n\nSasek\nCuartas\n\n\n1000\n\n\n950\n\n\n\u25b2 Babcock +\n\u25c6 Backman\n\n\n\u25b2 Wollastonite\n\n\n\u25c7 Babcock Cristobalite\n\n\n950\n\n\nTL\n\n\nBabcock Devitrite\n\n\n900\n\n\nA Babcock Wollastonite\n\n\n850\n\n\n900\n\n\n950 1000 1050 1100 1150 1200\n\n\n850\n\n\n900\n\n\n1050\nMeasured T (\u00b0C)\n\n\n900\n\n\n950\n\n\n1000\n\n\n1100\n\n\n1150\n\n\nMeasured T; (\u00b0C)\n\n\nFigure 17. Comparison of Experimental TL\nValues for VGS Glass with T Values Predicted\nby Various Models\n\n\nFigure 16. Comparison of Experimental T;\nValues for VGS Glass with Babcock's T\u2081 Values\nfor Soda-Lime Glass\n\n\nUsing Babcock's coefficients, T; values were predicted for experimental compositions normalized to\nthe Na\u2082O-CaO-SiO2 system. As seen in Figure 16, Babcock's models overestimate T; values for tridymite\nand underestimate T; values for devitrite and wollastonite. These differences can be attributed to the effect\nof MgO in the experimental glasses.\n\n\nAs a comparison of the coefficients in Table 8 with those in Table 1 shows, Cuartas' coefficients are\ncloser to our empirical coefficients than those by \u0160a\u0161ek et al. (1973). Hence, even though \u0160a\u0161ek et al.'s\n(1973) composition region is closer to our experimental composition region than that of Cuartas (1984),\nT\u2081 values from \u0160a\u0161ek et al.'s model less closely match the experimental Ts (Figure 17).\n\n\nThe main cause of the difference between our Al2O3 coefficients for T and those by other authors is\nprobably the confounding between substituting SiO2 for Al2O3 together with substituting K\u2082O for Na2O.\nOur Al2O3 coefficient accounts for this four-component change and thus cannot be compared with studies\nin which the Al2O3 fraction was an independent variable. Another factor is the extremely narrow range of\nAl2O3 content in our study (0.10 to 0.45 mass% as compared to 0.2 to 8.2 mass% of Silverman [1939]).\n\n\nThe experimental composition region of this study is shown in Figure 1 as a parallelogram. Figure 1\ndisplays isotherms and SiO2 fractions on the SiO2 saturation surface in the Na2O-CaO-MgO-SiO2 system.\nWhen the experimental glasses are normalized to this four-component system, the SiO2 mass fraction\nvaries from 0.737 to 0.747. Glasses with tridymite primary phase are those with a SiO2 mass fraction\nhigher than 0.741 (within the Na\u2082O-CaO-MgO-SiO2 system). This is in excellent agreement with SiO2\ncontours in Figure 1.\n\n\nFigure 1 suggests that glasses outside of the tridymite primary phase field are located in the devitrite\nprimary phase field, whereas the experimental glasses had wollastonite as their primary phase. As\nmentioned earlier, this shift was probably caused by Fe2O3 and Al2O3 in the experimental glass. These\ntwo oxides become network formers by immobilizing Na+ ions. Therefore, the effective Na\u2082O content is\n\n\n15\n\n\nlower in the experimental glasses, which would shift the experimental composition region into the\nwollastonite fields.\n\n\n16\n\n\nConclusions\n\n\nLow-alumina the standard glass has nearly identical TL (995\u00b0C) as the high-alumina the standard\nglass (992\u00b0C). Its primary phase is tridymite, while wollastonite is the primary phase of the latter glass.\nDecreasing SiO2 fraction by as little as 0.3 mass% (by replacing it with CaO, MgO, Na2O, or their\ncombination) shifts the glass into the wollastonite primary phase field. While in this field, the secondary\nphase is devitrite.\n\n\nNeither of the two standard glasses contains wollastonite when equilibrated at 900\u00b0C; both contain\ncristobalite and devitrite, the low-alumina glass at somewhat higher fraction. This, however, should not be\na problem in float glass manufacturing.\n\n\nThe effects of SiO2, CaO, MgO, and Na\u2082O on the T within each primary phase field were expressed\nas partial-specific Ts. These coefficients allow a further optimization on the low-alumina glass.\n\n\nThe partial-specific T coefficients published in the literature do not predict the T values\nsatisfactorily. It appears that two conditions are necessary for a correct prediction of T: 1) separate\nmodels are developed for each primary phase field and 2) the major components of the model are the\nsame as the major components of the composition region of interest.\n\n\n117\n\n\nReferences\n\n\nBabcock CL. 1977. Silicate Glass Technology Methods. John Wiley, New York.\n\n\nBackman R, KH Karlsson, M Cable, and NP Pennington. 1997. \u201cModel for Liquidus Temperature of\nMulti-Component Silicate Glasses.\" Physics and Chemistry of Glasses 38, 103-109.\n\n\nBartu\u0161ka M. 2001. \u201cCrystalline Inclusions.\u201d In Glass Flaws (in Czech), Bartu\u0161ka M, Editor, Pr\u00e1h,\nPrague, Czech Republic.\n\n\nCuartas R. 1984. \"Calculo teorico de propiedades del vidrio: viscosidad, parametros termicos y\nparametros de desvitrificacion.\" Ceram. Vidrio 23, 105-111.\n\n\nRoth RS, T Negas, and LP Cook. 1981. Phase Diagrams for Ceramists. Volume IV. American\nCeramic Society, Columbus, Ohio.\n\n\n\u0160a\u0161ek L, M Bartu\u0161ka, and V Van Thong. 1973. \u201cUtilization of Mathematico-Statistical Methods in\nSilicate Research. 2. Determination of Mathematical Relations for the Calculation of Crystallization\nProperties from Chemical Composition of Sheet and Container Glass\u201d (in Czech). Silikaty 17, 207-217.\nScholze H. 1990. Glass Nature, Structure, and Properties. Springer, New York.\n\n\nShahid KA and FP Glasser. 1972. Physics and Chemistry of Glasses 13, 27 (referenced in Roth et al.\n1981).\n\n\nSilverman WB. 1939. \u201cEffect of Alumina on Devitrification of Soda-Lime-Silica Glasses.\" Journal of\nthe American Ceramics Society 22, 378-384.\n\n\n19\n\n\nPNNL-13958\n\n\nDistribution\n\n\nNo. of\n\n\nNo. of\nCopies\n\n\nCopies\n\n\nOFFSITE\n\n\nONSITE\n\n\n10 Pacific Northwest National Laboratory\n\n\n1 James V. Jones\n\n\nD-S Kim\n\n\nTechnical Fellow\n\n\nK6-24\n\n\nAdvanced Technology and Engineering\n\n\nJ. D. Vienna\n\n\nK6-24\n\n\nVisteon Glass System\n\n\nP. R. Hrma (7)\n\n\nK6-24\n\n\n15000 Commerce Drive\n\n\nM. A. Khaleel\n\n\nK2-18\n\n\nN. Dearborn, MI 48120\n\n\nEdward N. Boulos, Ph.D. (3)\n\n\n3\n\n\nSenior Technical Fellow\n\n\nAdvanced Technology and Engineering\n\n\nVisteon Corporation\nVisteon Glass System\n\n\n15000 Commerce Drive\n\n\nN. Dearborn, MI 48120\n\n\n1\n\n\nDr. Theodore M. Besmann, Head\n\n\nSurface Processing and Mechanics Group\n\n\nOak Ridge National Laboratory\n\n\n1 Bethel Valley Road\n\n\nP.O. Box 2008\n\n\nOak Ridge, TN 37831-6063\n\n\nDistr. 1\n"}, "expected_output": {"claims": [{"unit": "mass%", "value": 0.03, "evidence": ["minor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O).", "0.10", "0.03", "the K2O content is 0.10 mass% in the high-alumina standard glass and 0.03 mass% in\nthe low-alumina standard glass."]}, {"unit": "mass%", "value": 0.1, "evidence": ["minor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O).", "0.10", "0.03", "the K2O content is 0.10 mass% in the high-alumina standard glass and 0.03 mass% in\nthe low-alumina standard glass."]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_7b57f5a0d8068ba3"}}
{"id": "5ea64655-7937-4917-b71f-c02926fad76c", "input": {"query": "What is the typical mass percentages of Fe2O3 in standard float glass composition? Answer in %.", "source_url": "https://www.pnnl.gov/main/publications/external/technical_reports/PNNL-13958.pdf", "document_text": "PNNL-13958\n\n\nPacific Northwest\nNational Laboratory\nOperated by Battelle for the\nU.S. Department of Energy\n\n\nThermochemical Optimization of\nFloat Glass Composition:\nLow-Alumina Glass Development\n\n\nAugust 2002\n\n\nPrepared for the U.S. Department of Energy\nunder Contract DE-AC06-76RL01830\n\n\nDEPARTMENT\n\n\nENERGY\n\n\n1. UNITED\n\n\nUNITED STA\n\n\nOF AMERIC\n\n\nDSTAT\n\n\nDISCLAIMER\n\n\nThis report was prepared as an account of work sponsored by an agency of the\nUnited States Government. Neither the United States Government nor any\nagency thereof, nor Battelle Memorial Institute, nor any of their employees,\nmakes any warranty, express or implied, or assumes any legal liability or\nresponsibility for the accuracy, completeness, or usefulness of any\ninformation, apparatus, product, or process disclosed, or represents that\nits use would not infringe privately owned rights. Reference herein to any\nspecific commercial product, process, or service by trade name, trademark,\nmanufacturer, or otherwise does not necessarily constitute or imply its\nendorsement, recommendation, or favoring by the United States Government\nor any agency thereof, or Battelle Memorial Institute. The views and opinions\nof authors expressed herein do not necessarily state or reflect those of the\nUnited States Government or any agency thereof.\n\n\nPACIFIC NORTHWEST NATIONAL LABORATORY\n\n\noperated by\nBATTELLE\n\n\nfor the\n\n\nUNITED STATES DEPARTMENT OF ENERGY\n\n\nunder Contract DE-ACO6-76RLO1830\n\n\nPrinted in the United States of America\n\n\nAvailable to DOE and DOE contractors from the\nOffice of Scientific and Technical Information,\nP.O. Box 62, Oak Ridge, TN 37831-0062;\nph: (865) 576-8401\nfax: (865) 576-5728\nemail: reports@adonis.osti.gov\n\n\nAvailable to the public from the National Technical Information Service,\nU.S. Department of Commerce, 5285 Port Royal Rd., Springfield, VA 22161\nph: (800) 553-6847\nfax: (703) 605-6900\nemail: orders@ntis.fedworld.gov\nonline ordering: http://www.ntis.gov/ordering.htm\n\n\nThis document was printed on recycled paper.\n(8/00)\n\n\nPNNL-13958\n\n\nThermochemical Optimization of Float Glass\nComposition: Low-Alumina Glass Development\n\n\nP. R. Hrma\n\n\nD. E. Smith\n\n\nJ. D. Yeager\n\n\nO. P. Lam\n\n\nAugust 2002\n\n\nPrepared for the U.S. Department of Energy\nunder Contract DE-AC06-76RL01830\n\n\nPacific Northwest National Laboratory\n\n\nRichland, Washington 99352\n\n\nAbstract\n\n\nThe liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to\n74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to\n14.2 mass% Na2O. Crystalline phases were identified at 900\u00b0C as cristobalite, wollastonite, and devitrite.\nThe primary phases were tridymite and wollastonite. Partial specific Ts were obtained from the data and\ncompared with the literature.\n\n\nE:\n\n\nContents\n\n\nAbstract....\n\n\n111\n\n\nIntroduction...\n\n\nExperimental..\n\n\n5\n\n\nResults\n\n\n7\n\n\n13\n\n\nDiscussion.\n\n\nConclusions..\n\n\n17\n\n\nReferences...\n\n\n19\n\n\nFigures\n\n\nFigure 1. Map of Tridymite Liquidus Surface in Na\u2082O-CaO-MgO-SiO2 System; Shahid and Glasser\n(1972), reproduced in Roth et al. (1981)..\n\n\n1\n\n\nFigure 2. XRD Pattern of Low-Al2O3-Low-Na2O Glass Heat-Treated at 900\u00b0 C for 24 h.\nFigure 3. Crystals of Cristobalite in Low-Al2O3-Low-CaO Glass, 1020\u00b0C...\nFigure 4. Crystals of Cristobalite in Low-Al2O3-Low-MgO Glass, 1000\u00b0C.\nFigure 5. Crystals of Devitrite in Low-Al2O3-High-CaO Glass, 970\u00b0C\nFigure 6. Crystals of Devitrite in Low-Al2O3-High-MgO Glass, 965\u00b0C\nFigure 7. Crystals of Wollastonite in Low-Al2O3-High-CaO Glass, 995\u00b0C.\nFigure 8. Crystals of Wollastonite in High-Al2O3-High-Na\u2082O Glass, 965\u00b0C.\n\n\n7\n\n\n9\n\n\n9\n\n\n10\n\n\n10\n\n\nFigure 9. Crystals of Wollastonite, Nucleated on a Bubble, in Low-Al2O3-High-MgO Glass, 995\u00b0C... 10\n\n\nFigure 10. Crystals of Wollastonite and Probably Devitrite, Nucleated on a Bubble, in High-Al2O3-\nBaseline Glass, 975\u00b0C\n\n\n10\n\n\nFigure 11. Crystals of Devitrite and Wollastonite in High-Al2O3-High-MgO Glass, 950\u00b0C.\n\n\n11\n\n\nFigure 12. The Effect of CaO on the Maximum Temperature at which Tridymite, Devitrite, and\nWollastonite occur in Float Glass at Equilibrium......\n\n\n13\n\n\nFigure 13. The Effect of MgO on the Maximum Temperature at which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium.....\n\n\n13\n\n\nFigure 14. The Effect of Na\u2082O on the Maximum Temperature at Which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium.\n\n\n14\n\n\nFigure 15. T;, Calculated vs. Measured.\n\n\n14\n\n\nFigure 16. Comparison of Experimental T; Values for VGS Glass with Babcock's TL Values for Soda-\nLime Glass..\n\n\n15\n\n\nFigure 17. Comparison of Experimental T Values for VGS Glass with T Values Predicted by Various\nModels\n\n\n15\n\n\nTables\n\n\nTable 1. Component Coefficients TL in \u00b0C\n\n\n2\n\n\nTable 2. Component Coefficients \u03c0 in \u00b0C in the Nonlinear Model by Backman et al. (a)\nTable 3. Composition Regions for VGS Glasses and for Published Models (mass%)(a).\nTable 4. Test Glass Compositions in Mass% of Oxides\n\n\n3\n\n\n5\n\n\nTable 5. Source Chemicals\n\n\nTable 6. Mass Fractions of Crystalline Phases in Glasses Heat-Treated at 900\u00b0 C for 24 h...\n\n\n8\n\n\nTable 7. Maximum Temperatures (\u00b0C) at Which Individual Crystalline Phases Occurred in VGS\nGlasses (a)(b)\n\n\n8\n\n\nTable 8. Component Coefficients (T;;) for the Maximum Temperatures at which Tridymite, Devitrite,\nWollastonite Occur in Float Glasses at Equilibrium (in \u00b0C)(a).\n\n\n13\n\n\nIntroduction\n\n\nSubstantial savings can be achieved if low-alumina float glass is produced. It has been proposed to\ndecrease the Al2O3 content in the float glass produced by the VGS Glass Systems (VGS) from the original\n0.45 mass% to 0.10 mass%. However, the removal of alumina from the current composition must not\nnegatively impact important glass properties, such as viscosity, liquidus temperature (TL), and chemical\ndurability.\n\n\nFloat glass is essentially a four-component mixture within the SiO2-Na2O-CaO-MgO system with\nminor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O). An\nimportant processability requirement is that the glass must avoid the tridymite primary phase field.\nChemical durability and thermal expansion requirements then place the composition to the neighborhood\nof the devitrite-wollastonite boundary\u2014see Figure 1. When Al2O3 and Fe2O3 are present in the mixture,\nthis boundary shifts towards the Na2O corner of the diagram because Al2O3 and Fe2O3 bound Na\u2082O for\ncharge compensation (this allows these oxides to be glass formers).\n\n\n(CaO)\n\n\n(CaO)\n\n\nSiO2\n\n\n61.4\n\n\n(SiO\u2082-Saturated)}\n\n\n1320\n\n\n-1300-\n\n\n(SiO2-Saturated)\n\n\nB\n\n\n1200-\n\n\n>\n\n\nCaSiO3\n\n\n62.5\n\n\nCaSiO3\n\n\n1375%\n\n\n73 73\n\n\n1035\u00b0\n\n\n---\n\n\n-1100-\n\n\nK\n\n\nDiopside\n\n\nDiopside\n\n\nNa2C03S16016\n\n\n-74\nLNa2CO3Si6016\n\n\n74\n\n\n980\u00b0\n\n\n827\u00b0\n\n\n900\n\n\n74 D\n\n\n75\n\n\n780\u00b0\n11:5\nC77271\nA 710\u00b0\n\n\nProtoenstatite\n\n\nProtoenstatite\n\n\n::5 74\n73 C\n75 75\n11:43:8\n\n\n755\u00b0\n\n\n995\u00b0\n\n\n1:5:12\n77\n\n\n800\u00b0\n\n\nM\n\n\n750\n\n\nNo\u2082Mg2SiO15\n\n\n1:5:12\n\n\nNa2Mg2Si6O15\n\n\n76 75 74.\n\n\n18:4\n\n\n3:8\n\n\n78.0\n\n\n77.0\n\n\n74\n\n\n74\n\n\n74\n\n\n1250\u00b0 1018\u00b0\n\n\n780\u00b0\n\n\n733\u00b0\n\n\n(MgO)\n\n\n(No\u2082O) (MgO)\n\n\n(Na\u2082O)\n\n\nWt %\n\n\nWt%\n\n\nFigure 1. Map of Tridymite Liquidus Surface in Na\u2082O-CaO-MgO-SiO2 System;\nShahid and Glasser (1972), reproduced in Roth et al. (1981)\n\n\nOptimizing glass composition with respect glass properties is done most effectively using glass\nproperty-composition relationship in the form of simple models. For small composition regions, these\nmodels have usually a form of linear functions. Thus, the maximum temperatures (T;) at which j-th\ncrystals are at equilibrium with glass can be expressed as a function of glass composition using the\nfollowing relationship:\n\n\nN\n\n\nT = \u03a3\u03a4\n\n\n(1)\n\n\nj\n\n\ni=1\n\n\n1\n\n\nwhere Tji is the i-th component j-th phase coefficient, x; is the i-th component mass fraction, and N is the\nnumber of components. If N is smaller then the actual number of glass components, mass fractions are\nnormalized in such a way that\n\n\nN\n\n\n\u03a3 x = 1\n\n\n(2)\n\n\ni=1\n\n\nSeveral TL models for commercial glasses exist in the literature. Scholze (1990) reports two such\nmodels: \u0160a\u0161ek et al. (1973) and Cuartas (1984). Other models were presented by Babcock (1977) and\nBackman et al. (1997). Component coefficients for these models are summarized in Tables 1 and 2. Table\n1 lists component coefficients for linear models represented by Equation (1). Models by Babcock and\n\u0160a\u0161ek et al. are linear in terms of mass fractions of components (x;). The original model by Cuartas, which\nis linear in x;/xsi02, was converted into the form of Equation (1) by setting Xsio2 = 0.73. Backman et al.'s\n(1997) model is nonlinear. Table 2 shows only coefficients relevant for the SiO2-Na2O-CaO-MgO-Al2O3\nsystem.\n\n\nTable 1. Component Coefficients TL in \u00b0C\nBabcock (a)\n\n\nTridymite Devitrite Wollastonite \u0160a\u0161ek et al. (a) Cuartas (b)\n\n\nBabcock\n\n\nSiO2\n\n\n2324\n\n\n833\n\n\n938\n\n\n1336\n\n\n1384\n\n\n1604\n\n\nAl2O3\n\n\n-7016\n\n\n2093\n\n\n2533\n\n\n1459\n\n\n304\n\n\n3834\n\n\n1229\n\n\nFe2O3\nCaO\n\n\n8936\n\n\n-342\n\n\n3307\n\n\n4349\n\n\n3061\n\n\n2365\n\n\n2060\n\n\n-1225\n\n\nMgO\n\n\n89\n\n\n-765\n\n\nNa\u2082O -4115\n\n\n-2105\n\n\n-2756\n\n\n-1619\n\n\n69\n\n\nK\u2082O\n\n\n3096\n\n\n-2200\n\n\n(a) Babcock's and \u0160a\u0161ek et al.'s models were transformed to the form of Equation (1); the resulting\n\n\ncoefficients are shown.\n\n\n(b)\n\n\nCuartas' coefficients were adjusted to float glass region with 73 mass% SiO2.\n\n\nTable 2. Component Coefficients T in \u00b0C in the Nonlinear Model by Backman et al.(a)\n\n\n(\u00b0C/mass%)\n\n\n8.90E+02\n\n\nIntercept\nCaO\n\n\n3.07E+01\n\n\n-1.83E+00\n\n\nNa\u2082OxNa\u2082O\n\n\nAl2O3 Na\u2082O\n\n\n2.69E+00\n\n\n2.83E-01\n\n\nNa\u2082OxCaOxMgO\n\n\nCaOxCaOxMgO\u00d7MgO\n\n\n-5.30E-02\n\n\nNa\u2082Ox Na\u2082O \u00d7 Na\u2082O \u00d7 Na\u2082O\n\n\n2.10E-03\n\n\n(a)\n\n\nOnly coefficients applicable to the VGS composition region are shown.\n\n\nUnfortunately, the usefulness of T coefficients that cover multiple primary phase fields is rather\ndubious. Because the effect of glass components on T substantially, even dramatically, changes from one\nprimary phase to another, the global models have a significant lack of fit. Global polynomial forms only\npartly remedy the problem. In addition, polynomial models cannot be extrapolated beyond their\nexperimental domain.\n\n\n2\n\n\nGenerally, it is not recommended to apply models beyond the composition region of their validity.\nComposition regions of the literature models are compared with the float-glass composition region in\nTable 3. Babcock used Silverman's (1939) data and a large database available at Owens-Illinois\nlaboratories. Babcock does not provide the composition region of his Owens-Illinois data. Therefore, only\nthe composition region of Silverman's study is shown in Table 3.\n\n\nTable 3. Composition Regions for VGS Glasses and for Published Models (mass%) (a)\n\n\n\u0160a\u0161ek et al.\n\n\nCuartas\n\n\nBackman\n(1997)\nmin max\n\n\nSilverman\n\n\n(1984)\nmin max\n\n\nExperimental\n\n\n(1973)\n\n\n(1939)\n\n\nFloat\n\n\nmin max\n\n\nmin max\n\n\nmin\n\n\nmin max\n\n\nmax\n\n\n67.5 77.2\n\n\n69.4 76\n\n\nSiO2\n\n\n73.1\n\n\n72.7\n\n\n74\n\n\n64.9\n\n\n50.2\n\n\n68.2\n\n\n59 79\n\n\n0.45\n\n\n0.3\n\n\nAl2O3\n\n\n0.4\n\n\n0.5\n\n\n0.1\n\n\n2\n\n\n0.1\n\n\n4\n\n\n0.5\n\n\n3\n\n\n0.2\n\n\n8.2\n\n\nFe2O3\n\n\n1.5\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0.1\n\n\n0.7\n\n\n0.7\n\n\n0\n\n\n0.5\n\n\n0.9\n\n\n12.2\n\n\n9\n\n\n9\n\n\n6.6\n\n\n14.2\n\n\nCaO\n\n\n7\n\n\n8\n\n\n7\n\n\n8.5\n\n\n2.7\n\n\n6\n\n\n18\n\n\nMgO\n\n\n3\n\n\n2.5\n\n\n4.5\n\n\n3\n\n\n0.1 4.9\n\n\n0.5\n\n\n0.2\n\n\n0.2\n\n\n4\n\n\n4\n\n\n5\n\n\n13.1\n\n\n11.8 14.8\n\n\n14.2\n\n\n11\n\n\nNa\u2082O\n\n\n12\n\n\n8.4 22.9\n\n\n15\n\n\n13 16\n\n\n17\n\n\n2 0.03\n\n\nK\u2082O\n\n\n0\n\n\n0 0.5 0.1 2.3 0.5\n\n\n7\n\n\n0.1\n\n\n(a) Values are highlighted yellow/green where maximum/minimum values of\ncomponent ranges were lower/higher than the corresponding maximum/minimum of\nthe experimental range.\n\n\n3\n\n\nExperimental\n\n\nA test matrix of 14 glasses, shown in Table 4, was developed by the VGS project.\n\n\nTable 4. Test Glass Compositions in Mass % of Oxides\n\n\nHigh Al2O3\n\n\nLow CaO High CaO\n\n\n(mass%)\n\n\nStandard\n73.18\n\n\nLow MgO\n\n\nHigh MgO\n\n\nLow Na\u2082O\n\n\nHigh Na\u2082O\n\n\n73.68\n\n\n73.68\n\n\nSiO2\nAl2O3\nTiO2\nFe2O3\n\n\n72.68\n\n\n72.68\n\n\n73.68\n\n\n72.68\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n8.95\n\n\nCaO\n\n\n8.45\n\n\n7.95\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n2.97\n\n\nMgO\nNa\u2082O\nK\u2082O\n\n\n3.97\n\n\n3.47\n\n\n13.63\n\n\n13.63\n\n\n13.13\n\n\n13.63\n\n\n13.63\n\n\n13.63\n\n\n14.13\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n100.00\n\n\n100.00\n\n\nSum\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\nLow Al2O3\n\n\nLow Na\u2082O\n\n\nHigh CaO\n\n\nHigh Na\u2082O\n\n\n(mass%)\n\n\nLow CaO\n\n\nLow MgO\n\n\nHigh MgO\n\n\nStandard\n\n\n73.53\n\n\n73.03\n\n\n74.03\n\n\n74.03\n\n\nSiO2\n\n\n73.03\n\n\n74.03\n\n\n73.03\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\nAl2O3\nTiO2\nFe2O3\nCaO\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n7.95\n\n\n8.95\n\n\n3.97\n\n\nMgO\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n2.97\n\n\n3.47\n\n\n3.47\n\n\n13.70\n\n\n13.70\n\n\n13.70\n\n\n13.20\n\n\nNa\u2082O\nK\u2082O\nSum\n\n\n13.70\n\n\n13.70\n\n\n14.20\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\nThe test matrix consists of two standard glasses one for high-alumina (0.45 mass%) glass and the\nother for low-alumina (0.10 mass%) glass. Both standards are equal in Al2O3+SiO2 content and in\nNa2O+K\u2082O content (the K2O content is 0.10 mass% in the high-alumina standard glass and 0.03 mass% in\nthe low-alumina standard glass. The remaining glasses are derived from the standard glasses by one-at-a-\ntime changes in CaO, MgO, and Na2O. These changes are compensated for by equal changes in the SiO2\nfraction. The composition regions covered by the VGS test matrix is shown as the experimental region in\nTable 3 and represented by a parallelogram in Figure 1.\n\n\nAs evident form Table 3, \u0160a\u0161ek et al.'s (1973) composition region is broader than the experimental\ncomposition region, except for Al2O3. The Cuartas (1984) composition region is wider than \u0160a\u0161ek's\ncomposition region, except for the upper limit for the range of SiO2. Backman et al. (1997) studied\nglasses with the SiO2 content <68.2 mass%. Their glasses also had at least 0.5 mass% of each of the\nfollowing oxides: Al2O3, K\u2082O, B2O3, SrO, and BaO. Composition regions in these studies covered several\nprimary phase fields, including tridymite, devitrite, wollastonite, and other phases, such as diopside.\n\n\n5\n\n\nBecause the composition variations are small (see bold numbers in Table 4), careful preparation of\nglasses and precise measurements of the T were required. Batches were prepared from chemicals listed\nin Table 5.\n\n\nTable 5. Source Chemicals\n\n\nChemical\n\n\nManufacturer Lot Number\n\n\nAl2O3\n\n\n006627\n\n\nFisher\n\n\nCaCO3\n\n\n007112\n\n\nFisher\n\n\nFisher\n\n\n005612\n\n\nFe2O3\nK2CO3\n\n\n031384\n\n\nAesar\n\n\nAldrich\n\n\n08629JF\n\n\nMgO\n\n\nNa2CO3\n\n\nFisher\n\n\n006825\n\n\nSiO2\n\n\nFisher\n\n\n010455\n\n\nTiO2\n\n\nJ.T. Baker\n\n\n525355\n\n\nThe chemicals were weighed to obtain batches for 250-g glass. Batches were blended, first by hand in\na plastic bag (roll and shake), and then milled in an agate disc mill for 2 min. The glass was melted twice,\neach time in a Pt-10%Rh crucible for 1 h at 1450\u00b0C. After the first melt, the glass was hand crushed and\nthen milled in a tungsten carbide disc mill for 2 min.\n\n\nThe TL was measured in a gradient temperature furnace. Pt boats filled with crushed glass (between\n20 and 40 mesh size) were heat treated for 24 h in the temperature gradient of 1.0\u00b0C/mm, spanning the\ntemperature interval from 800\u00b0C to 1050\u00b0C. Rods of heat-treated glass were mounted in epoxy, and the\nportion exposed to T >936\u00b0C was thin-sectioned for microscopic evaluation.\n\n\nTo identify phases, approximately 2.5-g samples of each glass were heat-treated in a Pt crucible for\n24 h at 900\u00b0C. X-ray diffraction (XRD) was performed on these samples (using Scintag DMC-008) to\ndetect the crystalline phases present and to measure their concentrations. CaF2 (5 mass%) was added to\nthe samples as the standard to evaluate fractions of crystalline phases. Jade software was used for the\nanalysis.\n\n\nResults\n\n\nXRD analysis detected three phases in the samples: cristobalite (SiO2), devitrite (Na2O.3CaO 6SiO2),\nand wollastonite (CaO-SiO2) (Figure 2). Table 6 shows the fractions of these phases in the glasses at\n900\u00b0C. The visual appearances of crystals of these phases were identified using both XRD data and the\nliterature (Bartu\u0161ka 2001). Using the optical microscope (Olympus PMG-3), the maximum temperatures\n(T;) at which j-th crystals were observed in the thin sections were determined (Table 7); typical examples\nare shown in Figures 3 to 11. The Ts (the highest T;) are shown in bold in Table 7.\n\n\nIVGS-13.rd] VGS-13\n\n\n5000-\n\n\n4500-\n\n\n4000-\n\n\n3500-\n\n\nIntensity (Counts)\n\n\n3000-\n\n\n2500-\n\n\n2000-\n\n\n1500-\n\n\n1000-\n\n\n500-\n\n\n0\n\n\n77-2245> CaF2 - Calcium Fluoride\n\n\n39-1425> Cristobalite, syn Si02\n\n\n23-0671> Na2Ca3 Si6016 - Sodium Calcium Silicate\n\n\n27-0088> Wollastonite-2M- CaSiO3\n\n\n10\n\n\n30\n\n\n40\n2-Theta()\n\n\n70\n\n\n20\n\n\n50\n\n\n60\n\n\nPNNL\n\n\n[JADE|h3085125]<c:\\program files Jade WVGS glasses> Thursday, Jul 25, 2002 02:39p (MDI/JADE6)\n\n\nFigure 2. XRD Pattern of Low-Al2O3-Low-Na\u2082O Glass Heat-Treated at 900\u00b0 C for 24 h\n\n\n7\n\n\nTable 6. Mass Fractions of Crystalline Phases in Glasses Heat-Treated at 900\u00b0 C for 24 h\n\n\nCristobalite Devitrite Wollastonite\n\n\nHigh Al2O3\n\n\nStandard\n\n\n0.033\n\n\n0.060\n\n\nLow CaO\nHigh CaO\nLow MgO\n\n\n0.041\n\n\n0.061\n\n\n0.043\n\n\n0.045\n\n\n0.047\n\n\n0.061\n\n\n0.061\n\n\nHigh MgO\n\n\nLow Na\u2082O\n\n\n0.072\n\n\n0.076\n\n\n0.076\n\n\nHigh Na\u2082O\n\n\n0.020\n\n\n0.079\n\n\nLow Al2O3\n\n\nStandard\n\n\n0.052\n\n\n0.066\n\n\nLow CaO\nHigh CaO\nLow MgO\n\n\n0.071\n\n\n0.046\n\n\n0.129\n\n\n0.061\n\n\n0.067\n\n\n0.055\n\n\n0.031\n\n\nHigh MgO\n\n\n0.057\n\n\nLow Na\u2082O\n\n\n0.079\n\n\n0.072\n\n\n0.092\n\n\nHigh Na\u2082O 0.027\n\n\n0.055\n\n\nTable 7. Maximum Temperatures (\u00b0C) at Which Individual Crystalline\nPhases Occurred in VGS Glasses (a)(b)\n\n\nTridymite\n\n\nDevitrite Wollastonite\n\n\nTL\n\n\nHigh Al2O3\n\n\nStandard\n\n\n967\n\n\n992\n\n\n992\n\n\nLow CaO\nHigh CaO\nLow MgO\nHigh MgO\n\n\n952\n\n\n973\n\n\n990\n\n\n990\n\n\n973\n\n\n1011 1011\n\n\n962\n\n\n997\n\n\n980\n\n\n997\n\n\n975\n\n\n1001 1001\n\n\nLow Na\u2082O\n\n\n1011\n\n\n971\n\n\n998 1011\n\n\nHigh Na\u2082O\n\n\n989 989\n\n\n965\n\n\nLow Al2O3\n\n\n995\n\n\nStandard\n\n\n965\n\n\n977 995\n\n\n968\n\n\nLow CaO\nHigh CaO\nLow MgO\nHigh MgO\n\n\n1024\n\n\n968\n\n\n1024\n\n\n1000\n\n\n1000\n\n\n965\n\n\n980 1026\n\n\n970\n\n\n1026\n\n\n965\n\n\n994 994\n\n\n1044(c)\n\n\nLow Na\u2082O\n\n\n962\n\n\n990 1044\n\n\nHigh Na\u2082O\n\n\n979\n\n\n970\n\n\n979\n\n\n(a)936\u00b0C was the minimum temperature at which glasses were evaluated.\n(b)Bold numbers indicate T (see also the last column)\n(1044\u00b0C was the maximum temperature at which glasses were evaluated; repeated tests showed that this was also\nthe TL value for the low-Al2O3-low-Na2O glass.\n\n\n8\n\n\nVGS-9, Cristobalite, 1020\u00b0C 0.010 mm\n\n\n0.010 mm\n\n\nVGS-11, Cristobalite, 1000\u00b0C\n\n\nFigure 4. Crystals of Cristobalite in Low-Al2O3-\nLow-MgO Glass, 1000\u00b0C\n\n\nFigure 3. Crystals of Cristobalite in Low-Al2O3-\nLow-CaO Glass, 1020\u00b0C\n\n\nVGS-10, Devitrite, 970\u00b0C\n\n\nVGS-12, Devitrite, 965\u00b0C\n\n\n0.1 mm\n\n\n0.1 mm\n\n\nFigure 5. Crystals of Devitrite in Low-Al2O3-\nHigh-CaO Glass, 970\u00b0C\n\n\nFigure 6. Crystals of Devitrite in Low-Al2O3-\nHigh-MgO Glass, 965\u00b0C\n\n\n9\n\n\nVGS-7, Wollastonite, 965\u00b0C\n\n\nVGS-10, Wollastonite, 995\u00b0C 0.010 mm\n\n\n0.1 mm\n\n\nFigure 7. Crystals of Wollastonite in Low-Al2O3-\nHigh-CaO Glass, 995\u00b0C\n\n\nFigure 8. Crystals of Wollastonite in High-Al2O3-\nHigh-Na\u2082O Glass, 965\u00b0C\n\n\nVGS-1, Devitrite, 975\u00b0C\n\n\n0.1 mm\n\n\nVGS-12, Wollastonite, 995\u00b0C\n\n\n0.010 mm\n\n\nFigure 10. Crystals of Wollastonite and Probably\nDevitrite, Nucleated on a Bubble, in High-Al2O3-\nBaseline Glass, 975\u00b0C\n\n\nFigure 9. Crystals of Wollastonite, Nucleated on a\nBubble, in Low-Al2O3-High-MgO Glass, 995\u00b0C\n\n\n10\n\n\nVGS-5, Devitrite, 950\u00b0C\n\n\n0.1 mm\n\n\nFigure 11. Crystals of Devitrite and Wollastonite\nin High-Al2O3-High-MgO Glass, 950\u00b0C\n\n\n11\n\n\nDiscussion\n\n\nFigures 12 to 14 show that T; is a nearly linear function of composition. To fit Equation (1) to data,\nmass fractions were normalized to the following five components: SiO2, Na2O, CaO, MgO, and Al2O3.\nThe resulting coefficient values are listed in Table 8. The T; values calculated with these coefficients,\nusing Equation (1), are compared with measured T;s in Figure 15.\n\n\nTable 8. Component Coefficients (T;;) for the Maximum Temperatures at which Tridymite,\nDevitrite, Wollastonite Occur in Float Glasses at Equilibrium (in \u00b0C)(a)\n\n\nDevitrite\n863\n\n\nWollastonite\n\n\nTL\n1669\n\n\nTridymite\n\n\nCCCC\n\n\nSiO2\nAl2O3\n\n\n3033\n\n\n763\n\n\n2787\n\n\n862\n\n\n-1936\n\n\n-7560\n\n\n4235\n2499\n\n\n-2525\n\n\nCaO\n\n\n1756\n\n\n1521\n\n\n-3418\n\n\nMgO\nNa\u2082O\n\n\n1260\n\n\n281\n\n\n-2646\n\n\n-230\n\n\n-6593\n\n\n962\n\n\nR\u00b2\n\n\n0.233\n\n\n0.608\n\n\n0.997\n\n\n0.950\n\n\nadj R\u00b2\n\n\n-0.108\n\n\n0.991\n\n\n0.434\n\n\n0.927\n\n\n3.3\n\n\n1.9\n7\n\n\n6.0\n14\n\n\n13.2\n\n\nS\n\n\n14\n\n\n14\n\n\nn\n\n\n(a) The coefficients are valid for the experimental composition region shown in Table 3.\nNote that the effect of Al2O3 is compounded with the effect of K\u2082O replacing Na\u2082O. The\nsymbols on the two bottom lines: s is the standard error, and n is the number of\nobservations.\n\n\n\u25b2 wollastoniteHiAl\n\n\n\u25b2 wollastonite HiAl\n\n\ndevitrite HiAI\ndevitriteLoAl\n\n\ndevitrite HiA\ndevitrite LoAl\n\n\ntridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\n\u25c6tridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\nMgO Effect\n\n\nCaO Effect\n\n\nA wollastoniteLoA\n\n\nA wollastoniteLoAl\n\n\n1030\n\n\n1030\n\n\n1020\n\n\n1020\n\n\n1010\n\n\n1010\n\n\nTemperature \u00b0C\n\n\nTemperature \u00b0C\n\n\n1000\n\n\n1000\n\n\n990\n\n\n990\n\n\n980\n\n\n980\n\n\n\u0394\n\n\n970\n\n\n970\n\n\n960\n\n\n960\n\n\n950\n\n\n940\n\n\n950\n\n\n7.8\n\n\n8\n\n\n8.2\n\n\n8.4\n\n\n8.6\n\n\n8.8\n\n\n9\n\n\n3.2\n\n\n3.6\n\n\n2.8\n\n\nMgO mass%\n\n\nCaO mass%\n\n\nFigure 13. The Effect of MgO on the Maximum\nTemperature at which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium\n\n\nFigure 12. The Effect of CaO on the Maximum\nTemperature at which Tridymite, Devitrite, and\nWollastonite occur in Float Glass at Equilibrium\n\n\n13\n\n\n\u25c6 tridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\ndevitrite HiAI\ndevitriteLoAI\n\n\n\u25b2 wollastoniteHiAl\n\n\n1060\n\n\nNa\u2082O Effect\n\n\nA wollastoniteLoAl\n\n\n1060\n\n\n1040\n\n\n1040\n\n\nTemperature \u00b0C\n\n\nCalculated T; (\u00b0C)\n\n\n1020\n\n\n1020\n\n\n1000\n\n\nA\n\n\n1000\n\n\n980\n\n\n\u25c6 Tridymite\n\n\n980\n\n\n\u25c9\n\n\n960\n\n\nDevitrite\n\n\n940\n\n\n\u25b2 Wollastonite\n\n\n960\n\n\n13.9\n\n\n14.3\n\n\n13.1\n\n\n13.5\n\n\nNa\u2082O mass%\n\n\n940\n\n\nFigure 14. The Effect of Na\u2082O on the Maximum\nTemperature at Which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium\n\n\n940\n\n\n990\n\n\n1040\n\n\nMeasured T; (\u00b0C)\n\n\nFigure 15. T;, Calculated vs. Measured\n\n\nAt 900\u00b0C, the crystalline form of silica as detected by XRD was cristobalite (Figure 2). At\ntemperatures below 1470\u00b0C, the stable form of silica in soda-lime glasses is tridymite; however, as\nreported by Bartuska (2001), cristobalite occurs in industrial glasses at lower temperatures, alone or\ntogether with tridymite. No attempt was made to distinguish between these two forms in this study.\nHowever, we assume that tridymite was the primary phase.\n\n\nAs Table 7 shows, tridymite was the primary phase in all glasses in which the SiO2 content was >75.3\nmass%. In all other glasses, the primary phase was wollastonite. In these glasses cristobalite did not form\nat T >936\u00b0C, but was detected in some glasses that were heat-treated at 900\u00b0C (Table 6). Devitrite and\nwollastonite were found in all glasses, though wollastonite disappeared at lower temperatures and was not\ndetected at 900\u00b0C in most glasses (Table 6). The highest temperature at which devitrite was observed was\nvirtually independent of composition. It varied within a narrow temperature interval of 950\u00b0C to 975\u00b0C,\nbut these variations in T; cannot be explained in terms of composition variation: note the low R\u00b2 value for\nTdevitrite in Table 8.\n\n\nTable 8 shows that the Tcristobalite increases with SiO2 fraction and decreases with the addition of any\nother component, most notably Al2O3 and Na\u2082O. The formation of wollastonite is promoted mainly by\nCaO, but also by MgO and Al2O3, whereas NaO2 suppresses its formation. The component coefficients\nfor tridymite and wollastonite are much different for each component, showing that the slopes of the\ncorresponding liquidus surfaces are substantially different and even opposite in Al2O3, CaO, and MgO\ncomponents. Therefore, the linear model for T (the last column in Table 8) is not a good representation\nof the real behavior. This is well illustrated in Figures 12 to 14.\n\n\nThe low-Al2O3 standard glass is close to the boundary between the tridymite and wollastonite primary\nphases. Its T is only 3\u00b0C higher than that of the high-Al2O3 standard; this difference is within the\nstandard deviation for Tj; of wollastonite (Table 8). It is, however, possible to decrease low Al2O3 glass TL\nby changing the MgO-to-CaO ratio.\n\n\nFigures 16 and 17 compare T; and T\u00cb measured values with those calculated using different models. It\nis not surprising, as seen in Figure 17, that linear or polynomial models are subjected to a large error.\n\n\n14\n\n\n1200\n\n\n1150\n\n\n1150\n\n\n1100\n\n\n1050\n1100\n650\n\n\nCalculated T; (\u00b0C)\n\n\nCalculated T\u2713 (\u00b0C)\n\n\n1050\n\n\n1000\n\n\n\u25c6 Tridymite\nDevitrite\n\n\nSasek\nCuartas\n\n\n1000\n\n\n950\n\n\n\u25b2 Babcock +\n\u25c6 Backman\n\n\n\u25b2 Wollastonite\n\n\n\u25c7 Babcock Cristobalite\n\n\n950\n\n\nTL\n\n\nBabcock Devitrite\n\n\n900\n\n\nA Babcock Wollastonite\n\n\n850\n\n\n900\n\n\n950 1000 1050 1100 1150 1200\n\n\n850\n\n\n900\n\n\n1050\nMeasured T (\u00b0C)\n\n\n900\n\n\n950\n\n\n1000\n\n\n1100\n\n\n1150\n\n\nMeasured T; (\u00b0C)\n\n\nFigure 17. Comparison of Experimental TL\nValues for VGS Glass with T Values Predicted\nby Various Models\n\n\nFigure 16. Comparison of Experimental T;\nValues for VGS Glass with Babcock's T\u2081 Values\nfor Soda-Lime Glass\n\n\nUsing Babcock's coefficients, T; values were predicted for experimental compositions normalized to\nthe Na\u2082O-CaO-SiO2 system. As seen in Figure 16, Babcock's models overestimate T; values for tridymite\nand underestimate T; values for devitrite and wollastonite. These differences can be attributed to the effect\nof MgO in the experimental glasses.\n\n\nAs a comparison of the coefficients in Table 8 with those in Table 1 shows, Cuartas' coefficients are\ncloser to our empirical coefficients than those by \u0160a\u0161ek et al. (1973). Hence, even though \u0160a\u0161ek et al.'s\n(1973) composition region is closer to our experimental composition region than that of Cuartas (1984),\nT\u2081 values from \u0160a\u0161ek et al.'s model less closely match the experimental Ts (Figure 17).\n\n\nThe main cause of the difference between our Al2O3 coefficients for T and those by other authors is\nprobably the confounding between substituting SiO2 for Al2O3 together with substituting K\u2082O for Na2O.\nOur Al2O3 coefficient accounts for this four-component change and thus cannot be compared with studies\nin which the Al2O3 fraction was an independent variable. Another factor is the extremely narrow range of\nAl2O3 content in our study (0.10 to 0.45 mass% as compared to 0.2 to 8.2 mass% of Silverman [1939]).\n\n\nThe experimental composition region of this study is shown in Figure 1 as a parallelogram. Figure 1\ndisplays isotherms and SiO2 fractions on the SiO2 saturation surface in the Na2O-CaO-MgO-SiO2 system.\nWhen the experimental glasses are normalized to this four-component system, the SiO2 mass fraction\nvaries from 0.737 to 0.747. Glasses with tridymite primary phase are those with a SiO2 mass fraction\nhigher than 0.741 (within the Na\u2082O-CaO-MgO-SiO2 system). This is in excellent agreement with SiO2\ncontours in Figure 1.\n\n\nFigure 1 suggests that glasses outside of the tridymite primary phase field are located in the devitrite\nprimary phase field, whereas the experimental glasses had wollastonite as their primary phase. As\nmentioned earlier, this shift was probably caused by Fe2O3 and Al2O3 in the experimental glass. These\ntwo oxides become network formers by immobilizing Na+ ions. Therefore, the effective Na\u2082O content is\n\n\n15\n\n\nlower in the experimental glasses, which would shift the experimental composition region into the\nwollastonite fields.\n\n\n16\n\n\nConclusions\n\n\nLow-alumina the standard glass has nearly identical TL (995\u00b0C) as the high-alumina the standard\nglass (992\u00b0C). Its primary phase is tridymite, while wollastonite is the primary phase of the latter glass.\nDecreasing SiO2 fraction by as little as 0.3 mass% (by replacing it with CaO, MgO, Na2O, or their\ncombination) shifts the glass into the wollastonite primary phase field. While in this field, the secondary\nphase is devitrite.\n\n\nNeither of the two standard glasses contains wollastonite when equilibrated at 900\u00b0C; both contain\ncristobalite and devitrite, the low-alumina glass at somewhat higher fraction. This, however, should not be\na problem in float glass manufacturing.\n\n\nThe effects of SiO2, CaO, MgO, and Na\u2082O on the T within each primary phase field were expressed\nas partial-specific Ts. These coefficients allow a further optimization on the low-alumina glass.\n\n\nThe partial-specific T coefficients published in the literature do not predict the T values\nsatisfactorily. It appears that two conditions are necessary for a correct prediction of T: 1) separate\nmodels are developed for each primary phase field and 2) the major components of the model are the\nsame as the major components of the composition region of interest.\n\n\n117\n\n\nReferences\n\n\nBabcock CL. 1977. Silicate Glass Technology Methods. John Wiley, New York.\n\n\nBackman R, KH Karlsson, M Cable, and NP Pennington. 1997. \u201cModel for Liquidus Temperature of\nMulti-Component Silicate Glasses.\" Physics and Chemistry of Glasses 38, 103-109.\n\n\nBartu\u0161ka M. 2001. \u201cCrystalline Inclusions.\u201d In Glass Flaws (in Czech), Bartu\u0161ka M, Editor, Pr\u00e1h,\nPrague, Czech Republic.\n\n\nCuartas R. 1984. \"Calculo teorico de propiedades del vidrio: viscosidad, parametros termicos y\nparametros de desvitrificacion.\" Ceram. Vidrio 23, 105-111.\n\n\nRoth RS, T Negas, and LP Cook. 1981. Phase Diagrams for Ceramists. Volume IV. American\nCeramic Society, Columbus, Ohio.\n\n\n\u0160a\u0161ek L, M Bartu\u0161ka, and V Van Thong. 1973. \u201cUtilization of Mathematico-Statistical Methods in\nSilicate Research. 2. Determination of Mathematical Relations for the Calculation of Crystallization\nProperties from Chemical Composition of Sheet and Container Glass\u201d (in Czech). Silikaty 17, 207-217.\nScholze H. 1990. Glass Nature, Structure, and Properties. Springer, New York.\n\n\nShahid KA and FP Glasser. 1972. Physics and Chemistry of Glasses 13, 27 (referenced in Roth et al.\n1981).\n\n\nSilverman WB. 1939. \u201cEffect of Alumina on Devitrification of Soda-Lime-Silica Glasses.\" Journal of\nthe American Ceramics Society 22, 378-384.\n\n\n19\n\n\nPNNL-13958\n\n\nDistribution\n\n\nNo. of\n\n\nNo. of\nCopies\n\n\nCopies\n\n\nOFFSITE\n\n\nONSITE\n\n\n10 Pacific Northwest National Laboratory\n\n\n1 James V. Jones\n\n\nD-S Kim\n\n\nTechnical Fellow\n\n\nK6-24\n\n\nAdvanced Technology and Engineering\n\n\nJ. D. Vienna\n\n\nK6-24\n\n\nVisteon Glass System\n\n\nP. R. Hrma (7)\n\n\nK6-24\n\n\n15000 Commerce Drive\n\n\nM. A. Khaleel\n\n\nK2-18\n\n\nN. Dearborn, MI 48120\n\n\nEdward N. Boulos, Ph.D. (3)\n\n\n3\n\n\nSenior Technical Fellow\n\n\nAdvanced Technology and Engineering\n\n\nVisteon Corporation\nVisteon Glass System\n\n\n15000 Commerce Drive\n\n\nN. Dearborn, MI 48120\n\n\n1\n\n\nDr. Theodore M. Besmann, Head\n\n\nSurface Processing and Mechanics Group\n\n\nOak Ridge National Laboratory\n\n\n1 Bethel Valley Road\n\n\nP.O. Box 2008\n\n\nOak Ridge, TN 37831-6063\n\n\nDistr. 1\n"}, "expected_output": {"claims": [{"unit": "%", "value": 0.71, "evidence": ["Float glass is essentially a four-component mixture within the SiO2-Na2O-CaO-MgO system with minor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O).", "0.71", "minor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O).", "Fe2O3"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_7b57f5a0d8068ba3"}}
{"id": "82a464c5-7de3-4b71-b562-0f787f1957c8", "input": {"query": "What is the exact material composition (mass fractions or percentages) of PVB resin in Trosifol PVB film? Provide the specific values as mass fractions or percentages.", "source_url": "https://www.trosifol.com/fileadmin/user_upload/about_us/sustainability/environmental-product-decleration-as-per-iso-14025-and-en-15804_a2-pvb-film-trosifol.pdf", "document_text": "ENVIRONMENTAL PRODUCT DECLARATION\n\n\nas per ISO 14025 and EN 15804+A2\n\n\nOwner of the Declaration Kuraray Europe GmbH\n\n\nInstitut Bauen und Umwelt e.V. (IBU)\n\n\nPublisher\n\n\nInstitut Bauen und Umwelt e.V. (IBU)\n\n\nProgramme holder\n\n\nEPD-KUR-20230072-CCI1-EN\n\n\nDeclaration number\n\n\nIssue date\n\n\n12/05/2023\n\n\nValid to\n\n\n11/05/2028\n\n\nPVB film (Trosifol\u24c7)\nKuraray Europe GmbH\n\n\nInstitut Bauen\nund Umwelt e.V.\n\n\nwww.ibu-epd.com | https://epd-online.com\n\n\nECO PLATFORM\n\n\nEPD\n\n\nVERIFIED\n\n\nALIEK\n\n\nkuraray\n\n\nGeneral Information\n\n\nKuraray Europe GmbH\nProgramme holder\n\n\nPVB film (Trosifol\u24c7)\n\n\nOwner of the declaration\nKuraray Europe GmbH\nPhilipp-Reis-Str. 4\n\n\nIBU - Institut Bauen und Umwelt e.V.\n\n\nHegelplatz 1\n10117 Berlin\n\n\n65795 Hattersheim\nGermany\n\n\nGermany\n\n\nDeclaration number\n\n\nDeclared product / declared unit\n\n\nTrosifol\u00ae PVB.\n\n\nEPD-KUR-20230072-CCI1-EN\n\n\nThe declared unit is 1 m\u00b2.\n\n\nThis declaration is based on the product category rules:\n\n\nScope:\n\n\nTrosifol PVB, manufactured in Troisdorf based on Mowital\u24c7 resin from\nFrankfurt.\n\n\nPlate glass for construction and interlayers, 01/09/2022\n(PCR checked and approved by the SVR)\n\n\nThe owner of the declaration shall be liable for the underlying information\nand evidence; the IBU shall not be liable with respect to manufacturer\ninformation, life cycle assessment data and evidences.\n\n\nIssue date\n\n\n12/05/2023\n\n\nThe EPD was created according to the specifications of EN 15804+A2. In\nthe following, the standard will be simplified as EN 15804.\n\n\nValid to\n\n\nVerification\n\n\n11/05/2028\n\n\nThe standard EN 15804 serves as the core PCR\nIndependent verification of the declaration and data according to ISO\n14025:2011\n\n\ninternally \u2611\n\n\nexternally\n\n\nNam Paten\n\n\nDipl.-Ing. Hans Peters\n\n\n(Chairman of Institut Bauen und Umwelt e.V.)\n\n\nHam Peter\nDipl.-Ing. Hans Peters\n\n\n\u041c\u0438\u043b\u0438\n\n\nDr. Matthew Fishwick,\n(Independent verifier)\n\n\n(Managing Director Institut Bauen und Umwelt e.V.)\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\nkuraray\n\n\nProduct\n\n\nschueren/\n\n\nProduct description/Product definition\n\n\nConstructional data\n\n\nKuraray's polyvinyl butyral (PVB) thermoplastic films are tough,\nresilient safety interlayers used in laminated architectural safety\nglass. These Trosifol\u24c7 PVB interlayers offer safety advantages\nby retaining dangerous shards in case of glass breakage. They\nare commonly used as safety glass interlayers available\nworldwide.\n\n\nUnit\n\n\nValue\n\n\nName\n\n\nRefractive index acc. to DIN EN ISO 489\n\n\n1.48\n\n\n0.21 W/mK\n\n\nThermal conductivity acc. to DIN EN 993-15\nThermal expansion coefficient acc. to ISO\n11359-2\n\n\n0.00017 K-1\n\n\nSpecific heat capacity\n\n\n1.9 kJ/kgK\n1E+13 \u03a9\n\n\nThis EPD covers all Trosifol\u24c7 PVB products produced in\nTroisdorf based on Mowital\u24c7 resin produced in Frankfurt.\nProduct codes Trosifol\u24c7 B2XX and Trosifol B8XX.\n\n\nSurface resistivity acc. to DIN 53482\nTensile strength acc. to ISO 527-3\nElongation at break acc. to ISO 527-3\n\n\nN/mm\u00b2\n\n\n20\n\n\n250\n\n\n%\n\n\nTg acc. to DMA, 3K/min, 1Hz\n\n\nFor the use and application of the product, the respective\nnational provisions at the place of use apply, in Germany, for\nexample, the building codes of the federal states and the\ncorresponding national specifications.\n\n\n\u00b0C\n\n\n32\n\n\nPerformance data of the product with respect to its\ncharacteristics in accordance with the relevant technical\nprovision (no CE-marking).\n\n\nApplication\n\n\nBase materials/Ancillary materials\n\n\nTrosifol\u24c7 PVB film needs to be laminated between two pieces\nof glass. This sandwich arrangement is called laminated safety\nglass according to EN ISO 14449. Special Trosifol\u24c7 PVB\ngrades offer additional decorative, acoustic, UV managing and\nstructural properties.\n\n\nThe main constituents of Trosifol\u24c7 PVB film are (in mass\npercentages):\n\n\n-PVB resin ~72 %\n\n\n-Plasticizer ~27-28 %\n\n\n-Additives and water <1%\n\n\n1) This product contains substances listed in the candidate list\n(date: 17.01.2023) exceeding 0.1 percentage by mass: no\n\n\nTechnical Data\n\n\nFor calculating the light, solar and heat parameters of glazing\nspecifically containing films from the Trosifol\u24c7 & SentryGlas\u24c7\nproduct range, please go to:\n\n\n2) This product contains other Carcinogenic, Mutagenic,\nReprotoxic (CMR) substances in categories 1A or 1B which are\nnot on the candidate list, exceeding 0.1 percentage by mass:\n\n\nhttps://www.trosifol.com/winslt-tool/\n\n\nno\n\n\nSound Control data can be found here:\n\n\n3) Biocide products were added to this construction product or it\nhas been treated with biocide products (this then concerns a\ntreated product as defined by the (EU) Ordinance on Biocide\nProducts No. 528/2012): no\n\n\nhttps://www.trosifol.com/soundlab-ai/\n\n\nThe following data are valid for Trosifol\u24c7 Clear / Trosifol\u24c7\nUltraClear. Other product's data can be found in our laminator\nbrochure:\n\n\nReference service life\n\n\nThe reference service life is typically determined by the glass\nand not by the interlayer.\n\n\nhttps://www.trosifol.com/de/salessupport/downloads/produktbro\n\n\nLCA: Calculation rules\n\n\nproducts and energy, as well as waste processing up to the\nend-of-waste state or disposal of final residues during the\nproduct stage.\n\n\nDeclared Unit\n\n\nThis declaration refers to the declared unit of 1 m\u00b2 of PVB film\n(Trosifol\u24c7). The grammage of the PVB film is 0.775 kg/m\u00b2.\n\n\nThese modules consider the manufacturing of system\ncomponents/raw materials, the transport to the production site\nand the production processes of the products under study. The\nimpact of packaging materials is included.\n\n\nDeclared unit - PVB film (Trosifol\u24c7)\n\n\nName\n\n\nValue\n\n\nUnit\nm\u00b2\nkg/m^2\n\n\nDeclared unit\n\n\n1\n0.775\n0.00076\n1.07\n\n\nGrammage\n\n\nModule A5:\n\n\nLayer thickness\n\n\nm\n\n\nTreatment and disposal of packaging material. Credits for\npotential avoided burdens due to energy substitution of\nelectricity and thermal energy generation are declared in\nmodule D and affect only the rate of primary material (no\nsecondary materials).\n\n\nDensity\n\n\ng/cm^3\n\n\nSystem boundary\n\n\nThe type of EPD is cradle-to-gate with options, modules C1-\nC4, and module D (A1-A3, C, D and additional module A5). In\nthe following section, a detailed description of the specific\nsystem boundaries is given:\n\n\nModule C1 to C4:\n\n\nThe end-of-life scenarios are as follows:\n\n\nC1 Deconstruction/demolition: Dismantling is manual\n(no environmental burden).\n\n\n\u2022\n\n\nModule A1 to A3:\n\n\nThe product stage includes the provision of all materials,\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n2\n\n\nkuraray\n\n\n\u2022 C2 - Transport to treatment/disposal site: Average\ntransport distance from the demolition site to waste\ntreatment is assumed as 50 km to the landfill.\n\n\nLand or region, in which the declared product system is\nmanufactured, used or handled at the end of the product's\nlifespan: Europe\n\n\n\u2022 C3 - Waste processing for reuse, recovery or recycling:\nNo waste processing (no environmental burden).\n\u2022 C4 - Disposal: PVB film is 100% landfilled.\nModule D:\n\n\nComparability\n\n\nBasically, a comparison or an evaluation of EPD data is only\npossible if all the data sets to be compared were created\naccording to EN 15804 and the building context, respectively\nthe product-specific characteristics of performance, are taken\ninto account. Background datasets: GaBi ts 10.6 software\nsystem and GaBi Professional 2022.1 LCI database.\n\n\nFor the thermal and electrical energy generated in Module A5\ndue to the thermal treatment of packaging and product waste,\navoided burdens have been calculated by the inversion of the\nelectricity grid mix and thermal energy from natural gas, using\nEuropean datasets.\n\n\nGeographic Representativeness\n\n\nLCA: Scenarios and additional technical information\n\n\nEnd of life (C1-C4)\n\n\nCharacteristic product properties biogenic carbon\nBiogenic carbon is only present in the packaging (wooden\npallets and cartons).\n\n\nThe end-of-life scenarios are as follows:\n\n\nC1 The deconstruction of the PVB film is assumed to be done\nmanually. Therefore, no environmental loads for the dismantling\nof this product are considered.\nC2 - Transport to treatment/disposal site: Average transport\ndistance from the demolition site to waste treatment is assumed\nas 50 km to landfill.\n\n\nAssumed water content in wooden pallets (packaging): 18 %.\nAssumed carbon content: dry wood mass consists of 50 %\nbiogenic carbon and paper/cardboard 43 %.\n\n\nC4 - Disposal: The PVB film is 100% landfilled.\nName\n\n\nThe biogenic carbon content of the packaging is thus: 0.129 kg\npallet/declared unit * 0.82 * 0.5 kg C / kg pallet (abs. dry) +\n0.082 kg *0.43 kg C/kg cardboard= 0.08815 kg C/declared unit.\nInformation on describing the biogenic Carbon Content at\nfactory gate\n\n\nValue Unit\n\n\nCollected as mixed construction waste [PVB film\nper FU]\n\n\n0.775 kg\n\n\n0.775 kg\n\n\nLandfilling [PVB film per FU]\n\n\nName\n\n\nValue Unit\n0.08815\n\n\nBiogenic carbon content in accompanying\npackaging\n\n\nkg\nC\n\n\nReuse, recovery and/or recycling potentials (D), relevant\nscenario information\n\n\nFor the thermal and electrical energy generated in Module A5\ndue to the thermal treatment of packaging, avoided burdens\nhave been calculated by the inversion of the electricity grid mix\nand thermal energy from natural gas, using European datasets.\n\n\nThe following technical scenario information is required for the\ndeclared modules.\n\n\nInstallation into the building (A5)\n\n\nThe packaging material treatment and disposal are also\nconsidered in module A5.\n\n\nValue Unit\n\n\nName\n\n\nOutput substances following waste treatment on\nsite [packaging materials per FU]\n\n\n0.2301 kg\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n3\n\n\nkuraray\n\n\nLCA: Results\n\n\nDESCRIPTION OF THE SYSTEM BOUNDARY (X = INCLUDED IN LCA; ND = MODULE OR INDICATOR NOT DECLARED; MNR =\nMODULE NOT RELEVANT)\n\n\nBenefits and\nloads beyond\nthe system\nboundaries\n\n\nConstruction\nprocess stage\n\n\nProduct stage\n\n\nUse stage\n\n\nEnd of life stage\n\n\nOperational energy\n\n\nTransport from the\n\n\nOperational water\n\n\n2 Waste processing\nC3\n\n\nC1\n2 De-construction\n\n\nRefurbishment\n\n\n\u2611 gate to the site\nA4\n\n\nManufacturing\n\n\nRaw material\n\n\nReplacement\n\n\nMaintenance\n\n\nRecovery-\n\n\nRecycling-\n\n\ndemolition\n\n\nAssembly\n\n\nTransport\n\n\nTransport\n\n\nDisposal\n\n\npotential\n\n\nReuse-\n\n\nsupply\n\n\nRepair\n\n\nUse\n\n\nuse\n\n\nuse\n\n\nDX\n\n\nA2\n\n\nA3\n\n\nA5\n\n\nC2\n\n\nC4\n\n\nA1\n\n\nB1\n\n\nB2\n\n\nB3\n\n\nB4\n\n\nB5\n\n\nB6\n\n\nB7\n\n\nMND X MND MND MNR MNR MNR MND MND\n\n\n\u2717\n\n\n\u2717\n\n\nX\n\n\n\u2717\n\n\nRESULTS OF THE LCA - ENVIRONMENTAL IMPACT according to EN 15804+A2: 1 m\u00b2 PVB film (Trosifol\u24c7)\nParameter\n\n\nUnit\n\n\nA1-A3\n\n\nA5\n\n\nC2\n\n\nC3\n\n\nC1\n\n\nC4\n\n\nD\n\n\nkg CO2 eq\n\n\n3.26E+00\n\n\n2.37E-03\n\n\n-1.19E-01\n\n\nGWP-total\n\n\n3.77E-01\n\n\n0\n\n\n0\n\n\n5.49E-02\n\n\nkg CO2 eq\n\n\nGWP-fossil\n\n\n2.37E-03\n\n\n3.55E+00\n\n\n5.77E-02\n\n\n0\n\n\n0\n\n\n5.49E-02\n\n\n-1.18E-01\n\n\nGWP-biogenic\n\n\nkg CO2 eq\n\n\n3.2E-01\n\n\n-2.95E-01\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n-6.05E-04\n\n\nkg CO2 eq\n\n\n9.09E-06\n\n\nGWP-luluc\n\n\n5.4E-04\n\n\n5.85E-06\n\n\n0\n\n\n2.67E-05\n\n\n-1.3E-05\n\n\n0\n\n\nkg CFC11 eq\n\n\n1.91E-09\n\n\nODP\n\n\n5.63E-14\n\n\n0\n\n\n4.84E-16\n\n\n7.38E-14\n\n\n-7.99E-13\n\n\n0\n\n\n4.67E-03\n\n\nmol H+ eq\n\n\n1.63E-04\n\n\nAP\n\n\n7.13E-05\n\n\n0\n\n\n7.48E-06\n\n\n0\n\n\n-1.56E-04\n\n\nEP-freshwater\n\n\nkg P eq\nkg N eq\nmol N eq\nkg NMVOC\neq\nkg Sb eq\nMJ\nm\u00b3 world eq\ndeprived\n\n\n-1.63E-07\n\n\n2.99E-05\n\n\n1.62E-08\n\n\n0\n\n\n4.71E-09\n\n\n0\n\n\n1.03E-05\n\n\nEP-marine\nEP-terrestrial\n\n\n3.5E-06\n\n\n-4.22E-05\n\n\n1.47E-03\n\n\n2.45E-05\n\n\n0\n\n\n3.6E-05\n\n\n0\n\n\n3.28E-04\n\n\n1.64E-02\n\n\n3.95E-04\n\n\n0\n\n\n3.9E-05\n\n\n0\n\n\n-4.52E-04\n\n\n6.56E-05\n\n\nPOCP\n\n\n5.78E-03\n\n\n0\n\n\n6.82E-06\n\n\n0\n\n\n1.16E-04\n\n\n-1.18E-04\n\n\n-1.78E-08\n\n\nADPE\n\n\n1.61E-06\n\n\n1.45E-09\n\n\n0\n\n\n2.36E-10\n\n\n0\n\n\n3.81E-09\n\n\nADPF\n\n\n7.95E+01\n\n\n1.39E-01\n\n\n0\n\n\n3.13E-02\n\n\n7.79E-01\n\n\n-2.01E+00\n\n\n0\n\n\nWDP\n\n\n3.87E-02\n\n\n-7.12E-01\n\n\n0\n\n\n1.01E-05\n\n\n0\n\n\n-5.42E-04\n\n\n-1.26E-02\n\n\nGWP = Global warming potential; ODP = Depletion potential of the stratospheric ozone layer; AP = Acidification potential of land and water; EP =\nEutrophication potential; POCP = Formation potential of tropospheric ozone photochemical oxidants; ADPE = Abiotic depletion potential for non-fossil\nresources; ADPF = Abiotic depletion potential for fossil resources; WDP = Water (user) deprivation potential)\nRESULTS OF THE LCA - INDICATORS TO DESCRIBE RESOURCE USE according to EN 15804+A2: 1 m\u00b2 PVB film (Trosifol\u24c7)\nParameter\n\n\nA5\n45\n3.42E+00\n\n\nA1-A3\n1.03E+01\n\n\nUnit\nMJ\n\n\nC1\n\n\nC2\n\n\nC4\n\n\nD\n\n\nC3\n\n\n2.06E-03\n\n\nPERE\n\n\n0\n\n\n6.41E-02\n0\n\n\n-5.52E-01\n\n\n0\n\n\n3.39E+00\n\n\nPERM\n\n\n-3.39E+00\n\n\n0\n\n\nMJ\n\n\n0\n\n\n0\n\n\n0\n-5.52E-01\n-2.01E+00\n0\n-2.01E+00\n\n\n|PERT\n\n\n1.37E+01\n\n\nMJ\n\n\n3.22E-02\n\n\n0\n\n\n2.06E-03\n\n\n0\n\n\n6.41E-02\n\n\n2.5E+01\n\n\nPENRE\n\n\nMJ\n\n\n5.47E+01\n\n\n6.94E-01\n\n\n0\n\n\n3.13E-02\n\n\n0\n\n\n-2.43E+01\n\n\nPENRM\n\n\n2.48E+01\n\n\n-5.55E-01\n\n\n0\n\n\n0\n\n\n0\n\n\nMJ\n\n\nPENRT\n\n\n7.96E+01\n\n\n1.39E-01\n\n\n0\n\n\n3.13E-02\n\n\n0\n\n\n7.8E-01\n\n\nMJ\n\n\nSM\n\n\nkg\nMJ\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nRSF\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nNRSF\n\n\nMJ\nm\u00b3\n\n\n0\n1.64E-02\n\n\n0\n\n\n0\n\n\n0\n0\n\n\n0\n\n\n0\n\n\n0\n9.16E-04\n\n\n1.6E-06\n\n\n1.01E-05\n\n\n-5.31E-04\n\n\nFW\n\n\n0\n\n\nPERE = Use of renewable primary energy excluding renewable primary energy resources used as raw materials; PERM = Use of renewable primary\nenergy resources used as raw materials; PERT = Total use of renewable primary energy resources; PENRE = Use of non-renewable primary energy\nexcluding non-renewable primary energy resources used as raw materials; PENRM = Use of non-renewable primary energy resources used as raw\nmaterials; PENRT = Total use of non-renewable primary energy resources; SM = Use of secondary material; RSF = Use of renewable secondary fuels;\nNRSF = Use of non-renewable secondary fuels; FW = Use of net fresh water\n\n\nRESULTS OF THE LCA \u2013 WASTE CATEGORIES AND OUTPUT FLOWS according to EN 15804+A2:\n\n\n1 m\u00b2 PVB film (Trosifol\u00ae)\n\n\nParameter\n\n\nUnit\n\n\nA1-A3\n8.63E-08\n\n\nA5\n1.17E-11\n\n\nC3\n\n\nC1\n\n\nC2\n\n\nC4\n\n\nD\n\n\n1.2E-10\n7.72E-01\n9.58E-06\n\n\nkg\n\n\n0\n\n\n1.37E-13\n\n\n-2.73E-10\n\n\nHWD\n\n\n0\n\n\n-1.02E-03\n\n\nkg\n\n\n3.89E-02\n\n\n2.29E-02\n\n\n0\n\n\n5.1E-06\n\n\nNHWD\n\n\n0\n\n\n3.94E-08\n\n\nkg\n\n\n7.58E-04\n\n\n6.07E-06\n\n\n0\n\n\n0\n\n\n-1.58E-04\n\n\nRWD\n\n\nCRU\n\n\nkg\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nMFR\n\n\nkg\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nMER\n\n\nkg\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n5.32E-01\n\n\nEEE\n\n\nMJ\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nEET\n\n\nMJ\n\n\n0\n\n\n9.58E-01\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n4\n\n\nkuraray\n\n\nHWD = Hazardous waste disposed; NHWD = Non-hazardous waste disposed; RWD = Radioactive waste disposed; CRU = Components for re-use;\nMFR = Materials for recycling; MER = Materials for energy recovery; EEE = Exported electrical energy; EET = Exported thermal energy\n\n\nRESULTS OF THE LCA - additional impact categories according to EN 15804+A2-optional:\n\n\n1 m\u00b2 PVB film (Trosifol\u00ae)\n\n\nParameter\n\n\nUnit\nDisease\nincidence\n\n\nA1-A3\n\n\nA5\n\n\nC1\n\n\nC2\n\n\nC3\n\n\nC4\n\n\nD\n\n\nPM\n\n\n6.26E-08\n\n\n5.26E-10\n\n\n0\n\n\n4.26E-11\n\n\n0\n\n\n1.56E-09\n\n\n-1.29E-09\n\n\nIR\n\n\nkBq U235 eq\nCTUe\n\n\n7.89E-02\n\n\n9.02E-04\n\n\n0\n\n\n0\n\n\n1.41E-03\n\n\n-2.68E-02\n\n\n4E-06\n\n\n8.63E-02\n\n\n7.63E-01\n\n\nETP-fw\n\n\n6.53E+01\n\n\n0\n\n\n2.4E-02\n\n\n0\n\n\n-4.41E-01\n\n\n-2.03E-11\n\n\nHTP-C\nHTP-nc\n\n\nCTUh\n\n\n9.32E-09\n\n\n4.13E-12\n\n\n0\n\n\n4.81E-13\n\n\n0\n\n\n3.42E-11\n\n\nCTUh\nSQP\n\n\n4.94E-08\n4.91E+01\n\n\n2.77E-11\n\n\n2.79E-10\n3.71E-02\n\n\n0\n\n\n0\n\n\n2.87E-09\n\n\n-7.8E-10\n\n\nSQP\n\n\n0\n\n\n9.38E-03\n\n\n0\n\n\n5.61E-02\n\n\n-3.59E-01\n\n\nPM = Potential incidence of disease due to PM emissions; IR = Potential Human exposure efficiency relative to U235; ETP-fw = Potential comparative\nToxic Unit for ecosystems; HTP-c = Potential comparative Toxic Unit for humans (cancerogenic); HTP-nc = Potential comparative Toxic Unit for humans\n(not cancerogenic); SQP = Potential soil quality index\n\n\nDisclaimer 1 - for the indicator \"Potential Human exposure efficiency relative to U235\". This impact category deals mainly with the\neventual impact of low-dose ionizing radiation on human health of the nuclear fuel cycle. It does not consider effects due to possible\nnuclear accidents, occupational exposure or radioactive waste disposal in underground facilities. Potential ionizing radiation from the\nsoil, radon and from some construction materials is also not measured by this indicator.\n\n\nDisclaimer 2 - for the indicators 'abiotic depletion potential for non-fossil resources', 'abiotic depletion potential for fossil resources',\n'water (user) deprivation potential, deprivation-weighted water consumption', 'potential comparative toxic unit for ecosystems', 'potential\ncomparative toxic unit for humans - cancerogenic', 'Potential comparative toxic unit for humans - not cancerogenic', 'potential soil\nquality index'. The results of this environmental impact indicator shall be used with care as the uncertainties on these results are high\nas there is limited experience with the indicator.\n\n\nReferences\n\n\nEN ISO 14025:2011, Environmental labels and declarations \u2014\nType III environmental declarations - Principles and\nprocedures.\n\n\nStandards\n\n\nDIN 53482\n\n\nDIN 53482:1967-01, Testing of Insulating Materials;\nDetermination of Electrical Resistances Values.\n\n\nFurther References\n\n\nDIN EN ISO 489\n\n\nDIN EN ISO 489:1999-08, Plastics - Determination of the\nrefractive index (ISO 489:1999).\n\n\nCandidate list\n\n\nCandidate List of substances of very high concern for\nAuthorisation, published on ECHA website, latest version\n17.01.2023 (https://echa.europa.eu/candidatelist-table)\n\n\nDIN EN 993-15\n\n\nDIN EN 993-15:2005-07, Methods of test for dense shaped\nrefractory products - Part 15: Determination of thermal\nconductivity by the hot-wire (parallel) method.\n\n\nGaBi\n\n\nGaBi Software System and Database for Life Cycle\nEngineering, 1992-2021, Sphera Solutions GmbH, Leinfelden-\nEchterdingen, with acknowledgement of\nLBP University of Stuttgart, program version GaBi 10; database\nversion 2022.1.\n\n\nDIN EN ISO 527-3\n\n\nDIN EN ISO 527-3:2019-02, Plastics - Determination of tensile\nproperties - Part 3: Test conditions for films and sheets (ISO\n527-3:2018).\n\n\nGaBi documentation\n\n\nGaBi dataset documentation for the software system and\ndatabases, LBP, University of Stuttgart and Sphera Solutions\nGmbH, Leinfelden-Echterdingen,\n\n\nEN 15804\n\n\nEN15804:2012+A1:2013, Sustainability of construction works\n- Environmental Product Declarations Core rules for the\nproduct category of construction products.\n\n\n-\n\n\n2021. (http://www.gabi-\n\n\nsoftware.com/support/gabi/gabi[1] database-2021-lci-\n\n\ndocumentation/)\n\n\nEN 15804\n\n\nEN 15804:2012+A2:2019+AC:2021, Sustainability of\nconstruction works - Environmental Product Declarations -\nCore rules for the product category of construction products.\n\n\nIBU 2021\n\n\nInstitut Bauen und Umwelt e.V.: General Instructions for the\nEPD programme of Institut Bauen und Umwelt e. V., Version\n2.0, Berlin: Institut Bauen und Umwelt e. V., 2021 HYPERLINK\n\"http://www.ibu-epd.\" www.ibu-epd.com\n\n\nEN ISO 14449\n\n\nEN 14449:2005/AC:2005, Glass in building - Laminated glass\nand laminated safety glass - Evaluation of conformity/Product\nstandard.\n\n\nOrdinance on Biocide Products No. 528/2012\nRegulation (EU) No 528/2012 of the European Parliament and\nof the Council of 22 May 2012 concerning the making available\non the market and use of biocidal products\n\n\nISO 11359-2\n\n\nISO 11359-2:2021-11, Plastics - Thermomechanical analysis\n(TMA) - Part 2: Determination of coefficient of linear thermal\nexpansion and glass transition temperature.\n\n\nPCR Part A\n\n\nPCR Part A: Calculation rules for the Life Cycle Assessment\nand Requirements on the Background Report according to EN\n\n\nISO 14025\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n5\n\n\nkuraray\n\n\n15 804+A2:2019, Version 1.3, Institut Bauen und Umwelt e.V.,\n2020.\n\n\nProduct Category Rules for Building Products, Part B:\n\n\nRequirements on the EPD for plate glass for construction and\ninterlayers, version 1.6, 2022 www.bau-umwelt.de\n\n\nPCR Part B\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n6\n\n\nkuraray\n\n\nPublisher\n\n\nInstitut Bauen und Umwelt e.V.\n\n\n+49 (0)30 3087748-0\ninfo@ibu-epd.com\nwww.ibu-epd.com\n\n\nHegelplatz 1\n10117 Berlin\nGermany\n\n\nInstitut Bauen\nund Umwelt e.V.\n\n\nProgramme holder\n\n\n+49 (0)30 3087748-0\ninfo@ibu-epd.com\nwww.ibu-epd.com\n\n\nInstitut Bauen und Umwelt e.V.\nHegelplatz 1\n\n\n10117 Berlin\nGermany\n\n\nInstitut Bauen\nund Umwelt e.V.\n\n\nAuthor of the Life Cycle Assessment\n\n\nSphera Solutions GmbH\n\n\n+49 711 341817-0\ninfo@sphera.com\nwww.sphera.com\n\n\nsphera\u24c7\n\n\nHauptstra\u00dfe 111-113\n\n\n70771 Leinfelden-Echterdingen\nGermany\n\n\nOwner of the Declaration\n\n\nKuraray Europe GmbH\nPhilipp-Reis-Str. 4\n\n\n+49 69 305 85 300\ntrosifol@kuraray.com\nhttps://www.kuraray.eu/\n\n\nkuraray\n\n\n65795 Hattersheim\nGermany\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n7\n"}, "expected_output": {"claims": [{"unit": "%", "value": 72, "evidence": ["-PVB resin ~72 %", "The main constituents of Trosifol\u24c7 PVB film are (in mass percentages):"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_4a571861e670e7da"}}
{"id": "d210b616-51c0-4aad-8439-c71f6e23112f", "input": {"query": "What is the keyboard unit weight (product weight excluding packaging). I need specific numeric values in pounds or grams/kg.", "source_url": "https://www.ergodirect.com/attachments/26497/Product-Sheet-Kensington-KB675-Pro-Fit-Ergo-EQ-TKL-Rechargeable-Keyboard.pdf", "document_text": "133\n\n\nPay\n\n\nKensington\n\n\nProduct\nSheet\n\n\nF10\n\n\nP\n\n\n0\n\n\n9\n\n\nA\n\n\n^\n\n\n9\n\n\nH\n\n\nF6\n\n\nn\n\n\nKensington\n\n\nF2\n\n\nR\n\n\nE\n\n\n2\n\n\nEsc\n\n\nW\n\n\nD\n\n\nC\n\n\nS\n\n\nA\n\n\nTab\n\n\nX\n\n\nCaps Lock\n\n\nZ\n\n\nea\n\n\nFn\n\n\nCtri\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nSculpted for\ncomfort and\nperformance.\n\n\nFor professionals wanting to reduce the fatigue associated with spending long\nhours typing each day, the Pro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\nsupports all-day comfort and productivity with style.\n\n\nThe sculpted ergonomic design and curved key layout promotes proper hand\nand wrist posture, accommodating the natural contour of your hands to provide\na more comfortable and productive typing experience. Performance-enhancing\nfeatures include multiple connectivity options, government-grade encryption\nsecurity, universal compatibility, and a long-lasting rechargeable battery.\n\n\nComplementing a modern workspace, the Pro Fit\u2122 Ergo KB675 EQ TKL\nRechargeable Keyboard is an ideal choice for professionals seeking to elevate\ntheir typing experience with style, comfort, and performance.\n\n\nErgonomic design\n\n\n\u2022\n\n\nSupports up to three devices\n\n\n\u2022\n\n\nreddot winner 2024\n\n\nWindows and macOS compatible\n\n\n\u2022\n\n\nSwift Pair Technology reduces pairing hassles for Windows devices\n\n\n\u2022\n\n\nPLASTIC\n\n\nCONTAINS\n\n\nVideo conferencing keys\n\n\n\u2022\n\n\n\u2022 RGB indicators for connection status, caps lock, and battery life\n\n\nRECYCLED\n\n\nCustomizable keys\n\n\n43%\n\n\n\u2022\n\n\nPOST\n\n\nIncludes 43% post-consumer recycled (PCR) content of total plastic components\n(Excludes PCB assembly, rechargeable battery, 2.4GHz dongle and charging cable)\n\n\n\u2022\n\n\n\u2022 3-year limited warranty\n\n\n\u2022\n\n\nKensington\n\n\nProduct\n\n\nSheet\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nCONTAINS\n\n\nPLASTIC\n\n\nRECYCLED\n\n\n43%\n\n\nPOST-\n\n\n2.4GHz\n\n\nSwift Pair Technology\n\n\nSustainability Attributes\nIncludes 43% post-\nconsumer recycled (PCR)\ncontent of total plastic\ncomponents (Excludes PCB\nassembly, rechargeable\nbattery, 2.4GHz dongle and\ncharging cable).\n\n\nErgonomic Design\n\n\nMultiple Connection\nOptions\n\n\nThe thoughtfully designed\ncompact curved key\nlayout, built-in wrist pad,\nand front keyboard legs\nsupport optimal hand\n\n\nEnhanced Bluetooth\u00ae\nconnection technology\ndesigned to reduce pairing\nhassles and simplify\nconnections for Windows\ndevices.\n\n\nSupports up to three\ndevices two via\n\n\nBluetooth connections,\none via the 2.4GHz USB\nreceiver with 128-bit\nAES government-grade\nencryption security.\nWindows and macOS\ncompatible.\n\n\nand wrist positioning for\nall-day ergonomic comfort\nand productivity.\n\n\n(4)\n500\nmAh\n\n\nF10 F11F12D\n\n\nE\n\n\nBackspace\n\n\nHome\n\n\nP\n\n\n]\n\n\n[\n\n\nO\n\n\nCustomizable Keys\n\n\nProductivity Enhancing\nKeys and Lights\n\n\nLong-Lasting Power\n\n\nThe built-in 500mAh\nrechargeable battery\nincludes a convenient\nUSB-C charging port and\ncable for uninterrupted\nproductivity.\n\n\nCustomize your experience\nby programming keys,\ncreating macros, adjusting\nkey mappings, managing\nprofiles, and more with\nthe optional Kensington\nKonnect\u2122 software.\n\n\nVideo conferencing keys\nsupport efficient meeting\ntransitions. Vibrant RGB\nindicators allow you\nto easily monitor your\nconnection status, caps\nlock, and battery life.\n\n\nKensington\n\n\nTechnical\nSpecifications\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nSystem Requirements\n\n\n1 x USB-A\n\n\nSystem Resource Requirement\n\n\nfor charging and wired connection.\nWindows, macOS\n\n\nOperating Systems Supported\n\n\nProduct Features\n\n\n2.4GHz, Bluetooth\n\n\nConnection\nMulti-device\n\n\nUp to 3 (2 x BT 5.2, 1 x 2.4GHz)\n\n\nTKL, 93 keys (89 + 4 meeting control buttons)\n\n\nKey Layout\n\n\nMeeting control button\nSwitches\n\n\n4 (Mute or Unmute/Video off or on/Share Screen/End call)\nMembrane\n\n\nOuter case and keycaps\n\n\n43% PCR ABS\n\n\nWrist pad\n\n\nYes (Non-Detachable)\n\n\nBacklight\n\n\n\u039d\u0391\n\n\nBattery\n\n\nRechargeable Battery\n\n\nKeyboard: USB-C / Host: USB-A\n\n\nCharging port\n\n\nCharging time\n\n\n3 hours to fully charged from low battery\nAt least 2.5 months after fully charged\nWindows and macOS\n\n\nBattery life\n\n\nKensington Konnect\u2122 Support\n\n\nProgrammable keys\nKeyboard Color\n\n\n20\n\n\nBlack\n\n\nProduct Dimensions\n\n\nKeyboard Unit Dimensions\n(Width x Length x Height)\nKeyboard Unit Weight\nKeyboard Color\n\n\n8.7 x 14.8 x 1.37\"\n\n\n(220.9 x 375.9 x 34.8mm)\n\n\n1.32lbs (0.6kg)\n\n\nBlack\n\n\nPackaging Dimensions\n\n\nPackaging Style\nPackaging Dimensions\n(Width x Length x Height)\nPackaging Weight\n\n\nRetail box (FSC) with inner brown box (FSC)\nUnit: 10.15 x 15.62 x 1.73\" (258 x 397 x 44mm)\nCarton: 11.5 x 16.3 x 8.07\" (292 x 414 x 205mm)\n\n\nUnit: 2.16lbs (0.98kg)\n\n\nMaster Carton: 9.92lbs (4.5kg)\n\n\nMaster Carton\nPackaging Content\n\n\n4pcs\n\n\n1 x Keyboard\n\n\n1 x 6' (1.8m) USB-C to USB-A cable\n\n\n1 x 2.4GHz Wireless receiver\n\n\n1 x Welcome card\n\n\n1 x Instruction guide\n\n\n1 x Battery warning insert\n\n\n1 x Compliance sheet\n\n\n1 x Warranty card\n\n\nKensington\n\n\nTechnical\nSpecifications\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nPower & Environmental\n\n\n500mAh rechargeable battery\n\n\nBattery\n\n\n32 to 104\u00b0F (0 to 40\u00b0C)\n\n\nOperating Temperature\n\n\n5 to 140\u00b0F (-15 to 60\u00b0C)\n\n\nStorage Temperature\n\n\nRelative Humidity\n\n\n0% to 85% non-condensing\n\n\nCountry of Origin, Regulatory, And Warranty\n\n\nCountry of Origin\n\n\nChina\n\n\n3-year Limited\n\n\nWarranty\n\n\nCertifications & Compliance\n\n\nFCC, IC, RCM, CB, China SRRC, RoHS, ACCO RSL, BC,\nNrCAN, CE, UKCA, WEEE, NOM, PSTI\n\n\nSustainability\n\n\nOuter Case and Keycaps\nPaper Packaging\n\n\nMade with 43% Post Consumer Recycled ABS\nFSC\u2122-certified\n\n\nKensington\n\n\nAll specifications are subject to change without notice. Products may not be available in all markets. Kensington and the ACCO\nname and design are registered trademarks of ACCO Brands. Kensington The Professionals' Choice is a trademark of ACCO Brands.\nAll other registered and unregistered trademarks are the property of their respective owners. 2024 Kensington Computer\nProducts Group, a division of ACCO Brands. All rights reserved.\n\n\nACCO\nBRANDS\n\n\nThe Professionals' Choice\u2122\n\n\nFOR MORE INFORMATION CONTACT: 1-855-692-0054 | sales@kensington.com\n"}, "expected_output": {"claims": [{"unit": "lbs", "value": 1.32, "evidence": ["1.32lbs (0.6kg)"]}]}, "metadata": {"product_category": "Furniture & other goods", "request_id": "req_9a170a87c1e30ebc"}}
{"id": "768e49d6c5ef6ac825725d11", "input": {"query": "What is the packaging weight for a keyboard unit? I need specific numeric values in pounds or grams/kg.", "source_url": "https://www.ergodirect.com/attachments/26497/Product-Sheet-Kensington-KB675-Pro-Fit-Ergo-EQ-TKL-Rechargeable-Keyboard.pdf", "document_text": "133\n\n\nPay\n\n\nKensington\n\n\nProduct\nSheet\n\n\nF10\n\n\nP\n\n\n0\n\n\n9\n\n\nA\n\n\n^\n\n\n9\n\n\nH\n\n\nF6\n\n\nn\n\n\nKensington\n\n\nF2\n\n\nR\n\n\nE\n\n\n2\n\n\nEsc\n\n\nW\n\n\nD\n\n\nC\n\n\nS\n\n\nA\n\n\nTab\n\n\nX\n\n\nCaps Lock\n\n\nZ\n\n\nea\n\n\nFn\n\n\nCtri\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nSculpted for\ncomfort and\nperformance.\n\n\nFor professionals wanting to reduce the fatigue associated with spending long\nhours typing each day, the Pro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\nsupports all-day comfort and productivity with style.\n\n\nThe sculpted ergonomic design and curved key layout promotes proper hand\nand wrist posture, accommodating the natural contour of your hands to provide\na more comfortable and productive typing experience. Performance-enhancing\nfeatures include multiple connectivity options, government-grade encryption\nsecurity, universal compatibility, and a long-lasting rechargeable battery.\n\n\nComplementing a modern workspace, the Pro Fit\u2122 Ergo KB675 EQ TKL\nRechargeable Keyboard is an ideal choice for professionals seeking to elevate\ntheir typing experience with style, comfort, and performance.\n\n\nErgonomic design\n\n\n\u2022\n\n\nSupports up to three devices\n\n\n\u2022\n\n\nreddot winner 2024\n\n\nWindows and macOS compatible\n\n\n\u2022\n\n\nSwift Pair Technology reduces pairing hassles for Windows devices\n\n\n\u2022\n\n\nPLASTIC\n\n\nCONTAINS\n\n\nVideo conferencing keys\n\n\n\u2022\n\n\n\u2022 RGB indicators for connection status, caps lock, and battery life\n\n\nRECYCLED\n\n\nCustomizable keys\n\n\n43%\n\n\n\u2022\n\n\nPOST\n\n\nIncludes 43% post-consumer recycled (PCR) content of total plastic components\n(Excludes PCB assembly, rechargeable battery, 2.4GHz dongle and charging cable)\n\n\n\u2022\n\n\n\u2022 3-year limited warranty\n\n\n\u2022\n\n\nKensington\n\n\nProduct\n\n\nSheet\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nCONTAINS\n\n\nPLASTIC\n\n\nRECYCLED\n\n\n43%\n\n\nPOST-\n\n\n2.4GHz\n\n\nSwift Pair Technology\n\n\nSustainability Attributes\nIncludes 43% post-\nconsumer recycled (PCR)\ncontent of total plastic\ncomponents (Excludes PCB\nassembly, rechargeable\nbattery, 2.4GHz dongle and\ncharging cable).\n\n\nErgonomic Design\n\n\nMultiple Connection\nOptions\n\n\nThe thoughtfully designed\ncompact curved key\nlayout, built-in wrist pad,\nand front keyboard legs\nsupport optimal hand\n\n\nEnhanced Bluetooth\u00ae\nconnection technology\ndesigned to reduce pairing\nhassles and simplify\nconnections for Windows\ndevices.\n\n\nSupports up to three\ndevices two via\n\n\nBluetooth connections,\none via the 2.4GHz USB\nreceiver with 128-bit\nAES government-grade\nencryption security.\nWindows and macOS\ncompatible.\n\n\nand wrist positioning for\nall-day ergonomic comfort\nand productivity.\n\n\n(4)\n500\nmAh\n\n\nF10 F11F12D\n\n\nE\n\n\nBackspace\n\n\nHome\n\n\nP\n\n\n]\n\n\n[\n\n\nO\n\n\nCustomizable Keys\n\n\nProductivity Enhancing\nKeys and Lights\n\n\nLong-Lasting Power\n\n\nThe built-in 500mAh\nrechargeable battery\nincludes a convenient\nUSB-C charging port and\ncable for uninterrupted\nproductivity.\n\n\nCustomize your experience\nby programming keys,\ncreating macros, adjusting\nkey mappings, managing\nprofiles, and more with\nthe optional Kensington\nKonnect\u2122 software.\n\n\nVideo conferencing keys\nsupport efficient meeting\ntransitions. Vibrant RGB\nindicators allow you\nto easily monitor your\nconnection status, caps\nlock, and battery life.\n\n\nKensington\n\n\nTechnical\nSpecifications\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nSystem Requirements\n\n\n1 x USB-A\n\n\nSystem Resource Requirement\n\n\nfor charging and wired connection.\nWindows, macOS\n\n\nOperating Systems Supported\n\n\nProduct Features\n\n\n2.4GHz, Bluetooth\n\n\nConnection\nMulti-device\n\n\nUp to 3 (2 x BT 5.2, 1 x 2.4GHz)\n\n\nTKL, 93 keys (89 + 4 meeting control buttons)\n\n\nKey Layout\n\n\nMeeting control button\nSwitches\n\n\n4 (Mute or Unmute/Video off or on/Share Screen/End call)\nMembrane\n\n\nOuter case and keycaps\n\n\n43% PCR ABS\n\n\nWrist pad\n\n\nYes (Non-Detachable)\n\n\nBacklight\n\n\n\u039d\u0391\n\n\nBattery\n\n\nRechargeable Battery\n\n\nKeyboard: USB-C / Host: USB-A\n\n\nCharging port\n\n\nCharging time\n\n\n3 hours to fully charged from low battery\nAt least 2.5 months after fully charged\nWindows and macOS\n\n\nBattery life\n\n\nKensington Konnect\u2122 Support\n\n\nProgrammable keys\nKeyboard Color\n\n\n20\n\n\nBlack\n\n\nProduct Dimensions\n\n\nKeyboard Unit Dimensions\n(Width x Length x Height)\nKeyboard Unit Weight\nKeyboard Color\n\n\n8.7 x 14.8 x 1.37\"\n\n\n(220.9 x 375.9 x 34.8mm)\n\n\n1.32lbs (0.6kg)\n\n\nBlack\n\n\nPackaging Dimensions\n\n\nPackaging Style\nPackaging Dimensions\n(Width x Length x Height)\nPackaging Weight\n\n\nRetail box (FSC) with inner brown box (FSC)\nUnit: 10.15 x 15.62 x 1.73\" (258 x 397 x 44mm)\nCarton: 11.5 x 16.3 x 8.07\" (292 x 414 x 205mm)\n\n\nUnit: 2.16lbs (0.98kg)\n\n\nMaster Carton: 9.92lbs (4.5kg)\n\n\nMaster Carton\nPackaging Content\n\n\n4pcs\n\n\n1 x Keyboard\n\n\n1 x 6' (1.8m) USB-C to USB-A cable\n\n\n1 x 2.4GHz Wireless receiver\n\n\n1 x Welcome card\n\n\n1 x Instruction guide\n\n\n1 x Battery warning insert\n\n\n1 x Compliance sheet\n\n\n1 x Warranty card\n\n\nKensington\n\n\nTechnical\nSpecifications\n\n\nPro Fit\u2122 Ergo KB675 EQ TKL Rechargeable Keyboard\n\n\nPart Number: K75491US | UPC Code: 0 85896 75491 6\n\n\nPower & Environmental\n\n\n500mAh rechargeable battery\n\n\nBattery\n\n\n32 to 104\u00b0F (0 to 40\u00b0C)\n\n\nOperating Temperature\n\n\n5 to 140\u00b0F (-15 to 60\u00b0C)\n\n\nStorage Temperature\n\n\nRelative Humidity\n\n\n0% to 85% non-condensing\n\n\nCountry of Origin, Regulatory, And Warranty\n\n\nCountry of Origin\n\n\nChina\n\n\n3-year Limited\n\n\nWarranty\n\n\nCertifications & Compliance\n\n\nFCC, IC, RCM, CB, China SRRC, RoHS, ACCO RSL, BC,\nNrCAN, CE, UKCA, WEEE, NOM, PSTI\n\n\nSustainability\n\n\nOuter Case and Keycaps\nPaper Packaging\n\n\nMade with 43% Post Consumer Recycled ABS\nFSC\u2122-certified\n\n\nKensington\n\n\nAll specifications are subject to change without notice. Products may not be available in all markets. Kensington and the ACCO\nname and design are registered trademarks of ACCO Brands. Kensington The Professionals' Choice is a trademark of ACCO Brands.\nAll other registered and unregistered trademarks are the property of their respective owners. 2024 Kensington Computer\nProducts Group, a division of ACCO Brands. All rights reserved.\n\n\nACCO\nBRANDS\n\n\nThe Professionals' Choice\u2122\n\n\nFOR MORE INFORMATION CONTACT: 1-855-692-0054 | sales@kensington.com\n"}, "expected_output": {"claims": [{"unit": "lbs", "value": 2.16, "evidence": ["Unit: 2.16lbs (0.98kg)", "Packaging Style\nPackaging Dimensions\n(Width x Length x Height)\nPackaging Weight"]}]}, "metadata": {"product_category": "Furniture & other goods", "request_id": "req_9a170a87c1e30ebc"}}
{"id": "f52b72aa2f78737a79088a78", "input": {"query": "What is the exact material composition (mass fractions or percentages) of plasticizer in Trosifol PVB film? Provide the specific values as mass fractions or percentages.", "source_url": "https://www.trosifol.com/fileadmin/user_upload/about_us/sustainability/environmental-product-decleration-as-per-iso-14025-and-en-15804_a2-pvb-film-trosifol.pdf", "document_text": "ENVIRONMENTAL PRODUCT DECLARATION\n\n\nas per ISO 14025 and EN 15804+A2\n\n\nOwner of the Declaration Kuraray Europe GmbH\n\n\nInstitut Bauen und Umwelt e.V. (IBU)\n\n\nPublisher\n\n\nInstitut Bauen und Umwelt e.V. (IBU)\n\n\nProgramme holder\n\n\nEPD-KUR-20230072-CCI1-EN\n\n\nDeclaration number\n\n\nIssue date\n\n\n12/05/2023\n\n\nValid to\n\n\n11/05/2028\n\n\nPVB film (Trosifol\u24c7)\nKuraray Europe GmbH\n\n\nInstitut Bauen\nund Umwelt e.V.\n\n\nwww.ibu-epd.com | https://epd-online.com\n\n\nECO PLATFORM\n\n\nEPD\n\n\nVERIFIED\n\n\nALIEK\n\n\nkuraray\n\n\nGeneral Information\n\n\nKuraray Europe GmbH\nProgramme holder\n\n\nPVB film (Trosifol\u24c7)\n\n\nOwner of the declaration\nKuraray Europe GmbH\nPhilipp-Reis-Str. 4\n\n\nIBU - Institut Bauen und Umwelt e.V.\n\n\nHegelplatz 1\n10117 Berlin\n\n\n65795 Hattersheim\nGermany\n\n\nGermany\n\n\nDeclaration number\n\n\nDeclared product / declared unit\n\n\nTrosifol\u00ae PVB.\n\n\nEPD-KUR-20230072-CCI1-EN\n\n\nThe declared unit is 1 m\u00b2.\n\n\nThis declaration is based on the product category rules:\n\n\nScope:\n\n\nTrosifol PVB, manufactured in Troisdorf based on Mowital\u24c7 resin from\nFrankfurt.\n\n\nPlate glass for construction and interlayers, 01/09/2022\n(PCR checked and approved by the SVR)\n\n\nThe owner of the declaration shall be liable for the underlying information\nand evidence; the IBU shall not be liable with respect to manufacturer\ninformation, life cycle assessment data and evidences.\n\n\nIssue date\n\n\n12/05/2023\n\n\nThe EPD was created according to the specifications of EN 15804+A2. In\nthe following, the standard will be simplified as EN 15804.\n\n\nValid to\n\n\nVerification\n\n\n11/05/2028\n\n\nThe standard EN 15804 serves as the core PCR\nIndependent verification of the declaration and data according to ISO\n14025:2011\n\n\ninternally \u2611\n\n\nexternally\n\n\nNam Paten\n\n\nDipl.-Ing. Hans Peters\n\n\n(Chairman of Institut Bauen und Umwelt e.V.)\n\n\nHam Peter\nDipl.-Ing. Hans Peters\n\n\n\u041c\u0438\u043b\u0438\n\n\nDr. Matthew Fishwick,\n(Independent verifier)\n\n\n(Managing Director Institut Bauen und Umwelt e.V.)\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\nkuraray\n\n\nProduct\n\n\nschueren/\n\n\nProduct description/Product definition\n\n\nConstructional data\n\n\nKuraray's polyvinyl butyral (PVB) thermoplastic films are tough,\nresilient safety interlayers used in laminated architectural safety\nglass. These Trosifol\u24c7 PVB interlayers offer safety advantages\nby retaining dangerous shards in case of glass breakage. They\nare commonly used as safety glass interlayers available\nworldwide.\n\n\nUnit\n\n\nValue\n\n\nName\n\n\nRefractive index acc. to DIN EN ISO 489\n\n\n1.48\n\n\n0.21 W/mK\n\n\nThermal conductivity acc. to DIN EN 993-15\nThermal expansion coefficient acc. to ISO\n11359-2\n\n\n0.00017 K-1\n\n\nSpecific heat capacity\n\n\n1.9 kJ/kgK\n1E+13 \u03a9\n\n\nThis EPD covers all Trosifol\u24c7 PVB products produced in\nTroisdorf based on Mowital\u24c7 resin produced in Frankfurt.\nProduct codes Trosifol\u24c7 B2XX and Trosifol B8XX.\n\n\nSurface resistivity acc. to DIN 53482\nTensile strength acc. to ISO 527-3\nElongation at break acc. to ISO 527-3\n\n\nN/mm\u00b2\n\n\n20\n\n\n250\n\n\n%\n\n\nTg acc. to DMA, 3K/min, 1Hz\n\n\nFor the use and application of the product, the respective\nnational provisions at the place of use apply, in Germany, for\nexample, the building codes of the federal states and the\ncorresponding national specifications.\n\n\n\u00b0C\n\n\n32\n\n\nPerformance data of the product with respect to its\ncharacteristics in accordance with the relevant technical\nprovision (no CE-marking).\n\n\nApplication\n\n\nBase materials/Ancillary materials\n\n\nTrosifol\u24c7 PVB film needs to be laminated between two pieces\nof glass. This sandwich arrangement is called laminated safety\nglass according to EN ISO 14449. Special Trosifol\u24c7 PVB\ngrades offer additional decorative, acoustic, UV managing and\nstructural properties.\n\n\nThe main constituents of Trosifol\u24c7 PVB film are (in mass\npercentages):\n\n\n-PVB resin ~72 %\n\n\n-Plasticizer ~27-28 %\n\n\n-Additives and water <1%\n\n\n1) This product contains substances listed in the candidate list\n(date: 17.01.2023) exceeding 0.1 percentage by mass: no\n\n\nTechnical Data\n\n\nFor calculating the light, solar and heat parameters of glazing\nspecifically containing films from the Trosifol\u24c7 & SentryGlas\u24c7\nproduct range, please go to:\n\n\n2) This product contains other Carcinogenic, Mutagenic,\nReprotoxic (CMR) substances in categories 1A or 1B which are\nnot on the candidate list, exceeding 0.1 percentage by mass:\n\n\nhttps://www.trosifol.com/winslt-tool/\n\n\nno\n\n\nSound Control data can be found here:\n\n\n3) Biocide products were added to this construction product or it\nhas been treated with biocide products (this then concerns a\ntreated product as defined by the (EU) Ordinance on Biocide\nProducts No. 528/2012): no\n\n\nhttps://www.trosifol.com/soundlab-ai/\n\n\nThe following data are valid for Trosifol\u24c7 Clear / Trosifol\u24c7\nUltraClear. Other product's data can be found in our laminator\nbrochure:\n\n\nReference service life\n\n\nThe reference service life is typically determined by the glass\nand not by the interlayer.\n\n\nhttps://www.trosifol.com/de/salessupport/downloads/produktbro\n\n\nLCA: Calculation rules\n\n\nproducts and energy, as well as waste processing up to the\nend-of-waste state or disposal of final residues during the\nproduct stage.\n\n\nDeclared Unit\n\n\nThis declaration refers to the declared unit of 1 m\u00b2 of PVB film\n(Trosifol\u24c7). The grammage of the PVB film is 0.775 kg/m\u00b2.\n\n\nThese modules consider the manufacturing of system\ncomponents/raw materials, the transport to the production site\nand the production processes of the products under study. The\nimpact of packaging materials is included.\n\n\nDeclared unit - PVB film (Trosifol\u24c7)\n\n\nName\n\n\nValue\n\n\nUnit\nm\u00b2\nkg/m^2\n\n\nDeclared unit\n\n\n1\n0.775\n0.00076\n1.07\n\n\nGrammage\n\n\nModule A5:\n\n\nLayer thickness\n\n\nm\n\n\nTreatment and disposal of packaging material. Credits for\npotential avoided burdens due to energy substitution of\nelectricity and thermal energy generation are declared in\nmodule D and affect only the rate of primary material (no\nsecondary materials).\n\n\nDensity\n\n\ng/cm^3\n\n\nSystem boundary\n\n\nThe type of EPD is cradle-to-gate with options, modules C1-\nC4, and module D (A1-A3, C, D and additional module A5). In\nthe following section, a detailed description of the specific\nsystem boundaries is given:\n\n\nModule C1 to C4:\n\n\nThe end-of-life scenarios are as follows:\n\n\nC1 Deconstruction/demolition: Dismantling is manual\n(no environmental burden).\n\n\n\u2022\n\n\nModule A1 to A3:\n\n\nThe product stage includes the provision of all materials,\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n2\n\n\nkuraray\n\n\n\u2022 C2 - Transport to treatment/disposal site: Average\ntransport distance from the demolition site to waste\ntreatment is assumed as 50 km to the landfill.\n\n\nLand or region, in which the declared product system is\nmanufactured, used or handled at the end of the product's\nlifespan: Europe\n\n\n\u2022 C3 - Waste processing for reuse, recovery or recycling:\nNo waste processing (no environmental burden).\n\u2022 C4 - Disposal: PVB film is 100% landfilled.\nModule D:\n\n\nComparability\n\n\nBasically, a comparison or an evaluation of EPD data is only\npossible if all the data sets to be compared were created\naccording to EN 15804 and the building context, respectively\nthe product-specific characteristics of performance, are taken\ninto account. Background datasets: GaBi ts 10.6 software\nsystem and GaBi Professional 2022.1 LCI database.\n\n\nFor the thermal and electrical energy generated in Module A5\ndue to the thermal treatment of packaging and product waste,\navoided burdens have been calculated by the inversion of the\nelectricity grid mix and thermal energy from natural gas, using\nEuropean datasets.\n\n\nGeographic Representativeness\n\n\nLCA: Scenarios and additional technical information\n\n\nEnd of life (C1-C4)\n\n\nCharacteristic product properties biogenic carbon\nBiogenic carbon is only present in the packaging (wooden\npallets and cartons).\n\n\nThe end-of-life scenarios are as follows:\n\n\nC1 The deconstruction of the PVB film is assumed to be done\nmanually. Therefore, no environmental loads for the dismantling\nof this product are considered.\nC2 - Transport to treatment/disposal site: Average transport\ndistance from the demolition site to waste treatment is assumed\nas 50 km to landfill.\n\n\nAssumed water content in wooden pallets (packaging): 18 %.\nAssumed carbon content: dry wood mass consists of 50 %\nbiogenic carbon and paper/cardboard 43 %.\n\n\nC4 - Disposal: The PVB film is 100% landfilled.\nName\n\n\nThe biogenic carbon content of the packaging is thus: 0.129 kg\npallet/declared unit * 0.82 * 0.5 kg C / kg pallet (abs. dry) +\n0.082 kg *0.43 kg C/kg cardboard= 0.08815 kg C/declared unit.\nInformation on describing the biogenic Carbon Content at\nfactory gate\n\n\nValue Unit\n\n\nCollected as mixed construction waste [PVB film\nper FU]\n\n\n0.775 kg\n\n\n0.775 kg\n\n\nLandfilling [PVB film per FU]\n\n\nName\n\n\nValue Unit\n0.08815\n\n\nBiogenic carbon content in accompanying\npackaging\n\n\nkg\nC\n\n\nReuse, recovery and/or recycling potentials (D), relevant\nscenario information\n\n\nFor the thermal and electrical energy generated in Module A5\ndue to the thermal treatment of packaging, avoided burdens\nhave been calculated by the inversion of the electricity grid mix\nand thermal energy from natural gas, using European datasets.\n\n\nThe following technical scenario information is required for the\ndeclared modules.\n\n\nInstallation into the building (A5)\n\n\nThe packaging material treatment and disposal are also\nconsidered in module A5.\n\n\nValue Unit\n\n\nName\n\n\nOutput substances following waste treatment on\nsite [packaging materials per FU]\n\n\n0.2301 kg\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n3\n\n\nkuraray\n\n\nLCA: Results\n\n\nDESCRIPTION OF THE SYSTEM BOUNDARY (X = INCLUDED IN LCA; ND = MODULE OR INDICATOR NOT DECLARED; MNR =\nMODULE NOT RELEVANT)\n\n\nBenefits and\nloads beyond\nthe system\nboundaries\n\n\nConstruction\nprocess stage\n\n\nProduct stage\n\n\nUse stage\n\n\nEnd of life stage\n\n\nOperational energy\n\n\nTransport from the\n\n\nOperational water\n\n\n2 Waste processing\nC3\n\n\nC1\n2 De-construction\n\n\nRefurbishment\n\n\n\u2611 gate to the site\nA4\n\n\nManufacturing\n\n\nRaw material\n\n\nReplacement\n\n\nMaintenance\n\n\nRecovery-\n\n\nRecycling-\n\n\ndemolition\n\n\nAssembly\n\n\nTransport\n\n\nTransport\n\n\nDisposal\n\n\npotential\n\n\nReuse-\n\n\nsupply\n\n\nRepair\n\n\nUse\n\n\nuse\n\n\nuse\n\n\nDX\n\n\nA2\n\n\nA3\n\n\nA5\n\n\nC2\n\n\nC4\n\n\nA1\n\n\nB1\n\n\nB2\n\n\nB3\n\n\nB4\n\n\nB5\n\n\nB6\n\n\nB7\n\n\nMND X MND MND MNR MNR MNR MND MND\n\n\n\u2717\n\n\n\u2717\n\n\nX\n\n\n\u2717\n\n\nRESULTS OF THE LCA - ENVIRONMENTAL IMPACT according to EN 15804+A2: 1 m\u00b2 PVB film (Trosifol\u24c7)\nParameter\n\n\nUnit\n\n\nA1-A3\n\n\nA5\n\n\nC2\n\n\nC3\n\n\nC1\n\n\nC4\n\n\nD\n\n\nkg CO2 eq\n\n\n3.26E+00\n\n\n2.37E-03\n\n\n-1.19E-01\n\n\nGWP-total\n\n\n3.77E-01\n\n\n0\n\n\n0\n\n\n5.49E-02\n\n\nkg CO2 eq\n\n\nGWP-fossil\n\n\n2.37E-03\n\n\n3.55E+00\n\n\n5.77E-02\n\n\n0\n\n\n0\n\n\n5.49E-02\n\n\n-1.18E-01\n\n\nGWP-biogenic\n\n\nkg CO2 eq\n\n\n3.2E-01\n\n\n-2.95E-01\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n-6.05E-04\n\n\nkg CO2 eq\n\n\n9.09E-06\n\n\nGWP-luluc\n\n\n5.4E-04\n\n\n5.85E-06\n\n\n0\n\n\n2.67E-05\n\n\n-1.3E-05\n\n\n0\n\n\nkg CFC11 eq\n\n\n1.91E-09\n\n\nODP\n\n\n5.63E-14\n\n\n0\n\n\n4.84E-16\n\n\n7.38E-14\n\n\n-7.99E-13\n\n\n0\n\n\n4.67E-03\n\n\nmol H+ eq\n\n\n1.63E-04\n\n\nAP\n\n\n7.13E-05\n\n\n0\n\n\n7.48E-06\n\n\n0\n\n\n-1.56E-04\n\n\nEP-freshwater\n\n\nkg P eq\nkg N eq\nmol N eq\nkg NMVOC\neq\nkg Sb eq\nMJ\nm\u00b3 world eq\ndeprived\n\n\n-1.63E-07\n\n\n2.99E-05\n\n\n1.62E-08\n\n\n0\n\n\n4.71E-09\n\n\n0\n\n\n1.03E-05\n\n\nEP-marine\nEP-terrestrial\n\n\n3.5E-06\n\n\n-4.22E-05\n\n\n1.47E-03\n\n\n2.45E-05\n\n\n0\n\n\n3.6E-05\n\n\n0\n\n\n3.28E-04\n\n\n1.64E-02\n\n\n3.95E-04\n\n\n0\n\n\n3.9E-05\n\n\n0\n\n\n-4.52E-04\n\n\n6.56E-05\n\n\nPOCP\n\n\n5.78E-03\n\n\n0\n\n\n6.82E-06\n\n\n0\n\n\n1.16E-04\n\n\n-1.18E-04\n\n\n-1.78E-08\n\n\nADPE\n\n\n1.61E-06\n\n\n1.45E-09\n\n\n0\n\n\n2.36E-10\n\n\n0\n\n\n3.81E-09\n\n\nADPF\n\n\n7.95E+01\n\n\n1.39E-01\n\n\n0\n\n\n3.13E-02\n\n\n7.79E-01\n\n\n-2.01E+00\n\n\n0\n\n\nWDP\n\n\n3.87E-02\n\n\n-7.12E-01\n\n\n0\n\n\n1.01E-05\n\n\n0\n\n\n-5.42E-04\n\n\n-1.26E-02\n\n\nGWP = Global warming potential; ODP = Depletion potential of the stratospheric ozone layer; AP = Acidification potential of land and water; EP =\nEutrophication potential; POCP = Formation potential of tropospheric ozone photochemical oxidants; ADPE = Abiotic depletion potential for non-fossil\nresources; ADPF = Abiotic depletion potential for fossil resources; WDP = Water (user) deprivation potential)\nRESULTS OF THE LCA - INDICATORS TO DESCRIBE RESOURCE USE according to EN 15804+A2: 1 m\u00b2 PVB film (Trosifol\u24c7)\nParameter\n\n\nA5\n45\n3.42E+00\n\n\nA1-A3\n1.03E+01\n\n\nUnit\nMJ\n\n\nC1\n\n\nC2\n\n\nC4\n\n\nD\n\n\nC3\n\n\n2.06E-03\n\n\nPERE\n\n\n0\n\n\n6.41E-02\n0\n\n\n-5.52E-01\n\n\n0\n\n\n3.39E+00\n\n\nPERM\n\n\n-3.39E+00\n\n\n0\n\n\nMJ\n\n\n0\n\n\n0\n\n\n0\n-5.52E-01\n-2.01E+00\n0\n-2.01E+00\n\n\n|PERT\n\n\n1.37E+01\n\n\nMJ\n\n\n3.22E-02\n\n\n0\n\n\n2.06E-03\n\n\n0\n\n\n6.41E-02\n\n\n2.5E+01\n\n\nPENRE\n\n\nMJ\n\n\n5.47E+01\n\n\n6.94E-01\n\n\n0\n\n\n3.13E-02\n\n\n0\n\n\n-2.43E+01\n\n\nPENRM\n\n\n2.48E+01\n\n\n-5.55E-01\n\n\n0\n\n\n0\n\n\n0\n\n\nMJ\n\n\nPENRT\n\n\n7.96E+01\n\n\n1.39E-01\n\n\n0\n\n\n3.13E-02\n\n\n0\n\n\n7.8E-01\n\n\nMJ\n\n\nSM\n\n\nkg\nMJ\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nRSF\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nNRSF\n\n\nMJ\nm\u00b3\n\n\n0\n1.64E-02\n\n\n0\n\n\n0\n\n\n0\n0\n\n\n0\n\n\n0\n\n\n0\n9.16E-04\n\n\n1.6E-06\n\n\n1.01E-05\n\n\n-5.31E-04\n\n\nFW\n\n\n0\n\n\nPERE = Use of renewable primary energy excluding renewable primary energy resources used as raw materials; PERM = Use of renewable primary\nenergy resources used as raw materials; PERT = Total use of renewable primary energy resources; PENRE = Use of non-renewable primary energy\nexcluding non-renewable primary energy resources used as raw materials; PENRM = Use of non-renewable primary energy resources used as raw\nmaterials; PENRT = Total use of non-renewable primary energy resources; SM = Use of secondary material; RSF = Use of renewable secondary fuels;\nNRSF = Use of non-renewable secondary fuels; FW = Use of net fresh water\n\n\nRESULTS OF THE LCA \u2013 WASTE CATEGORIES AND OUTPUT FLOWS according to EN 15804+A2:\n\n\n1 m\u00b2 PVB film (Trosifol\u00ae)\n\n\nParameter\n\n\nUnit\n\n\nA1-A3\n8.63E-08\n\n\nA5\n1.17E-11\n\n\nC3\n\n\nC1\n\n\nC2\n\n\nC4\n\n\nD\n\n\n1.2E-10\n7.72E-01\n9.58E-06\n\n\nkg\n\n\n0\n\n\n1.37E-13\n\n\n-2.73E-10\n\n\nHWD\n\n\n0\n\n\n-1.02E-03\n\n\nkg\n\n\n3.89E-02\n\n\n2.29E-02\n\n\n0\n\n\n5.1E-06\n\n\nNHWD\n\n\n0\n\n\n3.94E-08\n\n\nkg\n\n\n7.58E-04\n\n\n6.07E-06\n\n\n0\n\n\n0\n\n\n-1.58E-04\n\n\nRWD\n\n\nCRU\n\n\nkg\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nMFR\n\n\nkg\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nMER\n\n\nkg\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n5.32E-01\n\n\nEEE\n\n\nMJ\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nEET\n\n\nMJ\n\n\n0\n\n\n9.58E-01\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n4\n\n\nkuraray\n\n\nHWD = Hazardous waste disposed; NHWD = Non-hazardous waste disposed; RWD = Radioactive waste disposed; CRU = Components for re-use;\nMFR = Materials for recycling; MER = Materials for energy recovery; EEE = Exported electrical energy; EET = Exported thermal energy\n\n\nRESULTS OF THE LCA - additional impact categories according to EN 15804+A2-optional:\n\n\n1 m\u00b2 PVB film (Trosifol\u00ae)\n\n\nParameter\n\n\nUnit\nDisease\nincidence\n\n\nA1-A3\n\n\nA5\n\n\nC1\n\n\nC2\n\n\nC3\n\n\nC4\n\n\nD\n\n\nPM\n\n\n6.26E-08\n\n\n5.26E-10\n\n\n0\n\n\n4.26E-11\n\n\n0\n\n\n1.56E-09\n\n\n-1.29E-09\n\n\nIR\n\n\nkBq U235 eq\nCTUe\n\n\n7.89E-02\n\n\n9.02E-04\n\n\n0\n\n\n0\n\n\n1.41E-03\n\n\n-2.68E-02\n\n\n4E-06\n\n\n8.63E-02\n\n\n7.63E-01\n\n\nETP-fw\n\n\n6.53E+01\n\n\n0\n\n\n2.4E-02\n\n\n0\n\n\n-4.41E-01\n\n\n-2.03E-11\n\n\nHTP-C\nHTP-nc\n\n\nCTUh\n\n\n9.32E-09\n\n\n4.13E-12\n\n\n0\n\n\n4.81E-13\n\n\n0\n\n\n3.42E-11\n\n\nCTUh\nSQP\n\n\n4.94E-08\n4.91E+01\n\n\n2.77E-11\n\n\n2.79E-10\n3.71E-02\n\n\n0\n\n\n0\n\n\n2.87E-09\n\n\n-7.8E-10\n\n\nSQP\n\n\n0\n\n\n9.38E-03\n\n\n0\n\n\n5.61E-02\n\n\n-3.59E-01\n\n\nPM = Potential incidence of disease due to PM emissions; IR = Potential Human exposure efficiency relative to U235; ETP-fw = Potential comparative\nToxic Unit for ecosystems; HTP-c = Potential comparative Toxic Unit for humans (cancerogenic); HTP-nc = Potential comparative Toxic Unit for humans\n(not cancerogenic); SQP = Potential soil quality index\n\n\nDisclaimer 1 - for the indicator \"Potential Human exposure efficiency relative to U235\". This impact category deals mainly with the\neventual impact of low-dose ionizing radiation on human health of the nuclear fuel cycle. It does not consider effects due to possible\nnuclear accidents, occupational exposure or radioactive waste disposal in underground facilities. Potential ionizing radiation from the\nsoil, radon and from some construction materials is also not measured by this indicator.\n\n\nDisclaimer 2 - for the indicators 'abiotic depletion potential for non-fossil resources', 'abiotic depletion potential for fossil resources',\n'water (user) deprivation potential, deprivation-weighted water consumption', 'potential comparative toxic unit for ecosystems', 'potential\ncomparative toxic unit for humans - cancerogenic', 'Potential comparative toxic unit for humans - not cancerogenic', 'potential soil\nquality index'. The results of this environmental impact indicator shall be used with care as the uncertainties on these results are high\nas there is limited experience with the indicator.\n\n\nReferences\n\n\nEN ISO 14025:2011, Environmental labels and declarations \u2014\nType III environmental declarations - Principles and\nprocedures.\n\n\nStandards\n\n\nDIN 53482\n\n\nDIN 53482:1967-01, Testing of Insulating Materials;\nDetermination of Electrical Resistances Values.\n\n\nFurther References\n\n\nDIN EN ISO 489\n\n\nDIN EN ISO 489:1999-08, Plastics - Determination of the\nrefractive index (ISO 489:1999).\n\n\nCandidate list\n\n\nCandidate List of substances of very high concern for\nAuthorisation, published on ECHA website, latest version\n17.01.2023 (https://echa.europa.eu/candidatelist-table)\n\n\nDIN EN 993-15\n\n\nDIN EN 993-15:2005-07, Methods of test for dense shaped\nrefractory products - Part 15: Determination of thermal\nconductivity by the hot-wire (parallel) method.\n\n\nGaBi\n\n\nGaBi Software System and Database for Life Cycle\nEngineering, 1992-2021, Sphera Solutions GmbH, Leinfelden-\nEchterdingen, with acknowledgement of\nLBP University of Stuttgart, program version GaBi 10; database\nversion 2022.1.\n\n\nDIN EN ISO 527-3\n\n\nDIN EN ISO 527-3:2019-02, Plastics - Determination of tensile\nproperties - Part 3: Test conditions for films and sheets (ISO\n527-3:2018).\n\n\nGaBi documentation\n\n\nGaBi dataset documentation for the software system and\ndatabases, LBP, University of Stuttgart and Sphera Solutions\nGmbH, Leinfelden-Echterdingen,\n\n\nEN 15804\n\n\nEN15804:2012+A1:2013, Sustainability of construction works\n- Environmental Product Declarations Core rules for the\nproduct category of construction products.\n\n\n-\n\n\n2021. (http://www.gabi-\n\n\nsoftware.com/support/gabi/gabi[1] database-2021-lci-\n\n\ndocumentation/)\n\n\nEN 15804\n\n\nEN 15804:2012+A2:2019+AC:2021, Sustainability of\nconstruction works - Environmental Product Declarations -\nCore rules for the product category of construction products.\n\n\nIBU 2021\n\n\nInstitut Bauen und Umwelt e.V.: General Instructions for the\nEPD programme of Institut Bauen und Umwelt e. V., Version\n2.0, Berlin: Institut Bauen und Umwelt e. V., 2021 HYPERLINK\n\"http://www.ibu-epd.\" www.ibu-epd.com\n\n\nEN ISO 14449\n\n\nEN 14449:2005/AC:2005, Glass in building - Laminated glass\nand laminated safety glass - Evaluation of conformity/Product\nstandard.\n\n\nOrdinance on Biocide Products No. 528/2012\nRegulation (EU) No 528/2012 of the European Parliament and\nof the Council of 22 May 2012 concerning the making available\non the market and use of biocidal products\n\n\nISO 11359-2\n\n\nISO 11359-2:2021-11, Plastics - Thermomechanical analysis\n(TMA) - Part 2: Determination of coefficient of linear thermal\nexpansion and glass transition temperature.\n\n\nPCR Part A\n\n\nPCR Part A: Calculation rules for the Life Cycle Assessment\nand Requirements on the Background Report according to EN\n\n\nISO 14025\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n5\n\n\nkuraray\n\n\n15 804+A2:2019, Version 1.3, Institut Bauen und Umwelt e.V.,\n2020.\n\n\nProduct Category Rules for Building Products, Part B:\n\n\nRequirements on the EPD for plate glass for construction and\ninterlayers, version 1.6, 2022 www.bau-umwelt.de\n\n\nPCR Part B\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n6\n\n\nkuraray\n\n\nPublisher\n\n\nInstitut Bauen und Umwelt e.V.\n\n\n+49 (0)30 3087748-0\ninfo@ibu-epd.com\nwww.ibu-epd.com\n\n\nHegelplatz 1\n10117 Berlin\nGermany\n\n\nInstitut Bauen\nund Umwelt e.V.\n\n\nProgramme holder\n\n\n+49 (0)30 3087748-0\ninfo@ibu-epd.com\nwww.ibu-epd.com\n\n\nInstitut Bauen und Umwelt e.V.\nHegelplatz 1\n\n\n10117 Berlin\nGermany\n\n\nInstitut Bauen\nund Umwelt e.V.\n\n\nAuthor of the Life Cycle Assessment\n\n\nSphera Solutions GmbH\n\n\n+49 711 341817-0\ninfo@sphera.com\nwww.sphera.com\n\n\nsphera\u24c7\n\n\nHauptstra\u00dfe 111-113\n\n\n70771 Leinfelden-Echterdingen\nGermany\n\n\nOwner of the Declaration\n\n\nKuraray Europe GmbH\nPhilipp-Reis-Str. 4\n\n\n+49 69 305 85 300\ntrosifol@kuraray.com\nhttps://www.kuraray.eu/\n\n\nkuraray\n\n\n65795 Hattersheim\nGermany\n\n\nEnvironmental-Product Declaration - Kuraray Europe GmbH - PVB film (Trosifol\u24c7)\n\n\n7\n"}, "expected_output": {"claims": [{"unit": "%", "value": 27, "evidence": ["The main constituents of Trosifol\u24c7 PVB film are (in mass percentages):", "-Plasticizer ~27-28 %"]}, {"unit": "%", "value": 28, "evidence": ["The main constituents of Trosifol\u24c7 PVB film are (in mass percentages):", "-Plasticizer ~27-28 %"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_4a571861e670e7da"}}
{"id": "d4557cc7a4d7e001a7d88eb6", "input": {"query": "What is the typical mass percentages of Al2O3 in standard float glass composition? Answer in %.", "source_url": "https://www.pnnl.gov/main/publications/external/technical_reports/PNNL-13958.pdf", "document_text": "PNNL-13958\n\n\nPacific Northwest\nNational Laboratory\nOperated by Battelle for the\nU.S. Department of Energy\n\n\nThermochemical Optimization of\nFloat Glass Composition:\nLow-Alumina Glass Development\n\n\nAugust 2002\n\n\nPrepared for the U.S. Department of Energy\nunder Contract DE-AC06-76RL01830\n\n\nDEPARTMENT\n\n\nENERGY\n\n\n1. UNITED\n\n\nUNITED STA\n\n\nOF AMERIC\n\n\nDSTAT\n\n\nDISCLAIMER\n\n\nThis report was prepared as an account of work sponsored by an agency of the\nUnited States Government. Neither the United States Government nor any\nagency thereof, nor Battelle Memorial Institute, nor any of their employees,\nmakes any warranty, express or implied, or assumes any legal liability or\nresponsibility for the accuracy, completeness, or usefulness of any\ninformation, apparatus, product, or process disclosed, or represents that\nits use would not infringe privately owned rights. Reference herein to any\nspecific commercial product, process, or service by trade name, trademark,\nmanufacturer, or otherwise does not necessarily constitute or imply its\nendorsement, recommendation, or favoring by the United States Government\nor any agency thereof, or Battelle Memorial Institute. The views and opinions\nof authors expressed herein do not necessarily state or reflect those of the\nUnited States Government or any agency thereof.\n\n\nPACIFIC NORTHWEST NATIONAL LABORATORY\n\n\noperated by\nBATTELLE\n\n\nfor the\n\n\nUNITED STATES DEPARTMENT OF ENERGY\n\n\nunder Contract DE-ACO6-76RLO1830\n\n\nPrinted in the United States of America\n\n\nAvailable to DOE and DOE contractors from the\nOffice of Scientific and Technical Information,\nP.O. Box 62, Oak Ridge, TN 37831-0062;\nph: (865) 576-8401\nfax: (865) 576-5728\nemail: reports@adonis.osti.gov\n\n\nAvailable to the public from the National Technical Information Service,\nU.S. Department of Commerce, 5285 Port Royal Rd., Springfield, VA 22161\nph: (800) 553-6847\nfax: (703) 605-6900\nemail: orders@ntis.fedworld.gov\nonline ordering: http://www.ntis.gov/ordering.htm\n\n\nThis document was printed on recycled paper.\n(8/00)\n\n\nPNNL-13958\n\n\nThermochemical Optimization of Float Glass\nComposition: Low-Alumina Glass Development\n\n\nP. R. Hrma\n\n\nD. E. Smith\n\n\nJ. D. Yeager\n\n\nO. P. Lam\n\n\nAugust 2002\n\n\nPrepared for the U.S. Department of Energy\nunder Contract DE-AC06-76RL01830\n\n\nPacific Northwest National Laboratory\n\n\nRichland, Washington 99352\n\n\nAbstract\n\n\nThe liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to\n74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to\n14.2 mass% Na2O. Crystalline phases were identified at 900\u00b0C as cristobalite, wollastonite, and devitrite.\nThe primary phases were tridymite and wollastonite. Partial specific Ts were obtained from the data and\ncompared with the literature.\n\n\nE:\n\n\nContents\n\n\nAbstract....\n\n\n111\n\n\nIntroduction...\n\n\nExperimental..\n\n\n5\n\n\nResults\n\n\n7\n\n\n13\n\n\nDiscussion.\n\n\nConclusions..\n\n\n17\n\n\nReferences...\n\n\n19\n\n\nFigures\n\n\nFigure 1. Map of Tridymite Liquidus Surface in Na\u2082O-CaO-MgO-SiO2 System; Shahid and Glasser\n(1972), reproduced in Roth et al. (1981)..\n\n\n1\n\n\nFigure 2. XRD Pattern of Low-Al2O3-Low-Na2O Glass Heat-Treated at 900\u00b0 C for 24 h.\nFigure 3. Crystals of Cristobalite in Low-Al2O3-Low-CaO Glass, 1020\u00b0C...\nFigure 4. Crystals of Cristobalite in Low-Al2O3-Low-MgO Glass, 1000\u00b0C.\nFigure 5. Crystals of Devitrite in Low-Al2O3-High-CaO Glass, 970\u00b0C\nFigure 6. Crystals of Devitrite in Low-Al2O3-High-MgO Glass, 965\u00b0C\nFigure 7. Crystals of Wollastonite in Low-Al2O3-High-CaO Glass, 995\u00b0C.\nFigure 8. Crystals of Wollastonite in High-Al2O3-High-Na\u2082O Glass, 965\u00b0C.\n\n\n7\n\n\n9\n\n\n9\n\n\n10\n\n\n10\n\n\nFigure 9. Crystals of Wollastonite, Nucleated on a Bubble, in Low-Al2O3-High-MgO Glass, 995\u00b0C... 10\n\n\nFigure 10. Crystals of Wollastonite and Probably Devitrite, Nucleated on a Bubble, in High-Al2O3-\nBaseline Glass, 975\u00b0C\n\n\n10\n\n\nFigure 11. Crystals of Devitrite and Wollastonite in High-Al2O3-High-MgO Glass, 950\u00b0C.\n\n\n11\n\n\nFigure 12. The Effect of CaO on the Maximum Temperature at which Tridymite, Devitrite, and\nWollastonite occur in Float Glass at Equilibrium......\n\n\n13\n\n\nFigure 13. The Effect of MgO on the Maximum Temperature at which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium.....\n\n\n13\n\n\nFigure 14. The Effect of Na\u2082O on the Maximum Temperature at Which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium.\n\n\n14\n\n\nFigure 15. T;, Calculated vs. Measured.\n\n\n14\n\n\nFigure 16. Comparison of Experimental T; Values for VGS Glass with Babcock's TL Values for Soda-\nLime Glass..\n\n\n15\n\n\nFigure 17. Comparison of Experimental T Values for VGS Glass with T Values Predicted by Various\nModels\n\n\n15\n\n\nTables\n\n\nTable 1. Component Coefficients TL in \u00b0C\n\n\n2\n\n\nTable 2. Component Coefficients \u03c0 in \u00b0C in the Nonlinear Model by Backman et al. (a)\nTable 3. Composition Regions for VGS Glasses and for Published Models (mass%)(a).\nTable 4. Test Glass Compositions in Mass% of Oxides\n\n\n3\n\n\n5\n\n\nTable 5. Source Chemicals\n\n\nTable 6. Mass Fractions of Crystalline Phases in Glasses Heat-Treated at 900\u00b0 C for 24 h...\n\n\n8\n\n\nTable 7. Maximum Temperatures (\u00b0C) at Which Individual Crystalline Phases Occurred in VGS\nGlasses (a)(b)\n\n\n8\n\n\nTable 8. Component Coefficients (T;;) for the Maximum Temperatures at which Tridymite, Devitrite,\nWollastonite Occur in Float Glasses at Equilibrium (in \u00b0C)(a).\n\n\n13\n\n\nIntroduction\n\n\nSubstantial savings can be achieved if low-alumina float glass is produced. It has been proposed to\ndecrease the Al2O3 content in the float glass produced by the VGS Glass Systems (VGS) from the original\n0.45 mass% to 0.10 mass%. However, the removal of alumina from the current composition must not\nnegatively impact important glass properties, such as viscosity, liquidus temperature (TL), and chemical\ndurability.\n\n\nFloat glass is essentially a four-component mixture within the SiO2-Na2O-CaO-MgO system with\nminor addition, introduced either deliberately (Al2O3, Se2O3, C03O4) or as impurities (Fe2O3, K\u2082O). An\nimportant processability requirement is that the glass must avoid the tridymite primary phase field.\nChemical durability and thermal expansion requirements then place the composition to the neighborhood\nof the devitrite-wollastonite boundary\u2014see Figure 1. When Al2O3 and Fe2O3 are present in the mixture,\nthis boundary shifts towards the Na2O corner of the diagram because Al2O3 and Fe2O3 bound Na\u2082O for\ncharge compensation (this allows these oxides to be glass formers).\n\n\n(CaO)\n\n\n(CaO)\n\n\nSiO2\n\n\n61.4\n\n\n(SiO\u2082-Saturated)}\n\n\n1320\n\n\n-1300-\n\n\n(SiO2-Saturated)\n\n\nB\n\n\n1200-\n\n\n>\n\n\nCaSiO3\n\n\n62.5\n\n\nCaSiO3\n\n\n1375%\n\n\n73 73\n\n\n1035\u00b0\n\n\n---\n\n\n-1100-\n\n\nK\n\n\nDiopside\n\n\nDiopside\n\n\nNa2C03S16016\n\n\n-74\nLNa2CO3Si6016\n\n\n74\n\n\n980\u00b0\n\n\n827\u00b0\n\n\n900\n\n\n74 D\n\n\n75\n\n\n780\u00b0\n11:5\nC77271\nA 710\u00b0\n\n\nProtoenstatite\n\n\nProtoenstatite\n\n\n::5 74\n73 C\n75 75\n11:43:8\n\n\n755\u00b0\n\n\n995\u00b0\n\n\n1:5:12\n77\n\n\n800\u00b0\n\n\nM\n\n\n750\n\n\nNo\u2082Mg2SiO15\n\n\n1:5:12\n\n\nNa2Mg2Si6O15\n\n\n76 75 74.\n\n\n18:4\n\n\n3:8\n\n\n78.0\n\n\n77.0\n\n\n74\n\n\n74\n\n\n74\n\n\n1250\u00b0 1018\u00b0\n\n\n780\u00b0\n\n\n733\u00b0\n\n\n(MgO)\n\n\n(No\u2082O) (MgO)\n\n\n(Na\u2082O)\n\n\nWt %\n\n\nWt%\n\n\nFigure 1. Map of Tridymite Liquidus Surface in Na\u2082O-CaO-MgO-SiO2 System;\nShahid and Glasser (1972), reproduced in Roth et al. (1981)\n\n\nOptimizing glass composition with respect glass properties is done most effectively using glass\nproperty-composition relationship in the form of simple models. For small composition regions, these\nmodels have usually a form of linear functions. Thus, the maximum temperatures (T;) at which j-th\ncrystals are at equilibrium with glass can be expressed as a function of glass composition using the\nfollowing relationship:\n\n\nN\n\n\nT = \u03a3\u03a4\n\n\n(1)\n\n\nj\n\n\ni=1\n\n\n1\n\n\nwhere Tji is the i-th component j-th phase coefficient, x; is the i-th component mass fraction, and N is the\nnumber of components. If N is smaller then the actual number of glass components, mass fractions are\nnormalized in such a way that\n\n\nN\n\n\n\u03a3 x = 1\n\n\n(2)\n\n\ni=1\n\n\nSeveral TL models for commercial glasses exist in the literature. Scholze (1990) reports two such\nmodels: \u0160a\u0161ek et al. (1973) and Cuartas (1984). Other models were presented by Babcock (1977) and\nBackman et al. (1997). Component coefficients for these models are summarized in Tables 1 and 2. Table\n1 lists component coefficients for linear models represented by Equation (1). Models by Babcock and\n\u0160a\u0161ek et al. are linear in terms of mass fractions of components (x;). The original model by Cuartas, which\nis linear in x;/xsi02, was converted into the form of Equation (1) by setting Xsio2 = 0.73. Backman et al.'s\n(1997) model is nonlinear. Table 2 shows only coefficients relevant for the SiO2-Na2O-CaO-MgO-Al2O3\nsystem.\n\n\nTable 1. Component Coefficients TL in \u00b0C\nBabcock (a)\n\n\nTridymite Devitrite Wollastonite \u0160a\u0161ek et al. (a) Cuartas (b)\n\n\nBabcock\n\n\nSiO2\n\n\n2324\n\n\n833\n\n\n938\n\n\n1336\n\n\n1384\n\n\n1604\n\n\nAl2O3\n\n\n-7016\n\n\n2093\n\n\n2533\n\n\n1459\n\n\n304\n\n\n3834\n\n\n1229\n\n\nFe2O3\nCaO\n\n\n8936\n\n\n-342\n\n\n3307\n\n\n4349\n\n\n3061\n\n\n2365\n\n\n2060\n\n\n-1225\n\n\nMgO\n\n\n89\n\n\n-765\n\n\nNa\u2082O -4115\n\n\n-2105\n\n\n-2756\n\n\n-1619\n\n\n69\n\n\nK\u2082O\n\n\n3096\n\n\n-2200\n\n\n(a) Babcock's and \u0160a\u0161ek et al.'s models were transformed to the form of Equation (1); the resulting\n\n\ncoefficients are shown.\n\n\n(b)\n\n\nCuartas' coefficients were adjusted to float glass region with 73 mass% SiO2.\n\n\nTable 2. Component Coefficients T in \u00b0C in the Nonlinear Model by Backman et al.(a)\n\n\n(\u00b0C/mass%)\n\n\n8.90E+02\n\n\nIntercept\nCaO\n\n\n3.07E+01\n\n\n-1.83E+00\n\n\nNa\u2082OxNa\u2082O\n\n\nAl2O3 Na\u2082O\n\n\n2.69E+00\n\n\n2.83E-01\n\n\nNa\u2082OxCaOxMgO\n\n\nCaOxCaOxMgO\u00d7MgO\n\n\n-5.30E-02\n\n\nNa\u2082Ox Na\u2082O \u00d7 Na\u2082O \u00d7 Na\u2082O\n\n\n2.10E-03\n\n\n(a)\n\n\nOnly coefficients applicable to the VGS composition region are shown.\n\n\nUnfortunately, the usefulness of T coefficients that cover multiple primary phase fields is rather\ndubious. Because the effect of glass components on T substantially, even dramatically, changes from one\nprimary phase to another, the global models have a significant lack of fit. Global polynomial forms only\npartly remedy the problem. In addition, polynomial models cannot be extrapolated beyond their\nexperimental domain.\n\n\n2\n\n\nGenerally, it is not recommended to apply models beyond the composition region of their validity.\nComposition regions of the literature models are compared with the float-glass composition region in\nTable 3. Babcock used Silverman's (1939) data and a large database available at Owens-Illinois\nlaboratories. Babcock does not provide the composition region of his Owens-Illinois data. Therefore, only\nthe composition region of Silverman's study is shown in Table 3.\n\n\nTable 3. Composition Regions for VGS Glasses and for Published Models (mass%) (a)\n\n\n\u0160a\u0161ek et al.\n\n\nCuartas\n\n\nBackman\n(1997)\nmin max\n\n\nSilverman\n\n\n(1984)\nmin max\n\n\nExperimental\n\n\n(1973)\n\n\n(1939)\n\n\nFloat\n\n\nmin max\n\n\nmin max\n\n\nmin\n\n\nmin max\n\n\nmax\n\n\n67.5 77.2\n\n\n69.4 76\n\n\nSiO2\n\n\n73.1\n\n\n72.7\n\n\n74\n\n\n64.9\n\n\n50.2\n\n\n68.2\n\n\n59 79\n\n\n0.45\n\n\n0.3\n\n\nAl2O3\n\n\n0.4\n\n\n0.5\n\n\n0.1\n\n\n2\n\n\n0.1\n\n\n4\n\n\n0.5\n\n\n3\n\n\n0.2\n\n\n8.2\n\n\nFe2O3\n\n\n1.5\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0.1\n\n\n0.7\n\n\n0.7\n\n\n0\n\n\n0.5\n\n\n0.9\n\n\n12.2\n\n\n9\n\n\n9\n\n\n6.6\n\n\n14.2\n\n\nCaO\n\n\n7\n\n\n8\n\n\n7\n\n\n8.5\n\n\n2.7\n\n\n6\n\n\n18\n\n\nMgO\n\n\n3\n\n\n2.5\n\n\n4.5\n\n\n3\n\n\n0.1 4.9\n\n\n0.5\n\n\n0.2\n\n\n0.2\n\n\n4\n\n\n4\n\n\n5\n\n\n13.1\n\n\n11.8 14.8\n\n\n14.2\n\n\n11\n\n\nNa\u2082O\n\n\n12\n\n\n8.4 22.9\n\n\n15\n\n\n13 16\n\n\n17\n\n\n2 0.03\n\n\nK\u2082O\n\n\n0\n\n\n0 0.5 0.1 2.3 0.5\n\n\n7\n\n\n0.1\n\n\n(a) Values are highlighted yellow/green where maximum/minimum values of\ncomponent ranges were lower/higher than the corresponding maximum/minimum of\nthe experimental range.\n\n\n3\n\n\nExperimental\n\n\nA test matrix of 14 glasses, shown in Table 4, was developed by the VGS project.\n\n\nTable 4. Test Glass Compositions in Mass % of Oxides\n\n\nHigh Al2O3\n\n\nLow CaO High CaO\n\n\n(mass%)\n\n\nStandard\n73.18\n\n\nLow MgO\n\n\nHigh MgO\n\n\nLow Na\u2082O\n\n\nHigh Na\u2082O\n\n\n73.68\n\n\n73.68\n\n\nSiO2\nAl2O3\nTiO2\nFe2O3\n\n\n72.68\n\n\n72.68\n\n\n73.68\n\n\n72.68\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.45\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n8.95\n\n\nCaO\n\n\n8.45\n\n\n7.95\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n2.97\n\n\nMgO\nNa\u2082O\nK\u2082O\n\n\n3.97\n\n\n3.47\n\n\n13.63\n\n\n13.63\n\n\n13.13\n\n\n13.63\n\n\n13.63\n\n\n13.63\n\n\n14.13\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n100.00\n\n\n100.00\n\n\nSum\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\nLow Al2O3\n\n\nLow Na\u2082O\n\n\nHigh CaO\n\n\nHigh Na\u2082O\n\n\n(mass%)\n\n\nLow CaO\n\n\nLow MgO\n\n\nHigh MgO\n\n\nStandard\n\n\n73.53\n\n\n73.03\n\n\n74.03\n\n\n74.03\n\n\nSiO2\n\n\n73.03\n\n\n74.03\n\n\n73.03\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\nAl2O3\nTiO2\nFe2O3\nCaO\n\n\n0.10\n\n\n0.10\n\n\n0.10\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.01\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n0.71\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n8.45\n\n\n7.95\n\n\n8.95\n\n\n3.97\n\n\nMgO\n\n\n3.47\n\n\n3.47\n\n\n3.47\n\n\n2.97\n\n\n3.47\n\n\n3.47\n\n\n13.70\n\n\n13.70\n\n\n13.70\n\n\n13.20\n\n\nNa\u2082O\nK\u2082O\nSum\n\n\n13.70\n\n\n13.70\n\n\n14.20\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n0.03\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\n100.00\n\n\nThe test matrix consists of two standard glasses one for high-alumina (0.45 mass%) glass and the\nother for low-alumina (0.10 mass%) glass. Both standards are equal in Al2O3+SiO2 content and in\nNa2O+K\u2082O content (the K2O content is 0.10 mass% in the high-alumina standard glass and 0.03 mass% in\nthe low-alumina standard glass. The remaining glasses are derived from the standard glasses by one-at-a-\ntime changes in CaO, MgO, and Na2O. These changes are compensated for by equal changes in the SiO2\nfraction. The composition regions covered by the VGS test matrix is shown as the experimental region in\nTable 3 and represented by a parallelogram in Figure 1.\n\n\nAs evident form Table 3, \u0160a\u0161ek et al.'s (1973) composition region is broader than the experimental\ncomposition region, except for Al2O3. The Cuartas (1984) composition region is wider than \u0160a\u0161ek's\ncomposition region, except for the upper limit for the range of SiO2. Backman et al. (1997) studied\nglasses with the SiO2 content <68.2 mass%. Their glasses also had at least 0.5 mass% of each of the\nfollowing oxides: Al2O3, K\u2082O, B2O3, SrO, and BaO. Composition regions in these studies covered several\nprimary phase fields, including tridymite, devitrite, wollastonite, and other phases, such as diopside.\n\n\n5\n\n\nBecause the composition variations are small (see bold numbers in Table 4), careful preparation of\nglasses and precise measurements of the T were required. Batches were prepared from chemicals listed\nin Table 5.\n\n\nTable 5. Source Chemicals\n\n\nChemical\n\n\nManufacturer Lot Number\n\n\nAl2O3\n\n\n006627\n\n\nFisher\n\n\nCaCO3\n\n\n007112\n\n\nFisher\n\n\nFisher\n\n\n005612\n\n\nFe2O3\nK2CO3\n\n\n031384\n\n\nAesar\n\n\nAldrich\n\n\n08629JF\n\n\nMgO\n\n\nNa2CO3\n\n\nFisher\n\n\n006825\n\n\nSiO2\n\n\nFisher\n\n\n010455\n\n\nTiO2\n\n\nJ.T. Baker\n\n\n525355\n\n\nThe chemicals were weighed to obtain batches for 250-g glass. Batches were blended, first by hand in\na plastic bag (roll and shake), and then milled in an agate disc mill for 2 min. The glass was melted twice,\neach time in a Pt-10%Rh crucible for 1 h at 1450\u00b0C. After the first melt, the glass was hand crushed and\nthen milled in a tungsten carbide disc mill for 2 min.\n\n\nThe TL was measured in a gradient temperature furnace. Pt boats filled with crushed glass (between\n20 and 40 mesh size) were heat treated for 24 h in the temperature gradient of 1.0\u00b0C/mm, spanning the\ntemperature interval from 800\u00b0C to 1050\u00b0C. Rods of heat-treated glass were mounted in epoxy, and the\nportion exposed to T >936\u00b0C was thin-sectioned for microscopic evaluation.\n\n\nTo identify phases, approximately 2.5-g samples of each glass were heat-treated in a Pt crucible for\n24 h at 900\u00b0C. X-ray diffraction (XRD) was performed on these samples (using Scintag DMC-008) to\ndetect the crystalline phases present and to measure their concentrations. CaF2 (5 mass%) was added to\nthe samples as the standard to evaluate fractions of crystalline phases. Jade software was used for the\nanalysis.\n\n\nResults\n\n\nXRD analysis detected three phases in the samples: cristobalite (SiO2), devitrite (Na2O.3CaO 6SiO2),\nand wollastonite (CaO-SiO2) (Figure 2). Table 6 shows the fractions of these phases in the glasses at\n900\u00b0C. The visual appearances of crystals of these phases were identified using both XRD data and the\nliterature (Bartu\u0161ka 2001). Using the optical microscope (Olympus PMG-3), the maximum temperatures\n(T;) at which j-th crystals were observed in the thin sections were determined (Table 7); typical examples\nare shown in Figures 3 to 11. The Ts (the highest T;) are shown in bold in Table 7.\n\n\nIVGS-13.rd] VGS-13\n\n\n5000-\n\n\n4500-\n\n\n4000-\n\n\n3500-\n\n\nIntensity (Counts)\n\n\n3000-\n\n\n2500-\n\n\n2000-\n\n\n1500-\n\n\n1000-\n\n\n500-\n\n\n0\n\n\n77-2245> CaF2 - Calcium Fluoride\n\n\n39-1425> Cristobalite, syn Si02\n\n\n23-0671> Na2Ca3 Si6016 - Sodium Calcium Silicate\n\n\n27-0088> Wollastonite-2M- CaSiO3\n\n\n10\n\n\n30\n\n\n40\n2-Theta()\n\n\n70\n\n\n20\n\n\n50\n\n\n60\n\n\nPNNL\n\n\n[JADE|h3085125]<c:\\program files Jade WVGS glasses> Thursday, Jul 25, 2002 02:39p (MDI/JADE6)\n\n\nFigure 2. XRD Pattern of Low-Al2O3-Low-Na\u2082O Glass Heat-Treated at 900\u00b0 C for 24 h\n\n\n7\n\n\nTable 6. Mass Fractions of Crystalline Phases in Glasses Heat-Treated at 900\u00b0 C for 24 h\n\n\nCristobalite Devitrite Wollastonite\n\n\nHigh Al2O3\n\n\nStandard\n\n\n0.033\n\n\n0.060\n\n\nLow CaO\nHigh CaO\nLow MgO\n\n\n0.041\n\n\n0.061\n\n\n0.043\n\n\n0.045\n\n\n0.047\n\n\n0.061\n\n\n0.061\n\n\nHigh MgO\n\n\nLow Na\u2082O\n\n\n0.072\n\n\n0.076\n\n\n0.076\n\n\nHigh Na\u2082O\n\n\n0.020\n\n\n0.079\n\n\nLow Al2O3\n\n\nStandard\n\n\n0.052\n\n\n0.066\n\n\nLow CaO\nHigh CaO\nLow MgO\n\n\n0.071\n\n\n0.046\n\n\n0.129\n\n\n0.061\n\n\n0.067\n\n\n0.055\n\n\n0.031\n\n\nHigh MgO\n\n\n0.057\n\n\nLow Na\u2082O\n\n\n0.079\n\n\n0.072\n\n\n0.092\n\n\nHigh Na\u2082O 0.027\n\n\n0.055\n\n\nTable 7. Maximum Temperatures (\u00b0C) at Which Individual Crystalline\nPhases Occurred in VGS Glasses (a)(b)\n\n\nTridymite\n\n\nDevitrite Wollastonite\n\n\nTL\n\n\nHigh Al2O3\n\n\nStandard\n\n\n967\n\n\n992\n\n\n992\n\n\nLow CaO\nHigh CaO\nLow MgO\nHigh MgO\n\n\n952\n\n\n973\n\n\n990\n\n\n990\n\n\n973\n\n\n1011 1011\n\n\n962\n\n\n997\n\n\n980\n\n\n997\n\n\n975\n\n\n1001 1001\n\n\nLow Na\u2082O\n\n\n1011\n\n\n971\n\n\n998 1011\n\n\nHigh Na\u2082O\n\n\n989 989\n\n\n965\n\n\nLow Al2O3\n\n\n995\n\n\nStandard\n\n\n965\n\n\n977 995\n\n\n968\n\n\nLow CaO\nHigh CaO\nLow MgO\nHigh MgO\n\n\n1024\n\n\n968\n\n\n1024\n\n\n1000\n\n\n1000\n\n\n965\n\n\n980 1026\n\n\n970\n\n\n1026\n\n\n965\n\n\n994 994\n\n\n1044(c)\n\n\nLow Na\u2082O\n\n\n962\n\n\n990 1044\n\n\nHigh Na\u2082O\n\n\n979\n\n\n970\n\n\n979\n\n\n(a)936\u00b0C was the minimum temperature at which glasses were evaluated.\n(b)Bold numbers indicate T (see also the last column)\n(1044\u00b0C was the maximum temperature at which glasses were evaluated; repeated tests showed that this was also\nthe TL value for the low-Al2O3-low-Na2O glass.\n\n\n8\n\n\nVGS-9, Cristobalite, 1020\u00b0C 0.010 mm\n\n\n0.010 mm\n\n\nVGS-11, Cristobalite, 1000\u00b0C\n\n\nFigure 4. Crystals of Cristobalite in Low-Al2O3-\nLow-MgO Glass, 1000\u00b0C\n\n\nFigure 3. Crystals of Cristobalite in Low-Al2O3-\nLow-CaO Glass, 1020\u00b0C\n\n\nVGS-10, Devitrite, 970\u00b0C\n\n\nVGS-12, Devitrite, 965\u00b0C\n\n\n0.1 mm\n\n\n0.1 mm\n\n\nFigure 5. Crystals of Devitrite in Low-Al2O3-\nHigh-CaO Glass, 970\u00b0C\n\n\nFigure 6. Crystals of Devitrite in Low-Al2O3-\nHigh-MgO Glass, 965\u00b0C\n\n\n9\n\n\nVGS-7, Wollastonite, 965\u00b0C\n\n\nVGS-10, Wollastonite, 995\u00b0C 0.010 mm\n\n\n0.1 mm\n\n\nFigure 7. Crystals of Wollastonite in Low-Al2O3-\nHigh-CaO Glass, 995\u00b0C\n\n\nFigure 8. Crystals of Wollastonite in High-Al2O3-\nHigh-Na\u2082O Glass, 965\u00b0C\n\n\nVGS-1, Devitrite, 975\u00b0C\n\n\n0.1 mm\n\n\nVGS-12, Wollastonite, 995\u00b0C\n\n\n0.010 mm\n\n\nFigure 10. Crystals of Wollastonite and Probably\nDevitrite, Nucleated on a Bubble, in High-Al2O3-\nBaseline Glass, 975\u00b0C\n\n\nFigure 9. Crystals of Wollastonite, Nucleated on a\nBubble, in Low-Al2O3-High-MgO Glass, 995\u00b0C\n\n\n10\n\n\nVGS-5, Devitrite, 950\u00b0C\n\n\n0.1 mm\n\n\nFigure 11. Crystals of Devitrite and Wollastonite\nin High-Al2O3-High-MgO Glass, 950\u00b0C\n\n\n11\n\n\nDiscussion\n\n\nFigures 12 to 14 show that T; is a nearly linear function of composition. To fit Equation (1) to data,\nmass fractions were normalized to the following five components: SiO2, Na2O, CaO, MgO, and Al2O3.\nThe resulting coefficient values are listed in Table 8. The T; values calculated with these coefficients,\nusing Equation (1), are compared with measured T;s in Figure 15.\n\n\nTable 8. Component Coefficients (T;;) for the Maximum Temperatures at which Tridymite,\nDevitrite, Wollastonite Occur in Float Glasses at Equilibrium (in \u00b0C)(a)\n\n\nDevitrite\n863\n\n\nWollastonite\n\n\nTL\n1669\n\n\nTridymite\n\n\nCCCC\n\n\nSiO2\nAl2O3\n\n\n3033\n\n\n763\n\n\n2787\n\n\n862\n\n\n-1936\n\n\n-7560\n\n\n4235\n2499\n\n\n-2525\n\n\nCaO\n\n\n1756\n\n\n1521\n\n\n-3418\n\n\nMgO\nNa\u2082O\n\n\n1260\n\n\n281\n\n\n-2646\n\n\n-230\n\n\n-6593\n\n\n962\n\n\nR\u00b2\n\n\n0.233\n\n\n0.608\n\n\n0.997\n\n\n0.950\n\n\nadj R\u00b2\n\n\n-0.108\n\n\n0.991\n\n\n0.434\n\n\n0.927\n\n\n3.3\n\n\n1.9\n7\n\n\n6.0\n14\n\n\n13.2\n\n\nS\n\n\n14\n\n\n14\n\n\nn\n\n\n(a) The coefficients are valid for the experimental composition region shown in Table 3.\nNote that the effect of Al2O3 is compounded with the effect of K\u2082O replacing Na\u2082O. The\nsymbols on the two bottom lines: s is the standard error, and n is the number of\nobservations.\n\n\n\u25b2 wollastoniteHiAl\n\n\n\u25b2 wollastonite HiAl\n\n\ndevitrite HiAI\ndevitriteLoAl\n\n\ndevitrite HiA\ndevitrite LoAl\n\n\ntridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\n\u25c6tridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\nMgO Effect\n\n\nCaO Effect\n\n\nA wollastoniteLoA\n\n\nA wollastoniteLoAl\n\n\n1030\n\n\n1030\n\n\n1020\n\n\n1020\n\n\n1010\n\n\n1010\n\n\nTemperature \u00b0C\n\n\nTemperature \u00b0C\n\n\n1000\n\n\n1000\n\n\n990\n\n\n990\n\n\n980\n\n\n980\n\n\n\u0394\n\n\n970\n\n\n970\n\n\n960\n\n\n960\n\n\n950\n\n\n940\n\n\n950\n\n\n7.8\n\n\n8\n\n\n8.2\n\n\n8.4\n\n\n8.6\n\n\n8.8\n\n\n9\n\n\n3.2\n\n\n3.6\n\n\n2.8\n\n\nMgO mass%\n\n\nCaO mass%\n\n\nFigure 13. The Effect of MgO on the Maximum\nTemperature at which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium\n\n\nFigure 12. The Effect of CaO on the Maximum\nTemperature at which Tridymite, Devitrite, and\nWollastonite occur in Float Glass at Equilibrium\n\n\n13\n\n\n\u25c6 tridymiteHiAl\n\u25c7 tridymiteLoAl\n\n\ndevitrite HiAI\ndevitriteLoAI\n\n\n\u25b2 wollastoniteHiAl\n\n\n1060\n\n\nNa\u2082O Effect\n\n\nA wollastoniteLoAl\n\n\n1060\n\n\n1040\n\n\n1040\n\n\nTemperature \u00b0C\n\n\nCalculated T; (\u00b0C)\n\n\n1020\n\n\n1020\n\n\n1000\n\n\nA\n\n\n1000\n\n\n980\n\n\n\u25c6 Tridymite\n\n\n980\n\n\n\u25c9\n\n\n960\n\n\nDevitrite\n\n\n940\n\n\n\u25b2 Wollastonite\n\n\n960\n\n\n13.9\n\n\n14.3\n\n\n13.1\n\n\n13.5\n\n\nNa\u2082O mass%\n\n\n940\n\n\nFigure 14. The Effect of Na\u2082O on the Maximum\nTemperature at Which Tridymite, Devitrite, and\nWollastonite Occur in Float Glass at Equilibrium\n\n\n940\n\n\n990\n\n\n1040\n\n\nMeasured T; (\u00b0C)\n\n\nFigure 15. T;, Calculated vs. Measured\n\n\nAt 900\u00b0C, the crystalline form of silica as detected by XRD was cristobalite (Figure 2). At\ntemperatures below 1470\u00b0C, the stable form of silica in soda-lime glasses is tridymite; however, as\nreported by Bartuska (2001), cristobalite occurs in industrial glasses at lower temperatures, alone or\ntogether with tridymite. No attempt was made to distinguish between these two forms in this study.\nHowever, we assume that tridymite was the primary phase.\n\n\nAs Table 7 shows, tridymite was the primary phase in all glasses in which the SiO2 content was >75.3\nmass%. In all other glasses, the primary phase was wollastonite. In these glasses cristobalite did not form\nat T >936\u00b0C, but was detected in some glasses that were heat-treated at 900\u00b0C (Table 6). Devitrite and\nwollastonite were found in all glasses, though wollastonite disappeared at lower temperatures and was not\ndetected at 900\u00b0C in most glasses (Table 6). The highest temperature at which devitrite was observed was\nvirtually independent of composition. It varied within a narrow temperature interval of 950\u00b0C to 975\u00b0C,\nbut these variations in T; cannot be explained in terms of composition variation: note the low R\u00b2 value for\nTdevitrite in Table 8.\n\n\nTable 8 shows that the Tcristobalite increases with SiO2 fraction and decreases with the addition of any\nother component, most notably Al2O3 and Na\u2082O. The formation of wollastonite is promoted mainly by\nCaO, but also by MgO and Al2O3, whereas NaO2 suppresses its formation. The component coefficients\nfor tridymite and wollastonite are much different for each component, showing that the slopes of the\ncorresponding liquidus surfaces are substantially different and even opposite in Al2O3, CaO, and MgO\ncomponents. Therefore, the linear model for T (the last column in Table 8) is not a good representation\nof the real behavior. This is well illustrated in Figures 12 to 14.\n\n\nThe low-Al2O3 standard glass is close to the boundary between the tridymite and wollastonite primary\nphases. Its T is only 3\u00b0C higher than that of the high-Al2O3 standard; this difference is within the\nstandard deviation for Tj; of wollastonite (Table 8). It is, however, possible to decrease low Al2O3 glass TL\nby changing the MgO-to-CaO ratio.\n\n\nFigures 16 and 17 compare T; and T\u00cb measured values with those calculated using different models. It\nis not surprising, as seen in Figure 17, that linear or polynomial models are subjected to a large error.\n\n\n14\n\n\n1200\n\n\n1150\n\n\n1150\n\n\n1100\n\n\n1050\n1100\n650\n\n\nCalculated T; (\u00b0C)\n\n\nCalculated T\u2713 (\u00b0C)\n\n\n1050\n\n\n1000\n\n\n\u25c6 Tridymite\nDevitrite\n\n\nSasek\nCuartas\n\n\n1000\n\n\n950\n\n\n\u25b2 Babcock +\n\u25c6 Backman\n\n\n\u25b2 Wollastonite\n\n\n\u25c7 Babcock Cristobalite\n\n\n950\n\n\nTL\n\n\nBabcock Devitrite\n\n\n900\n\n\nA Babcock Wollastonite\n\n\n850\n\n\n900\n\n\n950 1000 1050 1100 1150 1200\n\n\n850\n\n\n900\n\n\n1050\nMeasured T (\u00b0C)\n\n\n900\n\n\n950\n\n\n1000\n\n\n1100\n\n\n1150\n\n\nMeasured T; (\u00b0C)\n\n\nFigure 17. Comparison of Experimental TL\nValues for VGS Glass with T Values Predicted\nby Various Models\n\n\nFigure 16. Comparison of Experimental T;\nValues for VGS Glass with Babcock's T\u2081 Values\nfor Soda-Lime Glass\n\n\nUsing Babcock's coefficients, T; values were predicted for experimental compositions normalized to\nthe Na\u2082O-CaO-SiO2 system. As seen in Figure 16, Babcock's models overestimate T; values for tridymite\nand underestimate T; values for devitrite and wollastonite. These differences can be attributed to the effect\nof MgO in the experimental glasses.\n\n\nAs a comparison of the coefficients in Table 8 with those in Table 1 shows, Cuartas' coefficients are\ncloser to our empirical coefficients than those by \u0160a\u0161ek et al. (1973). Hence, even though \u0160a\u0161ek et al.'s\n(1973) composition region is closer to our experimental composition region than that of Cuartas (1984),\nT\u2081 values from \u0160a\u0161ek et al.'s model less closely match the experimental Ts (Figure 17).\n\n\nThe main cause of the difference between our Al2O3 coefficients for T and those by other authors is\nprobably the confounding between substituting SiO2 for Al2O3 together with substituting K\u2082O for Na2O.\nOur Al2O3 coefficient accounts for this four-component change and thus cannot be compared with studies\nin which the Al2O3 fraction was an independent variable. Another factor is the extremely narrow range of\nAl2O3 content in our study (0.10 to 0.45 mass% as compared to 0.2 to 8.2 mass% of Silverman [1939]).\n\n\nThe experimental composition region of this study is shown in Figure 1 as a parallelogram. Figure 1\ndisplays isotherms and SiO2 fractions on the SiO2 saturation surface in the Na2O-CaO-MgO-SiO2 system.\nWhen the experimental glasses are normalized to this four-component system, the SiO2 mass fraction\nvaries from 0.737 to 0.747. Glasses with tridymite primary phase are those with a SiO2 mass fraction\nhigher than 0.741 (within the Na\u2082O-CaO-MgO-SiO2 system). This is in excellent agreement with SiO2\ncontours in Figure 1.\n\n\nFigure 1 suggests that glasses outside of the tridymite primary phase field are located in the devitrite\nprimary phase field, whereas the experimental glasses had wollastonite as their primary phase. As\nmentioned earlier, this shift was probably caused by Fe2O3 and Al2O3 in the experimental glass. These\ntwo oxides become network formers by immobilizing Na+ ions. Therefore, the effective Na\u2082O content is\n\n\n15\n\n\nlower in the experimental glasses, which would shift the experimental composition region into the\nwollastonite fields.\n\n\n16\n\n\nConclusions\n\n\nLow-alumina the standard glass has nearly identical TL (995\u00b0C) as the high-alumina the standard\nglass (992\u00b0C). Its primary phase is tridymite, while wollastonite is the primary phase of the latter glass.\nDecreasing SiO2 fraction by as little as 0.3 mass% (by replacing it with CaO, MgO, Na2O, or their\ncombination) shifts the glass into the wollastonite primary phase field. While in this field, the secondary\nphase is devitrite.\n\n\nNeither of the two standard glasses contains wollastonite when equilibrated at 900\u00b0C; both contain\ncristobalite and devitrite, the low-alumina glass at somewhat higher fraction. This, however, should not be\na problem in float glass manufacturing.\n\n\nThe effects of SiO2, CaO, MgO, and Na\u2082O on the T within each primary phase field were expressed\nas partial-specific Ts. These coefficients allow a further optimization on the low-alumina glass.\n\n\nThe partial-specific T coefficients published in the literature do not predict the T values\nsatisfactorily. It appears that two conditions are necessary for a correct prediction of T: 1) separate\nmodels are developed for each primary phase field and 2) the major components of the model are the\nsame as the major components of the composition region of interest.\n\n\n117\n\n\nReferences\n\n\nBabcock CL. 1977. Silicate Glass Technology Methods. John Wiley, New York.\n\n\nBackman R, KH Karlsson, M Cable, and NP Pennington. 1997. \u201cModel for Liquidus Temperature of\nMulti-Component Silicate Glasses.\" Physics and Chemistry of Glasses 38, 103-109.\n\n\nBartu\u0161ka M. 2001. \u201cCrystalline Inclusions.\u201d In Glass Flaws (in Czech), Bartu\u0161ka M, Editor, Pr\u00e1h,\nPrague, Czech Republic.\n\n\nCuartas R. 1984. \"Calculo teorico de propiedades del vidrio: viscosidad, parametros termicos y\nparametros de desvitrificacion.\" Ceram. Vidrio 23, 105-111.\n\n\nRoth RS, T Negas, and LP Cook. 1981. Phase Diagrams for Ceramists. Volume IV. American\nCeramic Society, Columbus, Ohio.\n\n\n\u0160a\u0161ek L, M Bartu\u0161ka, and V Van Thong. 1973. \u201cUtilization of Mathematico-Statistical Methods in\nSilicate Research. 2. Determination of Mathematical Relations for the Calculation of Crystallization\nProperties from Chemical Composition of Sheet and Container Glass\u201d (in Czech). Silikaty 17, 207-217.\nScholze H. 1990. Glass Nature, Structure, and Properties. Springer, New York.\n\n\nShahid KA and FP Glasser. 1972. Physics and Chemistry of Glasses 13, 27 (referenced in Roth et al.\n1981).\n\n\nSilverman WB. 1939. \u201cEffect of Alumina on Devitrification of Soda-Lime-Silica Glasses.\" Journal of\nthe American Ceramics Society 22, 378-384.\n\n\n19\n\n\nPNNL-13958\n\n\nDistribution\n\n\nNo. of\n\n\nNo. of\nCopies\n\n\nCopies\n\n\nOFFSITE\n\n\nONSITE\n\n\n10 Pacific Northwest National Laboratory\n\n\n1 James V. Jones\n\n\nD-S Kim\n\n\nTechnical Fellow\n\n\nK6-24\n\n\nAdvanced Technology and Engineering\n\n\nJ. D. Vienna\n\n\nK6-24\n\n\nVisteon Glass System\n\n\nP. R. Hrma (7)\n\n\nK6-24\n\n\n15000 Commerce Drive\n\n\nM. A. Khaleel\n\n\nK2-18\n\n\nN. Dearborn, MI 48120\n\n\nEdward N. Boulos, Ph.D. (3)\n\n\n3\n\n\nSenior Technical Fellow\n\n\nAdvanced Technology and Engineering\n\n\nVisteon Corporation\nVisteon Glass System\n\n\n15000 Commerce Drive\n\n\nN. Dearborn, MI 48120\n\n\n1\n\n\nDr. Theodore M. Besmann, Head\n\n\nSurface Processing and Mechanics Group\n\n\nOak Ridge National Laboratory\n\n\n1 Bethel Valley Road\n\n\nP.O. Box 2008\n\n\nOak Ridge, TN 37831-6063\n\n\nDistr. 1\n"}, "expected_output": {"claims": [{"unit": "%", "value": 0.1, "evidence": ["Al2O3", "0.45", "0.10", "The test matrix consists of two standard glasses one for high-alumina (0.45 mass%) glass and the\nother for low-alumina (0.10 mass%) glass.", "The liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to\n74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to\n14.2 mass% Na2O. Crystalline phases were identified at 900\u00b0C as cristobalite, wollastonite, and devitrite.\nThe primary phases were tridymite and wollastonite. Partial specific Ts were obtained from the data and\ncompared with the literature.\n", "Substantial savings can be achieved if low-alumina float glass is produced. It has been proposed to decrease the Al2O3 content in the float glass produced by the VGS Glass Systems (VGS) from the original 0.45 mass% to 0.10 mass%.", "The liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to 74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to 14.2 mass% Na2O."]}, {"unit": "%", "value": 0.45, "evidence": ["Al2O3", "0.45", "0.10", "The test matrix consists of two standard glasses one for high-alumina (0.45 mass%) glass and the\nother for low-alumina (0.10 mass%) glass.", "The liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to\n74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to\n14.2 mass% Na2O. Crystalline phases were identified at 900\u00b0C as cristobalite, wollastonite, and devitrite.\nThe primary phases were tridymite and wollastonite. Partial specific Ts were obtained from the data and\ncompared with the literature.\n", "Substantial savings can be achieved if low-alumina float glass is produced. It has been proposed to decrease the Al2O3 content in the float glass produced by the VGS Glass Systems (VGS) from the original 0.45 mass% to 0.10 mass%.", "The liquidus temperature (TL) was measured for a float-glass-type composition region with 72.7 to 74.0 mass% SiO2, 0.1 to 0.45 mass% Al2O3, 8.0 to 9.0 mass% CaO, 3.0 to 4.0 mass% MgO, and 13.1 to 14.2 mass% Na2O."]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_7b57f5a0d8068ba3"}}
{"id": "76e78c2e94ddea3fbf028e90", "input": {"query": "What are the typical scrap rates for injection molding operations? Please provide specific percentages for different materials, machine types, or process conditions.", "source_url": "https://minds.wisconsin.edu/bitstream/handle/1793/41389/2004nguyend.pdf?sequence=1&isAllowed=y", "document_text": "INJECTION MOLDING SCRAP REDUCTION: A STUDY IN\nTHE RELATIONSHIPS OF PLASTICS PROCESSING METHODS\nBy\nDuoc T. Nguyen\nA Research Paper\nSubmitted in Partial Fulfillment of the\nRequirements for the\nMaster of Science Degree in\nManagement Technology\nApproved for Completion of 3 Semester Credits\nINMGT 735\nResearch Advisors\nThe Graduate College\nUniversity of Wisconsin\nMay 2004\nThe Graduate School\nUniversity of Wisconsin - Stout\nMenomonie, WI 54751\nABSTRACT\nNguyen\n(Writer)\n(Title)\n(Last Name)\nInjection Molding Scrap Reduction: a Study in the\nRelationships of Plastics Processing Methods\nDuoc\n(First)\nT.\n(Initial)\nManagement Technology Robert H. Feirn & Linards Stradins\n57 Pages\n(Graduate Major)\n(Research Advisor) (Month/Year) (No. of Pages)\n5/2004\nPublication Manual of the American Psychological Association\n(Name of Style Manual Used in this Study)\nTo maintain competitiveness and maximize profits in today's marketplace, one of\nthe most important aspects that organizations must focus on is scrap reduction. In the\nplastics industry Phillips Plastics-Short Run, an injection molding facility, methodically\ncollects data in an effort to better understand and control scrap generation.\nThe purpose of this research is to analyze the scrap levels associated with\nvariances occurring in the injection molding machines, processes, materials, and\noperators at the Short Run facility. The results help to identify the possible causes of\nscrap and will lead to an appropriate solution to support Short Run in reducing scrap.\nAcknowledgements\nThanks to John Ahlbrecht, Pete Posch, and Matt Rominski, the Plant Manager,\nProduction Manager, and Maintenance Manager, respectively, at Phillips Plastics-Short\nRun in directing and supporting the data collection for this research.\nThanks to Mike Cran, Linda Whitcome, and Ron Watrud, 1st, 2nd, and 3rd shift\nSupervisors at Short Run for supporting the collection of the necessary data related with\noperators and their years of experience.\nThanks to Mr. Robert H. Feirn and Mr. Linards Stradins, my dedicated advisors,\nfor all their patience and knowledge to assist me through careful reading and support\nthrough valuable feedback to edit this paper.\nParticular thanks go to Mr. James Keyes, my wonderful teacher at UW-Stout in\nManagement Technology in putting his time in reviewing the research and giving me the\ngreat comments and feedback to improve the text of this research.\nFinally, special thanks to my wife, my daughter, and my son whose care and\nencouragement lifted me over a tough spot.\nTable of Contents\nAbstract..\nChapter 1.\nIntroduction..\nStatement of Problems\nObjectives of the Study.\nPurpose.........\nLimitations.\nDefinitions...\nChapter 2..........\nLiterature Review.\nIntroduction.....\nMachine......\nProcess\nMaterial.\nOperator\nChapter 3..\nMethodology.\nResearch Design\u2026........\nData Collection\nAnalysis............\nMachine...\nProcess\nMaterials\nOperator\niv\n.ii\n6\n6\n6\n6\n6\n7\n7\n10\n10\n10\n11\n12\n14\n17\n19\n19\n20\n20\n21\n21\n22\n222\n22\n23\nV\nChapter 4\nResults..\nFigure 1 Scrap Percentage Related to Machine Age.\nFigure 2 Scrap Percentage Related to Machine size..\nFigure 3 Scrap Levels Associated to Melt Temperature\nFigure 4 Scrap Levels Related to Mold Temperature\u2026..\nFigure 5 Scrap Level Related to Pack Pressure...\nTable 1 Scrap Level Related to Calibre..\nTable 2 Scrap Level Related to Radel.\nTable 3 Scrap Level Related to Ultem\nTable 4 Scrap Level Related to Cycolac.\nTable 5 Scrap Level Related to RTP .\nFigure 6 Scrap Levels Related to 1st Shift Operators..\nFigure 7 Scrap Levels Related to 2nd Shift Operators.\nFigure 8 Scrap Levels Related to 3rd Shift Operators........\nFigure 9 Scrap Levels Related with Each Production Line.\nFigure 10 Scrap Levels Related to 1st Shift's A-Line Based on Experience.\nFigure 11 Scrap Levels of 1st Shift's B-Line Operators Based on Experience..\nFigure 12 Scrap Levels of 1st Shift's C-Line Operators Based on Experience..\nFigure 13 Scrap Levels of 2nd Shift's A-Line Operators Based on Experience..\nFigure 14 Scrap Levels of 2nd Shift's B-Line Operators Based on Experience..\nFigure 15 Scrap Levels of 2nd Shift's C-Line Operators Based on Experience ...\nFigure 16 Scrap Levels of 3rd Shift's A-Line Operators Based on Experience.\nFigure 17 Scrap Levels of 3rd Shift's B-Line Operators Based on Experience.\nFigure 18 Scrap Levels of 3rd Shift's C-Line Operators Based on Experience ..\nValidity\nDiscussion.\nChapter 5 .\u2026\u2026\u2026\u2026\u2026\u2026\u2026\u2026\u2026\u2026\nConclusion\nRecommendations....\nReferences...\nAppendix A.\nAppendix B\nAppendix C..\n23\n22\n23\n24\n25\n.25\n.26\n.27\n28\n28\n29\n29\n30\n30\n31\n32\n.32\n33\n34\n.35\n.36\n.37\n.38\n38\n39\n.40\n.41\n41\n42\n52\n52\n53\n.54\n55\n56\n57\nChapter 1\n6\nIntroduction\nIn today's business, controlling scrap is one factor that companies focus on in\norder to remain competitive and maximize benefit. More specifically, in the injection\nmolding area, Phillips Plastics-Short Run is trying to improve scrap control. This can be\ndone through identifying the scrap associated with injection molding machines, plastic\nprocesses, materials, and operators. The purpose of this research is to analyze and\nidentify the scrap levels related with those factors to help Short Run in better controlling\ntheir scrap rate.\nStatement of Problems\nThis study analyzes scrap produced at Philips Plastics-Short Run, a plastic\ninjection molding facility, to discover the scrap levels associated with machines,\nprocesses, materials, and operators to help improve Short Run's scrap control.\nObjectives of the Study\nThe objective of this research was to collect data from Phillips Plastics-Short\nRun's scrap history log, analyze the scrap data, and identify the scrap levels to help\ncontrol scrap rates at Short Run.\nPurpose\nThe purpose of this study is to help Short Run, an injection molding facility, to\nidentify a better solution toward the reduction of scrap. This study focuses on analyzing\nthe scrap related with injection molding machines, plastic processes, materials, and\noperators. The results will help to identify the causes of scrap and will lead to an\nappropriate solution to support Short Run in reducing scrap.\n7\nLimitations\nThis study analyzes the scrap data from Short Run's scrap history log. The results\nare based on 100 production samples taken at Short Run. Part mold design is not\ndiscussed in this study. The time factor in each process will not be discussed in this study.\nThe post-mold operation scrap and start-up scrap are also not included in this research.\nThis research focuses only on the key parameters related with plastics processing.\nDefinitions\nAmorphous Polymers A family of polymers characterized by the randomness of\nentangled polymer chains.\nAnisotropic Shrinkage A shrinkage that is not the same in all directions. It occurs in\nfiber filled materials due to the restriction of shrinkage along the fiber length, which\ntends to be in the flow direction.\nBranched Polymer A polymer that has additional monomer chains protruding from its\nprimary chain.\nBubble Air or other gas trapped within the plastic leading to a void in the part.\nBurning Showing evidence of thermal decomposition through some discoloration,\ndistortion, or localized destruction of the surface of the plastic.\nCalibre The trade name of a polycarbonate material.\n8\nClamping Force The maximum holding force, expressed in tons, that a machine is\ncapable of maintaining.\nContamination Imperfections caused by foreign material molded into the part.\nCrystalline Polymers A family of polymers characterized by areas of order in which the\nmolecular chains line up and lay tightly together in an otherwise amorphous mass.\nCrystallization Temperature The temperature at which a crystalline resin begins to\ncrystallize upon cooling.\nCycle Time The time that has elapsed between the starting point in one cycle of\nproduction and the starting point in the next cycle.\nCycolac The trade name of acrylonitrile-butadience-styrene (ABS) material.\nDegradation A reduction in the physical properties of polymers due primarily to\nbreaking of the long chained molecules. It occurs when the resin is heated at too high a\ntemperature, or for too long, and can result in substandard parts.\nDelamination The splitting of a plastic material along the plane of its layers. It is a\nphysical separation or loss of bond between laminate plies.\nEngineering Thermoplastics A group of thermoplastics generally considered as high\nperformance materials.\nFlash Extra plastic attached to a molding along the parting line.\nFlow Marks A mark on a molded piece made by the meeting of two flow fronts during\nmolding.\nGlass Transition Temperature (Tg) The temperature at which a material turns rubbery\nupon heating and glassy upon cooling.\n9\nInjection Molding A polymer processing method that produces intricate, high\nperformance, precise parts with very little secondary labor operations and with minimal\nwaste.\nMelting Point (Tm) The temperature at which the crystalline regions soften and begin to\nflow.\nMolded-In Stress The stress that has been \u201cbuilt into\u201d the part during processing.\nMold Shrinkage The amount of material shrinkage in the mold that must be\naccommodated for in the tooling design and is reported in inches/inch, mils/inch, or as a\npercentage.\nMolecular Weight The weight of a molecule, calculated as the sum of the atomic\nweights of its atoms.\nMolecule The smallest unit of a substance, able to exist by itself and retain all of the\nproperties of the original substance. Molecules are composed of one or more atoms.\nPlasticize A combination of electrical and mechanical heat energy that is used to soften\npolymer pellets until they flow.\nPlastics A material containing one or more organic polymeric substances of large\nmolecular weight.\nPolymer A chemical compound formed by many small molecular units joined together to\nform a large, chain-like molecule.\nPolymerization A chemical reaction in which two or more small molecules combine to\nform large molecules that contain repeating structural units of the original molecules.\nRadel The trade name of polyerylsulfone (PES) material.\nResins Any of various materials made from polymers or plastics.\n10\n10\nRheology The study of material flow.\nRTP the trade name of a polycarbonate material manufactured by RTP.\nShort Shot A part that is not completely filled with plastic.\nShrinkage A volume reduction of polymers that occurs during cooling due to a reduction\nin space between the molecules.\nSink Marks A depression, or \u201cdimple\u201d due to shrinkage on a part, usually found on the\nopposite side of where a thick wall section exists, caused by internal stress.\nSplay Silver-streaked appearance caused by gasses in the plastic when parts are filled,\nusually caused by moisture or material degradation.\nThermoplastics A family of polymers characterized by its ability to be reprocessed.\nUltem The trade name of a polyetherimide (PEI) material.\nViscosity The resistance to flow of a plastics material. It is the ratio of shearing stress to\nrate of a shear.\nVoid Caused by internal stresses pulling plastic molecules apart.\nWarpage Dimensional distortion in a plastic object after molding.\nWeld Line The line in a part which results when two flow fronts meet and \u201cknit.\u201d\nChapter 2\nLiterature Review\nIntroduction\nThe analysis of most injection molding problems usually focuses on the molding\ncycle (Reinhold, 1986). Typically these types of analyses are concerned with three major\n\n\n11\nelements of the molding operation: injection molding machine, mold, and material. The\nperformance of these operating elements is influenced by the three variables controlling\nthe injection molding process: time, pressure, and temperature. They are all interrelated\nvariables. A change in one or more variables in the molding operation can affect the\nwhole process. Molding problems may also occur due to the action of operators\n(Reinhold, 1986). For example, jobs that require loading inserts, pick-outs, or applying\nmold release before starting the cycle, are hard to keep on a consistent cycle time.\nHowever, through using CAD/CAM, injection molding simulation and prototyping, the\nproblems due to mold design are minimized (Beaumont, Nagel, Sherman, 2002). Cycle\ntime also depends on the response of hydraulic valves and the mechanical system of the\ninjection molding machine. For this reason, this research does not include process time in\nthe analysis. The major operating elements, machines, processes, materials, and operators\nand their associated problems that cause scrap are presented in this chapter.\nMachine\nThe basic process requirements that each injection molding machine must meet\nare based on process time, temperature, and pressure (Reinhold, 1987). However,\nmachine characteristics may change with age (Beaumont, Nagel, & Sherman, 2002).\nScrap occurs due to inappropriate machine conditions such as a malfunctioning feed\nsystem, inconsistent screw stop action, inconsistent screw speed, uneven back\nadjustment, malfunctioning temperature control system, insufficient plasticizing capacity,\ninconsistent cycle, clamp pressure not maintained, and so forth (Reinhold, 1986).\nInconsistent machine control can cause material degradation, part delamination, burning,\nflash, short shot, sink marks, flow marks, etc (Reinhold, 1986).\npressure\nProcess\n12\nInjection molding is a process by which plastic raw materials are converted into\nuseful products meeting acceptable standards (Beaumont, Nagel, & Sherman, 2002).\nEach plastic material, based on type and grade, can be correctly processed within a\ncertain range of temperatures and pressures (Whelan, Anthon, Craft, 1978). These are the\nkey parameters of plastic processing. The purpose of this section is to review these key\nparameters and their associated problems that can lead to scrap.\nMelt temperature is the temperature at which a resin changes from a solid to a\nliquid. At this point, the crystalline regions of a plastic material soften and begin to flow\n(Reinhold, 1991). Melt temperatures range from 248\u00b0F to 662\u00b0F (Beaumont, Nagel, &\nSherman, 2002) and are controlled by barrel temperature, nozzle temperature, screw\nspeed, back pressure, and residence time.\nAbout 70% of the heat needed to melt the plastic is generated from shear heating\nthat occurs within the plastic itself (Beaumont, Nagel, & Sherman, 2002). Therefore, the\nmelt temperature is difficult to measure and cannot be directly controlled by the\nthermostats on the control panel. The injection molding machine barrel and nozzle\ntemperatures are maintained by electrical heating elements. These elements provide the\nremaining 30% of the heat required to maintain a plastic temperature high enough to\nguarantee plastic flow.\nThe thermodynamic properties of the molten plastic, such as viscosity, enthalpy,\nand specific volume, change simultaneously with melt temperature (Johannaber, 1994).\nTherefore, it is important to try and maintain a constant melt temperature in the molding\nprocess. Any change in melt temperature leads to a change in cavity pressure, cycle time,\n13\nor injection time. The problems associated with inconsistent melt temperatures may cause\nsome common part defects such as splay, burning, flash, flow line, short shot, sink marks,\nbubbles, etc (Reinhold, 1996).\nMold temperature is the temperature that maintains the cavity surface at a\nspecified temperature to cool a hot plastic to a solid state. Mold temperatures range from\n32\u00b0F to 302\u00b0F, and are dependent on the coolant temperature, flow rate, proximity of the\ncoolant channels to the plastic, and the rate of the heat input from the plastic. The coolant\nflow rate reaches its maximum if the fluid is in the state of turbulent flow. The flow can\nbe characterized by calculating the Reynold's number for the stream. It is the ratio of\ninertia forces to viscous forces within the flow field. The Reynold's number is a\ndimensionless number based on the channel diameter, flow rate, density, and viscosity of\nthe fluid. If this number is greater than 10,000 (Reinhold, 1996), turbulent flow is fully\ndeveloped and the co-efficient of the heat transfer is maximized (Beaumont, Nagel, &\nSherman, 2002). Heat transfer may be satisfactory when the mold is new, but it will\ndecrease as the mold ages due to corrosion and lime deposits. The heat input from the\nplastic varies from location to location. It depends on the local part thickness and the\nplastic flow past the location. The mold wall temperature is not uniform, and hence the\ncooling rate will not be uniform. This can affect the melt viscosity, cavity pressure, or\nhold pressure (Johannaber, 1994). These problems can cause part defects such as sink\nmarks, void, bubble, warpage, short shot, flow line, surface defects, etc (Reinhold, 1986).\nIn many plastic injection molding machines, a hydraulic pressure is applied to\nforce the screw forward against the melted plastic. The plastic is forced to flow through\nthe nozzle, the sprue, the runner system, and finally into the cavity to form the part\n14\n(Beaumont, Nagel, & Sherman, 2002). This pressure ranges from 7,250 pounds per\nsquare inch (PSI) to 36, 260 PSI. Due to the way a plastic fills and solidifies within the\nmold cavity, pressure is not uniform throughout the part. The highest pressure occurs at\nthe gate; the lowest at the point in the cavity to fill last. The total volume of plastic in the\ncavity tends to be reduced as it cools and solidifies. Pack and hold pressures are used to\nprovide a compensation flow into the cavity to make up for lost volume as the plastic\nshrinks with decreasing temperature and while the gate remains open. However, to\nminimize volumetric shrinkage in the cavity, pressures are minimized to the point where\nthere is enough pressure remaining to provide a compensating flow but not to where it\ntends to over-pack the material near the gate. Insufficient pressures can cause part\nbrittleness, bubbles, sink marks, cracking, flashing, short shot, etc. (Reinhold, 1986).\nMaterial\nPlastic materials consist of two basic groups, thermoplastics and thermosets.\nThermoplastic materials are made of the linear or branched polymer units comprised of\nrepeating monomers. They comprise about 94% of the volume of the material used in the\nplastic industry and can be repeatedly heated, melted, and formed into a product.\n(Beaumont, Nagel, & Sherman, 2002).\nThermoplastics include amorphous polymers and semi-crystalline polymers.\nAmorphous resins have a wide processing temperature window. An amorphous polymer\nchain is randomly entangled. Typical amorphous polymers are polystyrene, acrylonitrile-\nbutadience-styrene (ABS), polyethersulfone, etc. Crystalline resins, on the other hand,\nhave a narrow processing temperature window. Their molecular chains lie side by side in\n15\na highly oriented fashion. Well-known examples of crystalline polymers are\npolypropylene, acetal, and polytetrafluoroethylene (Brydson, 1999)\nThermoset materials react chemically during processing to form a cross linked\nstructure. They cannot be melted and reprocessed (Reinhold, 1991). Polyimides,\npolyesters, and epoxies are examples of thermoset materials.\nThe transition between solid and liquid phases is a primary concern when\nprocessing plastics. In crystalline materials, the change from solid to liquid is abrupt and\neasily discernible. In an amorphous polymer, the material softens over a wide\ntemperature range. In polymer science, there is a point called the glass transition point\n(Tg). At temperatures below this point, plastic is stiff, stable, and behaves like a solid. In\nenvironments at temperatures above the Tg, the polymer will behave as a viscous liquid\n(Reinhold, 1991).\nWithin a material, there are two forms of energy (Reinhold, 1991) used to\nmaintain the physical structure. One is potential energy, which is a measure of the forces\nof attraction between the molecules. The other is kinetic energy, the energy of motion\ntending to separate the molecules. As more energy is put into the system, it turns into a\nliquid. That is where the potential and kinetic energies are equal.\nThermoplastic materials are made of strong covalent bonds (primary bonds) along\npolymer molecules and weaker secondary bonds (Van der Waals' forces) between\npolymer molecules (Beaumont, Nagel, & Sherman, 2002). A covalent bond exists when\ntwo atoms share the electrons in their outer shells in order to be stable. It has a\ndisassociation energy of 83 Kcal/mole. The Van der Waals' forces are electrostatic in\nnature and have a disassociation energy of 2-5 Kcal/mole. This is the energy that attracts\n16\nmolecules (cohesive energy), and is also the energy required to move a molecule a large\ndistance from its neighbor (Reinhold, 1991).\nIf the temperature of a plastic rises, the distance between the molecules increases.\nBecause the Van der Waals' forces decrease with the sixth power of the distance, the\nmolecules and their segments become more mobile. As these forces decrease in an\nexponential manner, there is a relatively narrow range in which the polymer changes\nfrom solid to a liquid. Therefore, polymer properties are quite temperature-dependent\n(Reinhold, 1991). For this reason, it is important to know the processing temperature\nrange for each plastic material in order to make good parts.\nDuring the injection molding process, a polymer mass is heated to a point at\nwhich it melts. In this molten phase, the material can flow and is forced into a cold mold\nwhere it will take the shape of the cavity (Beaumont, Nagel, & Sherman, 2002).\nHowever, a resistance to the flow of plastic material, exists due to viscosity (Brydson,\n1999), which is determined by shear stress divided by shear rate. Shear stress is a\nmeasure of the resistance to flow of the molecules sliding over each other (Reinhold,\n1991). Shear rate is the rate of velocity change of a flowing material.\nFor a specific polymer, the viscosity is dependent on temperature, molecular\nweight, and shear rate (Brydson, 1999). The higher the temperature, the lower the\nviscosity. The higher the molecular weight, the greater the entanglements and the greater\nthe melt viscosity. To maintain the same volume of material in the cavity, shot after shot,\nit is necessary to maintain the consistency of pressure and viscosity. However, the\ncharacteristics of materials are very complex (Beaumont, Nagel, & Sherman, 2002) and\nmost of them are non-constant. Non-constant viscosity is affected by the molecular\n17\nweight distribution of a polymer. This is a factor that makes it difficult for a process\nengineer to create an optimum process.\nState-of-the-art tools such as material databases and injection molding simulation\nsoftware are available to aid in the prevention of common problems. However, the non-\nconstant and complexity in characteristics of plastic material, combined with the process\nconditions in manufacturing, still result in unacceptable parts. Common part defects\ninclude contamination, burning, bubbles, splay, sink marks, voids, short shot, and\nwarpage (Reinhold, 1986).\nOperator\nReinhold wrote, \u201cIf the problem appears, disappears, or changes with the\noperators, look for the differences in actions of operators.\" (1986, p 665) Operator actions\nare the methods by which operators perform their jobs, and these actions can be related to\nmachine performance. This research reviews the operator experience-performance\nrelationship and uses it as a tool for analyzing the relationship between operator and scrap\nlevels. The only factor relating to operators that this research focuses on is work\nexperience based on the time spent on the job.\nMost literature on work experience has focused on time (79.5%) or job level of\nspecificity (68.2%), (Kolz, McFarland, & Silverman, 1998). Job experience was defined\nas a number of years an employee had worked in the same job for the same company.\nThis is the most commonly used definition of job experience. However, the jobs\nconsisted of a limited number of activities and rarely deviated from their routine; the\nnumber of years on the job was highly related to the number of times each work task was\nperformed. This agreed with the results showing that the number of times a person\n18\nperforms a job task is more strongly correlated with work performance than the time\nspent on the job. Data on the number of years on the job is readily available. Therefore, it\nis a more practical measure of experience rather than counting the number of times a task\nis performed. For this reason, the years on the job are likely to continue to be an\nimportant measure of experience.\nWork experience can be measured based on time, amount, or type (Kolz,\nMcFarland, 1998). In a time-based mode, experience is measured by the time spent in a\nparticular job, company, or given occupation. In an amount-based measure, the\nexperience of an operator can be measured as the number of times he/she performs a\nparticular task. Some studies measure an individual's work experience based on type.\nThis measurement mode defined experience as the degree of similarity between a\nperson's previous job and current job. The more similar the previous job is to the current\none, the more relevant work experience the person is presumed to bring to the current\njob.\nWork experience has also been measured at the task, job, or organizational level\nof specificity (Kolz, McFarland, 1998). At a task level of specificity, work experience is\nmeasured as the number of times an individual performed a task. At a job level,\nexperience is measured by the amount of time an individual spent on a job. At the\norganizational level of specificity, experience is measured as the time a person spent in\nan organization.\nThe results showed the strongest relationship between experience and\nperformance was when work experience was measured at an amount or task level of\nspecificity. The meta-analysis also reveals that the relationship between work experience\n19\nand job performance was positive regardless of the work experience measure used\n(Qui\u00f1ones, Ford, & Teachout, 1995). In other research, conducted by Hofman, Jacob, &\nBaratta, It was concluded that individuals may rate an employee's performance relating\nto his/her level of experience (1993). In addition, a meta-analysis by McDaniel, Schmidt,\n& Hunter (1998) found a mean corrected correlation of 32% between work experience\nand job performance across a number of occupations.\nIn the injection molding process, there are cases where the operators need to\nmanually operate the injection molding machines. For example, an operator has to open\nthe machine door to take the part out of the mold. They need to put in inserts, pick-outs,\nor apply mold release before starting the next cycle. Such actions by operators may result\nin changing the cycle time, residence time, temperature, and viscosity of the plastic\nmaterial. These problems will result in several part defects such as splay, burning, short\nshot, sink marks, surface defects, and so forth (Reinhold, 1986).\nChapter 3\nMethodology\nThis chapter will present the methods and procedures used to identify the scrap\nlevels associated with each major factor: machine, process, material, and operator. The\ninformation is based on scrap data collected from the scrap history log at the Phillips\nPlastics-Short Run facility. Data was organized and used to calculate average scrap levels\nrelated to each factor for this analysis. Results will be discussed to address the causes of\nthe problems that lead to scrap.\n20\n20\nResearch Design\nThe data from hundreds of production jobs was collected randomly from\nSeptember through November 2003. However, to assure that a good statistical sampling\nof data was gathered, at least three jobs for each machine were collected. To guarantee\nconsistency when gathering data, each production job was run through all shifts. The data\ncollected includes material type and key process parameters: melt temperature, mold\ntemperature, and pressure. Information about machine age and size or tonnage for all 32\ninjection molding machines was collected. Operator data for the facility's first, second,\nand third shifts was obtained according to their years of experience.\nData Collection\nWith permission from the plant and production managers, maintenance\ncoordinator, and supervisors of all three shifts at Phillips Plastics-Short Run facility, data\nwas collected from Short Run's scrap history logs for September to November, 2003. The\ndata included: operator initial and shift run, total good part, total scrap, and scrap caused\nby common problems like contamination, partial fill, void/bubbles, sink marks, burning,\nand splay. The material type, process key parameter, melt temperature, mold temperature,\nand pressure were also included with the data (Appendix A). In addition, age and tonnage\nfrom all injection molding machines on the floor were collected (Appendix B). The list of\noperators and their years of experience was also collected (Appendix C).\nFor this study the data was collected randomly from one hundred production jobs.\nHowever, to assure that a good statistical sampling of data was taken, at least three\nproduction samples were collected from each machine in an effort to include every\n\n\n21\nmachine in this research. Although limited data can be gathered from a real world\nbusiness, the information collected satisfied basic statistical and research requirements.\nAnalysis\nThe data collected was organized and grouped in a manner that addressed each of\nthe major factors covered in this research: machine, process, material, and operator. Then\nthe scrap was organized and analyzed based on the appropriate requirements of each\nfactor to yield their scrap levels.\nMachine\nTo analyze the scrap related with the injection molding machines, the percentage\nscrap related with each machine was calculated. Then the machines were organized into\nfour groups based on their tonnage and age at their associated scrap level. Group one\nconsisted of machines two years old or less. Group two included machines from three to\nfour years of age. Group three included machines from five to eight years of age. Finally,\ngroup four included the oldest machines, which consisted of eleven to fourteen years\noperation. The average scrap percentage related with these groups was calculated for an\noverview of the scrap level trends associated with the age of machines.\nTo analyze the scrap level related with the machine tonnage, machines were\norganized into following five groups:\n\u2022 Group one: 20-40-ton machines.\nof\n\u00b7\nGroup two: 55-80-ton machines.\n. Group three: 110-150-ton machines.\n.\nGroup four: 200-300-ton machines.\n\u2022\nGroup five: only one machine of 400 tons.\n22\n22\nThe average scrap level, in percentage, was calculated and yielded the scrap trend\nassociated with the machine tonnage from low to high.\nProcess\nThis area of research focused only on the analysis of key process parameters in\nthe injection molding process conditions. These parameters include melt temperature,\nmold temperature, and pressure.\nMelt Temperature\nThe melt temperature of the process in this research is classified into three levels:\nunder 500\u00b0F, 500-600\u00b0F, and over 600\u00b0F. The scrap levels associated with each range of\ntemperature were calculated to identify trends.\nMold Temperature\nThe mold temperatures are presented at three levels: under 100\u00b0F, 100-200\u00b0F, and\nover 200\u00b0F. The scrap levels related with the mold temperatures were calculated for each\nrange to identify trends.\nPressure\nThe pressures were organized into three ranges: pressures under 10,000 PSI,\n10,000-20,000 PSI, and over 20,000 PSI. Good and scrap part rates associated with each\nrange of pressure were calculated to get the average scrap percentage. The results identify\na trend of scrap levels associated with the process pressure.\nMaterials\nThe majority of materials used in the one hundred production samples collected\nfor this research consisted of the following materials: Calibre, Radel, Ultem, Cycolac,\nand RTP. Other materials used with less frequency in the production run were Lexan,\n23\n23\nBayblend, LNP, Lustran, Cycoloy, Zytel, Santoprene, etc. Only the first five materials are\nanalyzed for this study. The data related with each of these five materials was organized\nto yield their specific scrap levels. The scrap rate related to common defects, such as\ncontamination, partial fill, void/bubbles, burn/streaks, sink marks, and splay, was also\ncalculated to identify relationships between these material defects.\nOperator\nTo analyze the scrap associated by operator; this research organized operators on\neach production line based on their years of experience. Operators were divided into three\ngroups: less than one year of experience, one to two years of experience, and over two\nyears of experience. The scrap percentages for each operator were calculated, and then\nidentified for each group. The results show a relationship exists between scrap levels and\noperator experience. The average scrap levels associated with operators on each\nproduction line and shift were also calculated and analyzed.\nChapter 4\nResults\nThis chapter presents a detailed overview of the scrap levels associated with\nmachine, process, material, and operator. The scrap levels are related to the tonnage or\nage of injection molding machines. The scrap percentages are also related to plastic\nprocess key parameters: melt temperature, mold temperature, and pressure. The scrap\nlevels associated with the five most common materials, used by Phillips Plastics-Short\nRun facility, were used for this research. And the scrap associated with operators, on each\nshift and production line, was based on their years of experience. These results, based on\n24\n24\nthe analysis of actual scrap data, can be used by Phillips Plastics-Short Run to quickly\npredict the levels of scrap that may occur at each work station. Therefore, they will be\nable to better control the generation of scrap at their work stations.\nThe scrap rates associated with machine age are shown in Figure 1. The scrap\nrates related with using the newest machine (1-2 yrs) were quite high. The scrap rates\nwere lowest using the 2-4 year old machines. Scrap level related with using the machines\nover 5 years of operating age were similar, at above 5%. Except for the machines under 2\nyears of age, scrap level tends to increase as machines get older.\nScrap (%)\n18\n16\n14\n8 6420\n12\n10\n8\n16.5\nScrap % Related to Machine Age\nCO\n6\n5.1\n5.4\n3.6\n4\n2\n0\n1-2\n2-4\n5-8\n11-14\nAge (yr)\nFigure 1 Scrap Percentage Related to Machine Age\nThe results show in figure 2 that the scrap level increases as the tonnage of the\nmachine increases. The machines less than 40 tons in size resulted in the lowest level of\n25\nscrap. Machines from 55-80 tons in size resulted in 5.9% of scrap. Machines 110-150\ntons in size resulted in a scrap level of 5.8%. Machines 200-300 tons in size resulted in a\n7% scrap level. Only one 400-ton machine operates on the shop floor. The scrap related\nwith it was high, up to 34%. However, just one machine at that clamp force does not\nresult in a good statistical sampling. In general the trend of scrap increases as the\nmachines' tonnage gets higher.\nScrap (%)\nScrap Level Related to Machine size\n40%\n30%\n20%\n10%\n5.90% 5.80%\n7.00%\n4.20%\n\u2610\n0%\n34.0%\n20-40 55-80 110-150 200-300 40\nSize (ton)\nFigure 2 Scrap Percentage Related to Machine size\nFigure 3 shows the scrap levels associated with the melt temperature. In the range\nof melt temperatures under 500\u00b0F, the scrap level was 0.8%. The scrap jumped to a\nhigher level of 3.9% with the melt temperatures between 500-600\u00b0F. And with the melt\ntemperature over 600\u00b0F, the scrap again increased to a level of 5.3%. In the common\nrange for melt temperature (under 500\u00b0F), the scrap was at an acceptable level. At higher\n26\n46\ntemperatures, the scrap levels increased. The apparent trend is that the higher the melt\ntemperature, the higher the scrap level.\nScrap (%)\nScrap Level Related to Melt Temperature\n6%\n5.3%\n5%\n3.9%\n4%\n3%\n2%\n0.8%\n1%\n0%\nunder 500\n500-600\nover 600\nTemperature (\u00b0F)\nFigure 3 Scrap Levels Associated to Melt Temperature\nFigure 4 presents the average scrap percentage results associated with mold\ntemperatures. At lower mold temperatures (under 100\u00b0F), the scrap level reaches a very\nhigh level of 8.6%. The scrap percentage reduces to about half (at 4.7%) in the mid-range\nof mold temperatures (100 \u00b0F -200\u00b0F). At higher mold temperatures (200\u00b0F and up) the\nscrap level was 2.9%. Therefore, it is apparent that the higher the mold temperature, the\nlower the scrap level.\nScrap (%)\nScrap Level Related to Mold Temperature\n10%\n8.6%\n8%\n6%\n4.7%\n4%\n2.9%\n2%\n0%\nunder 100 100-200 over 200\nTemperature (\u00b0F)\nFigure 4 Scrap Levels Related to Mold Temperature\nAs Figure 5 shows, at the lower range of pressures (under 10,000 PSI), the scrap\nlevel was very high at 9.3%. At the mid-range of pressures (10,000-20,000 PSI), the\nscrap was lower at 3.7%. The process at higher pressures, 20,000 PSI or higher, resulted\nin a scrap level of 4.8%. In general, the scrap level was lowest at pressures of 10,000PSI\nto 20,000PSI. The processes that require higher pressures produce higher level of scrap.\nHowever, at a lower pressure (under 10,000 PSI) the scrap level also becomes very high.\nThe five most commonly used materials for this research were Calibre, Radel,\nUltem, Cycolac, and RTP. Calibre was the most commonly used material in the data\ncollected for this research. Table 1 show that the over all scrap level associated with\nusing this material was 3%. The contamination level related with this material resulted in\nthe highest specific scrap level. The scrap level due to burning, void/bubble, and splay\nwere also monitored. Scrap levels due to partial fill were very low using this material,\n27\n28\nand scrap due to sink marks was nonexistent. Generally, Short Run should pay very close\nattention to contamination of material, then consider scrap due to burning, void/bubble,\nand splay when using this material.\nScrap(%)\nScrap Level Related to Pack Pressure\n10%\n9.3%\n8%\n6%\n4.8%\n3.7%\n4%\n2%\n0%\nunder 10,000 10,000-20,000 over 20,000\nPressure (PSI)\nFigure 5 Scrap Level Related to Pack Pressure\nTotal Parts Contamination Partial Void/Bubble Burn/Streak Sink Splay Scrap Rates\n203,707\n1844\n52\n256\n318\n0 203\n3.0%\nTable 1 Scrap Level Related to Calibre\nRadel was the second most commonly used material at the Short Run facility, as\nidentified by this study. The over all scrap level related with the use of this material was\n29\n4.6%. Splay resulted in the most scrap. Scrap due to contamination was the second most\nimportant issue when using this material.\nTotal Parts Contamination Partial Void/Bubble Burn/Streak Sink Splay Scrap Rates\n8,387\n59\n14\n7\n1\n0\n138\n4.6%\nTable 2 Scrap Level Related to Radel\nUltem was another material used often by Short Run. The study shows that the\nover all scrap level related with using this material was 5.3%. Most scrap problems\noccurred due to splay. The next problem occurred due to contamination scrap level.\nVoid/Bubble and burn/streak scrap levels were much lower than splay and contamination\nscrap levels. For this material, scrap levels due to sink marks and partial fill were very\nlow. Therefore, splay and contamination should be closely monitored when using Ultem.\nTotal Parts Contamination Partial Void/Bubble Burn/Streak Sink Splay Scrap Rates\n102,519\n245\n26\n97\n63\n18 840\n5.3%\nTable 3 Scrap Level Related to Ultem\nTable 4 shows that the over all scrap level associated with using Cycolac was\n2.5%. The most common scrap problem presented with using this material proved to be\nsplay. Contamination was far less a problem, but nevertheless also needs to be\nconsidered. Similar to Ultem, when Cycolac was used, the splay scrap levels were\nhighest.\n30\nTotal Parts Contamination Partial Void/Bubble Burn/Streak Sink Splay Scrap Rates\n39,056 72 15 3 2 3 420 2.5%\nTable 4 Scrap Level Related to Cycolac\nThe over all scrap level related with using RTP is presented in Table 5, and was\nfound to be 9.3%. The table also shows the biggest problem was scrap due to\ncontamination. Sink marks and partial fill scrap levels should also be considered. Splay,\nbubble, and burning scrap levels were rarely problematic when using this material.\nTherefore, most scrap reduction efforts should be focused on contamination during the\nuse of RTP.\nTotal Parts Contamination Partial Void/Bubble Burn/Streak Sink Splay Scrap Rates\n13,184\n454\n178\n0\n13\n193\n5\n9.3%\nTable 5 Scrap Level Related to RTP\nFigure 6 shows the scrap percentage related with three lines of 1st shift production\nat Short Run. The scrap level on A-line was found to be 2.9%. On B-line, the scrap\npercentage jumped to a higher level of 4.6%. And on C-line, the scrap level resulted in a\nrate of 3.5%. These scrap rates resulted in an average scrap rate of 3.7% for the first shift.\n\n\nScrap(%)\n5%\nScrap Level Related to 1st Shift Operator\n4.6%\n4%\n3.5%\n2.9%\n3%\n2%\n1%\n0%\nA-Line\nB-Line\nC-Line\nProduction Line\nFigure 6 Scrap Levels Related to 1st Shift Operators\nThe result in figure 7 show that the scrap percentages for 2nd shift as related to\noperators were: 6.8% (A-line), 9.6% (B-line), and 7.5% (C-line). B-line, again garnered\nthe highest percentage of scrap level. The next lower scrap level was C-line, and at yet a\nlower scrap level was the A-line. On average, the scrap level on 2nd shift was 8%.\nThe results for 3rd shift, shown in figure 8, at Short Run were: 4.7%, 8%, and\n4.2% associated with A, B, and C-lines, respectively. The highest level of scrap still\noccurred with the B-line. However, A and C-lines were extremely close to each other, at\n4.7% and 4.2%, respectively. The results also show an average scrap percentage of 5.6%\nfor 3rd shift.\n31\nScrap(%)\nScrap(%)\n12%\nScrap Level Related to 2nd Shift Operator\n9.6%\n10%\n7.5%\n8% 6.8%\n6%\n4%\n2%\n0%\nA-Line\nB-Line\nC-Line\nProduction Line\nFigure 7 Scrap Levels Related to 2nd Shift Operators\n10%\n8%\n6%\n4%\n2%\nScrap Level Related to 3rd Shift Operator\n4.7%\n8.0%\n\u0644\u062a\n4.2%\nili\n0%\nA-Line\nB-Line\nC-Line\nProduction Line\nFigure 8 Scrap Levels Related to 3rd Shift Operators\nThe average scrap level for each production line on all three shift is shown in\nfigure 9. It can be seen that the average percentage scrap level for the A-lines on all shifts\nwas 3.7%. The B-lines reached an 8% average and the C-lines yielded a 5.6% scrap level\n32\n32\n33\nassociated with all three shifts. Generally, Figure 9 shows the highest scrap level occurs\nfor B-line, lowest on A-line, and C-line falls in between.\nScrap(%)\nScrap Level Related to each Production Line\n10%\n8%\n8.0%\n5.6%\n6%\n3.7%\n4%\n2%\n0%\nA-Lines B-Lines\nC-Lines\nProduction Line\nFigure 9 Scrap Levels Related with Each Production Line\nData gathered on operator experience for A-Line 1st shift is shown in figure 10.\nThis data was organized into three groups based on their years of experience:\n\u2022 Group 1: includes the operators that had less than a year experience.\nGroup 2: includes the operators that had 1 to 2 years of experience.\n\u2022 Group 3: includes the operators that had over 2 years of experience.\nFigure 10 shows only two groups because there were no operators with less than one year\nof experience. The group of operators having from 1 to 2 years experience resulted in a\nscrap level of 3.2%. The other group having over 2 years of experience was associated\nwith a scrap level of 2.1%. Figure 10 provides evidence that operators with more than 2\nyears of experience generated less scrap.\nScrap (%)\nScrap Level Related to Operator Experience\n5.00%\n4.00%\n3.2%\n3.00%\n2.10%\n2.00%\n1.00%\n0.00%\n1-2 Years Over 2 Years\nExperience (Yr)\nFigure 10 Scrap Levels Related to 1st Shift's A-Line Based on Experience\n34\n==\nFigure 11 shows the scrap levels related to three operator groups for 1st shift B-\nline. This study was organized into the following groups:\n\u2022 Group 1: operators that had less than 1 year of experience resulted in 3% scrap.\n\u2022\nGroup 2: operators that had from 1- 2 years of experience resulted in 3.4% scrap.\n\u2022 Group 3: operators that had over 2 years of experience resulted in 4.4% scrap.\nSurprisingly, Figure 11 shows that the more experienced operator generated the higher\nscrap level. The fact that a more experienced operator generated the highest scrap level\nwas an unexpected trend.\nScrap(%)\nScrap Level Related to Operator Experience\n5%\n4.4%\n4%\n3.4%\n3.0%\n3%\n2%\n1%\n0%\nUnder 1 year 1-2 years\nOver 2 years\nExperience (Yr)\nFigure 11 Scrap Levels of 1st Shift's B-Line Operators Based on Experience\nFigure 12 shows the relationship between operator experience and scrap levels for\n1st shift C-line. This data was organized into the following groups:\n\u2022\n\u2022\nGroup 1: operators that had under 1 year of experience resulted in 5.2% scrap.\nGroup 2: operators that had 1-2 years of experience resulted in 3.7% scrap.\n\u2022 Group 3: operators that had over 2 years of experience resulted in 2.0% scrap.\nThe results presented in Figure 12 show that the higher the operator's experience, the\nlower the scrap level. Based on the operator capabilities due to experience, this is an\nexpected trend in the generation of scrap.\n35\nScrap(%)\nScrap Level Related to Operator Experience\n6%\n5.2%\n5%\n3.7%\n4%\n3%\n2%\n1%\n0%\n2.0%\nUnder 1 year 1-2 years Over 2 years\nExperience (Yr)\nFigure 12 Scrap Levels of 1st Shift's C-Line Operators Based on Experience\nFigure 13 presents the scrap levels related to 2nd shift A-line, based on operator\nexperience. The data was organized in the following groups:\n\u2022\nGroup 1: operators that had under 1 year of experience resulted in 9.4% scrap.\nGroup 2: operators that had 1-2 years of experience resulted in 7.2% scrap.\n\u2022 Group 3: operators that had over 2 years of experience resulted 1.3% scrap\nThe trend shows a good relationship between operator's experience and their scrap\nlevels. As expected, the more experienced operators produced less scrap.\n36\n36\nScrap(%)\n12%\nScrap Level Related to Operator Experience\n10% 9.4%\n8%\n6%\n4%\n2%\n0%\n7.2%\n1.3%\nUnder 1 year 1-2 years Over 2 years\nExperience (Yr)\nFigure 13 Scrap Levels of 2nd Shift's A-Line Operators Based on Experience\nFigure 14 presents two groups of operators for 2nd shift B-line and their related\nscrap percentage, based on their experience. In this particular case, there was no group of\noperators with 1 to 2 years of experience. The results show a scrap level of 6.3% for the\ngroup of operators having less than 1 year of experience. The group of operators having\nover two years of experience generated scrap at a 11.3% level. Once again, the trend here\nis that the higher experienced operators produced more scrap. This trend is counter\nintuitive to what was expected, scrap levels should decrease as operator experience\nincreases.\nFigure 15 presents the scrap levels associated with two groups of operator for 2nd\nshift C-line. One group had from 1 to 2 years experience. The scrap level for this group\nresulted in an 11.2% scrap level. The other group consisted of operators with over 2 years\nof experience. The scrap level for this group resulted in a 6.6%. Figure 15 also shows an\n37\nScrap(%)\nexpected trend. The 2nd shift C-line operators with the most experience generated the\nleast scrap.\nScrap(%)\nScrap Level Related to Operator Experience\n14%\n11.3%\n12%\n10%\n8%\n6.3%\n6%\n4%\n2%\n0%\nUnder 1 year\nOver 2 years\nExperience (Yr)\nFigure 14 Scrap Levels of 2nd Shift's B-Line Operators Based on Experience\n14%\nScrap Level Related to Operator Experience\n12%\n10%\n8%\n6%\n4%\n2%\n0%\n11.2%\n6.6%\nli\n1-2 years Over 2 years\nExperience (Yr)\nFigure 15 Scrap Levels of 2nd Shift's C-Line Operators Based on Experience\n38\n39\nFigure 16 shows the scrap levels related with the operators on 3rd shift A-line. The\ndata was organized by operator experience into the three following groups:\n\u2022\n\u2022\nGroup 1: operators that had under 1 year of experience resulted in 2.2% scrap.\nGroup 2: operators that had 1-2 years of experience resulted in 5.7% scrap.\nGroup 3: operators that had over 2 years of experience resulted 5.2% scrap\nThese results are inconsistent in regards to the expected relation between operator\nexperience and scrap levels. The group of operators having the least experience resulted\nin the lowest level of scrap. The group having 1-2 years of experience resulted in the\nhighest level. The group of operators having over 2 years of experience resulted in a\nslightly lower scrap level. In general, the trend witnessed here is that the more\nexperienced operators made more scrap.\nScrap(%)\nScrap Level Related to Operator Experience\n8%\n7%\n5.7%\n6%\n5.2%\n5%\n4%\n3%\n2.2%\n2%\n1%\n11\n0%\nUnder 1 year\n1-2 years\nOver 2 years\nExperience (Yr)\nFigure 16 Scrap Levels of 3rd Shift's A-Line Operators Based on Experience\n40\nFigure 17 illustrates the relationship between operators and their related scrap\nlevels for 3rd shift B-line. An irregular trend can be witnessed here. The least and most\nexperienced groups generated similar levels of scrap, 11.0% and 10.4% respectively. The\ngroup of operators that had 1-2 years of experience resulted in the scrap percentage of\n4.3%. Again, these results do not show a consistent relationship between operator's\nexperience and scrap.\nScrap(%)\nScrap Level Related to Operator Experience\n14%\n12% -11.0%\n10.4%\n10%\n8%\n6%\n4.3%\n4%\n2%\n0%\nUnder 1 year 1-2 years Over 2 years\nExperience (Yr)\nFigure 17 Scrap Levels of 3rd Shift's B-Line Operators Based on Experience\nFigure 18 presents the scrap levels related with three groups of operators on 3rd\nshift C-line. Based on their experience, this data is summarized below:\n\u2022\n\u2022\n\u2022\n4.6% scrap level for the group with under a year of experience\n3.0% scrap level for the group with 1-2 years of experience\n10.8% scrap level for the group that has over 2 years of experience\n\n\n41\nThe first two groups depict an expected trend about the relationship between operator's\nexperience and scrap percentage. However, as seen previously, the highest scrap level of\n10.8% was generated by the most experienced operators. This result again is counter\nintuitive to what is expected in the generation of scrap and based on operator experience.\nScrap(%)\nScrap Level Related to Operator Experience\n12%\n10%\n8%\n6%\n4.6%\n4%\n3.0%\n2%\n0%\n10.8%\nUnder 1 year 1-2 years Over 2 years\nExperience (Yr)\nFigure 18 Scrap Levels of 3rd Shift's C-Line Operators Based on Experience\nValidity\nThe research presented here is based on real data gathered from a Phillips\nPlastics-Short Run facility. The data was organized and analyzed based on the statistical\nprocess control that has been applied in the real-world manufacturing operations at the\nShort Run facility. Due to the fact that this Short Run facility continually monitors their\nscrap levels, it is possible to have a high level of confidence in the data gathered.\nThis research can be used by Short Run production to better understand why scrap\nlevels vary for production jobs run with the specific process parameters identified in this\n42\nwork. Short Run production personnel, such as machine operators, mold technicians, and\nsupervisors, will be able to use this work to minimize the generation of scrap. Every\nproduction job will be prepared in a way that will reduce the problems by\nacknowledgement of the related scrap levels with the type of injection machine, process,\nmaterial, and operator involved in each production job at specific work stations.\nDiscussion\nThis section of the research discusses the reasons for the varying scrap levels seen\nin production related to the machine, process, material, and operator. By understanding\nthe root causes of the problem, Short Run can determine more predictable ways in using\nthese production resources, ultimately resulting in a better control of scrap generation.\nMachines\nThe findings of this research indicated that the scrap levels vary based on machine\nage and tonnage. The data gathered supports the concept that the longer a machine is in\noperation, the less efficient it becomes. The machine response time becomes inconsistent\nwith age. Therefore, the machine itself changes the processing conditions, resulting in an\nincrease of scrap. At Short Run, an injection molding facility, scrap increased from a rate\nof 3.6% to 5.4% with machines that had been in operation from 2 to 14 years. An\nunexpected result identified when evaluating the production data was that newer\nmachines (1-2 years old) resulted in a very high scrap level of 16.5%. These newer\ninjection molding machines included a vertical machine and a 400-ton machine. Most of\nthe time, these machines must run in a semi-automatic function mode. The reason for the\nhigher scrap levels associated with using these machines was due to the way operators\noperate the machines. Differing operator actions made it difficult to keep the cycle times\n43\nconsistent. This inconsistency in cycle time was the reason for processing condition\nchanges, resulting in an increase of scrap generation.\nAnother factor identified in this research that lead to scrap generation involved\nmachine size or tonnage. This factor is an indicator of the machine's strength. Jobs\nrequiring a stronger machine are usually complicated in shape, mold design, and part\ndesign. The polymer flow pattern of larger parts is usually quite complex, requiring\nhigher processing pressures to fill mold cavities and greater machine strength. Due to the\ncomplexity of these larger parts, they must be made in a semi-automatic function mode\nbecause of the necessity of loading the pick-outs, inserts, etc. For these reasons it was\ndifficult for the operators running these jobs to keep cycle times consistent. When one or\na combination of these machine factors existed in production, the scrap level may have\nrisen. Based on the recently discussed items it is possible to make a rationalization as to\nthe reason that the use of larger machines can often lead to increased levels of scrap\ngeneration.\nProcess\nThis portion of the research focuses on analyzing the key process parameters of\nmelt temperature, mold temperature, and pressure. Though process time can lead to other\nissues of concern, it is easy to control. Problems can occur based on how machines\nrespond to each change in setting. If the machine does not respond well, possibly due to\nits age, the time setting may not work resulting in molding problems.\nMelt Temperature\nThe research findings shows that the higher the melt temperature, the higher the\nscrap level. Most of the melt temperature induced problems can be related to melt\n44\ntemperatures higher than 500\u00b0F, especially when the temperature was higher than 600 \u00b0F.\nThis occurs because most thermoplastics require that the processing temperatures be less\nthan 500\u00b0F. Melt temperatures higher than 500\u00b0F are considered to be excessively high\ntemperatures. With the higher temperatures, the covalent bonds of the resin molecules are\nmore likely to degrade in molded materials. Therefore, the chances of splay, burning, or\nshrinkage problems occurring may be higher than when processing at lower temperatures.\nIt is for this reason that higher melt temperatures typically lead to higher scrap levels.\nMold Temperature\nThe research findings show that when higher mold temperatures were used the\nresulting scrap level was lower. Scrap levels were lowered because the use of higher\nmold temperatures lead to a reduction in material shear rate. Therefore, the in-mold\nmaterial flows occur more smoothly. Problems such as short-shot, surface finish, over-\nsizing, etc, may be reduced by increasing mold temperatures.\nPressure\nMachine pressures between 10,000 PSI to 20,000 PSI are applied for processing\nof most thermoplastics. Most parts require pressures in this range to guarantee adequate\nin-mold material flow. Pressures should be great enough to keep the cavity filled but not\nso much that over-packing occurs. If a particular process requires a higher pressure (over\n20,000 PSI) to meet the requirements of a complex flow pattern, it usually causes higher\nmaterial shear rates or more material crystallization around the gate. These problems can\ncause part splay, crack, or warpage. For this reason higher machine pressures typically\nlead to higher scrap rates. The results also show a trend towards high scrap levels (9.3%)\nrelated to using machine pressures that are less than 10,000 PSI. Too low a machine\n45\npressure usually leads to slower material fill rates and improperly filled or packed parts.\nNot enough plastic is added during the mold fill process to compensate for part shrinkage\nwhen too low a pressure is used. This can be the reason of material short shot, shrinkage,\nor sink marks.\nMaterials\nIn a real-world plastics manufacturing facility, like Short Run, quite a few\nmaterial types are used to produce plastic parts. In an attempt to keep the analyses\nperformed for this research manageable, this study covers the five most commonly used\nmaterials at Short Run. These materials consist of Calibre, Radel, Ultem, Cycolac, and\nRTP. The scrap levels associated with each of these five materials, listed from worst to\nleast, were as followed:\n.\nRTP at 9.3% scrap\n\u2022\nUltem at 5.3% scrap\n\u2022\nRadel at 4.6% scrap\n\u2022\nCalibre at 3.0% scrap\n\u2022 Cycolac at 2.5% scrap\nCycolac\nIt is\neasy to understand why the use of Cycolac as a material resulted in a low\nscrap level of 2.5%. By examining the data collected, it can be seen that the processing of\nCycolac required a temperature range from 400 \u00b0F to 500\u00b0F. As discussed earlier in the\nliterature review section, this is a common processing temperature range that is\napplicable for most plastic materials.\n46\nThe mold temperature required for processing this material ranged from 50 \u00b0F to\n150\u00b0F. Most of the productions runs, based on the data collected for this research, were\ncompleted using a mold temperature of about 100 \u00b0F. This was another advantage of using\nthis type of material. The pressures applied for processing this material ranged from\n5,500 PSI to 21,000 PSI. For most of the productions runs pressure ranged from about\n10,000 PSI to 15,000 PSI. This range of processing pressures is common for pack and\nhold pressures used to process most plastic materials. In general, the key parameters used\nto control the processing of Cycolac, melt temperature, mold temperature, and pressure,\nwere within common process conditions. Therefore, the 2.5% of scrap associated with\nprocessing Cycolac can be explained by looking at its key processing parameters.\nBecause Cycolac is hygroscopic it is susceptible to the occurrence of splay,\ncontamination, or short shot when used to produce parts, and ultimately is the primary\nreason for producing degraded and undesired parts.\nCalibre\nCalibre is the most commonly used production material at Short Run. A scrap\nlevel of 3.0% was seen with the use of Calibre. The most common problems to occur\nwhen using this material consisted of contamination, burning, splay, and voids/bubbles.\nThese issues can be explained by taking a look at material processing parameters. The\nmelt temperature used to process this material ranged from 530\u00b0F to 600\u00b0F. Mold\ntemperatures ranged from 70\u00b0F to 190\u00b0F. Pressures ranged from 11,000 PSI to 25,000\nPSI. In comparing these processing parameters to those used for Cycolac, it can be seen\nthat the values for all three parameters had increased slightly. These results imply that the\nuse of a higher mold temperature resulted in a lower scrap level. Therefore, it can be seen\n47\nthat the mold temperature-scrap relation is not an issue when processing Calibre.\nHowever, the melt temperatures and pressures required for processing Calibre can be\nrelated to the generation of scrap. The higher that the melt temperature or pressures are\nthe higher the resulting scrap levels. Again, by closely examining the process parameters\nused to produce Calibre parts, it is possible to explain the low scrap level of 3%.\nRadel\nThe use of Radel in producing parts resulted in a scrap level of 4.6%. The most\ncommon reason for scrap generation was the occurrence of splay. The next most common\nreason for problems occurred due to contamination and short shot. It is easy to\nunderstand why the scrap levels related to using this material were higher than the scrap\nlevels that occurred when using Cycolac or Calibre. By examining the process melt\ntemperature, ranging in value from 670\u00b0F to 740\u00b0F, required for producing Radel parts, it\nis possible to gain insight into the reasons for an increase in scrap generation. It can be\nseen that fairly high temperatures were used to produce Radel parts, causing the material\nin the barrel to easily degrade or burn. For this reason, an increase in the occurrence of\nsplay or contamination was seen with the use of Radel, resulting in higher scrap levels.\nA similar situation to that seen with the processing of Radel also applied to the\nprocessing of Ultem. The processing of this material resulted in a 5.3% scrap level.\nUltem processing also required the use of higher temperatures, ranging in value from\n680\u00b0F to 740\u00b0F. However, the pressures used to process Utem ranged in values from\n7,500 PSI to 28,000 PSI. This resulted in the use of a greater pressure range than that\nused to process Radel. The use of a greater pressure range led to higher in-mold flow\nrates (Reinhold, 1991). A phenomenon such as this can lead to part defects such as splay\n48\nor contamination. These results tend to support the premise that higher pressures result in\nhigher scrap levels.\nRTP\nThe highest scrap level as seen in the collection of data for this research resulted\nin the use of a material called RTP. When processing RTP, a polycarbonate material, the\nrecorded melt temperatures ranged from 490\u00b0F to 530\u00b0F. The mold temperatures used to\nprocess RTP ranged from 70\u00b0F to 180\u00b0F and the pressures ranged from 4,000 PSI to\n17,000 PSI. If the processing of this material is compared to that of Cycolac or Calibre, it\ncan be seen that the processing parameters are identical. Since this is the case, why does\nthe processing of RTP result in a very high scrap level, consisting mostly of defects due\nto contamination, sink marks, and short shot? Based on the literature review performed\nfor this research, the processing temperature of RTP or polycarbonate, ranges from a\nvalue of 446\u00b0F to 572\u00b0F. At processing temperatures just above this range, the material\ndegrades quite rapidly (Brydson, 1999). The melt viscosity of this material is very high.\nDuring processing, this material's flow path ratios fall in the range of 30:1 to 70:1. These\nratios are substantially less than many of the more general-purpose thermoplastics such as\npolypropylene at 175:1 to 350:1, ABS at 80:1 to 150:1, and Nylon 6/6 at 180:1 to 350:1.\nThe characteristics of RTP just discussed can be used to try and explain the cause\nof part defects like contamination, sink marks, or short shot. The high scrap level could\nbe related to the injection machines used to produce RTP products. Based on the data\ncollected for this research, most of the time these products were molded using injection\nmolding machines sized in the range from 200 ton to 400 ton. These are considered high\nclamp-force machines at the Short Run facility. These injection molding machines are\n49\ntypically used to produce complicated part designs with complex flow patterns. Due to\nthe difficulty in moving material through these molds, several types of part defects can\noccur. Another reason for high scrap levels could be due to the operator's actions. Again\nbecause of the complexity of some part designs, machine operators often have a difficult\ntime in identifying the correct processing parameters. This difficulty in setting processing\nparameters results in inconsistent cycle times or longer material residence times. These\nproblems can be the reasons for higher scrap when RTP is used.\nOperators\nThe scrap level associated with each production shift can be explained by relating\nthem with operators' experience. Most operators working on the A and B lines of 1st shift\nhave over two years of machine operation experience. These experienced operators are\nuniformly distributed throughout the lines. The 1st shift generated a total scrap level of\n2.9%. This trend is similar to that predicted by the reading discussed in the literature\nreview; the more experienced an operator is the less scrap generated. The scrap level of\n4.6% associated with 2nd shift was higher than the 3.5% scrap level associated with 3rd\nshift. This phenomenon cannot be explained based solely on the experience of each\noperator because the operators on 2nd shift and 3rd shift were uniformly distributed with\nsimilar years of experience. Also with each shift on the same production line, the same\nmachines, processes, and materials were used. The highest scrap level, which was seen\non 2nd shift, can be hypothesized as being a function of something other than operator\nexperience. By applying practical reasoning to identifying the causes of this problem, it\nappears that this higher scrap level could have been caused by outside non-machine\nrelated variables such as time of day, operator alertness, or non-machine related\n50\n50\nresponsibilities. These conditions may have an extreme affect on the performance of a 2nd\nshift machine operator.\nThe scrap levels associated to the A-line, B-line, and C-line can be explained\nbased on machine use and the operators' actions. All A-line injection machines are\nsmaller in size, from 20 to 40 tons. Most of these machines are relatively new and\nmodern. The production jobs run on these machines are typically operated in an auto-\nfunction mode, hence consistent cycle times are possible. Also, since these are small\nmachines, they are used to produce small parts. Due to the smaller sized parts, the flow\npatterns are simple and result in less part fill type problems. For these reasons the A-line\nscraps levels are low for all three shifts.\nMost B-line scrap generation levels can be explained based on machine age and\nuse. Almost half of the B-line machines are 8 to 14 years old, fairly close to the end of\ntheir life cycle. In addition, the B-line also consists of a vertical injection molding\nmachine that requires all jobs be operated in a semi-automatic mode. Problems typically\nencountered in using this machine can usually be correlated back to the operator. Many\nof the jobs produced on the B-line also require that inserts and pick-outs be loaded into\nthe mold prior to starting a cycle. Therefore, operator experience is a key factor when\nproblems arise in the B-line.\nThe C-line scrap level of 5% was better than B-line's scrap level at 7.4%. This\ndifference in scrap levels can also be explained based on machine age and use. Except for\none 12-year-old machine, most of the machines on C-line are relatively new. These\nnewer machines are more dependable and require less operator interaction during\n\n\n51\noperation. For these reasons the C-line had a better scrap level rate on all shifts than the\nB-line.\nThe scrap levels associated with the production of the 1st, 2nd, and 3rd shifts of A-\nline were 2.9%, 6.8%, and 4.7%, respectively. Because 1st shift of A-line had more\nexperienced operators than the 2nd and 3rd shifts, operator inexperience was a reasonable\ncause for the higher scrap levels seen on the later shifts.\nB-line 1st, 2nd, and 3rd shift scrap levels were 4.6%, 9.6%, and 8.0%, respectively.\nSome of the differences in scrap levels between shifts can be explained due to the more\nexperienced operators working on 1st shift, most having more than two years of\nexperience, and ultimately resulting in the lowest scrap level. The 2nd shift B-line scrap\nlevel was higher than that of the 3rd shift. Again, this phenomenon cannot be explained\nbased solely on operator experience. Both shifts consist of a similarly uniform\ndistribution of experienced operators. Since these shifts consist of the same production\nlines, machines, processes, and materials, the variation in scrap levels between the B-line\n2nd and 3rd shifts may be caused by non process related variables, as presented earlier in\nthis section.\nThe scrap levels associated with all of C-line shifts, in chronological order, were\n3.5%, 7.5%, and 4.2%. The distribution of operator experience was identical for all shifts.\nTherefore, it appears once again that operator experience was not a factor in the scrap\nlevel variances seen between all of the C-line shifts. These variances in all probability\noccurred due to non-process dependent variables, as presented earlier.\n52\nChapter 5\nConclusion\nThe generation of scrap is an important problem that must be addressed by\ntoday's businesses. It needs to be controlled for the purpose of lowering costs in an\nattempt to gain a competitive advantage in business. Despite several methods currently\nused to control scrap generation, a simple way that can help Phillips Plastics-Short Run\ncontrol scrap generation is to understand the scrap levels as they relate to injection\nmolding machines, processes, materials, and machine operators. By understanding the\neffects of these process parameters it becomes possible for a production facility to gain\ninsight into the generation of scrap due to the production conditions for each workstation\nas it applies to each job. In this way production facilities can control scrap levels by\nmaking a better informed decisions in the selection of injection machine, process,\nmaterial, and operator required to produce a particular job, ultimately minimizing scrap\nlevels.\nThis research presented scrap level data associated to the major processing\nparameters: injection machine, process, material, and operator. The analysis performed in\nthis work was subdivided further by examining the scrap levels associated to several\nproduction lines and work shifts. Understanding the reasons for variance in production\nlines, shifts, or operators can help Short Run plan for the rise in scrap levels for specific\nproduction jobs. Therefore, it will become possible for Short Run to better adjust\nproduction conditions to minimize scrap generation.\nThis research analyzes the scrap generated from 100 samples jobs collected at\nPhillips Plastics Short Run facility. Most of the research results can be explained based\n53\non an understanding of process parameters in regards to injection machine, process,\nmaterial, or operator assigned to the production job. Some results, such as the differences\nseen between scrap levels from 2nd shift and 3rd shift, and A, B, and C-lines, requires\nfurther study.\nRecommendations\nGenerally, this research identifies the causes for most problems related with the\nscrap rates that occur at the Phillips Plastics-Short Run facility based on injection\nmachines, processes, materials, and operators. Some findings could be refined and\nresearched further.\nThe 100 samples collected for use in this research only satisfy basic statistical\nanalysis requirements, but yield acceptable results for use in this research. For more\naccuracy, future research should be performed by collecting more data samples.\nFuture research should also involve gathering data about mold design. Bad mold\ndesign is another common cause of scrap generation. Including this data in the research\nanalysis may yield more comprehensive results and help better explain research findings.\nCollection of process time variables, such as fill time, pack/holding time, cooling\ntime, and cycle time, was outside the scope of this research. However, inconsistency in\nprocess time is a common cause of part defects. Including process time in future research\nanalysis can provide more comprehensive explanations in regards to scrap generation.\nSeveral production jobs required the implementation of secondary operations to\ncomplete part fabrication. These additional operations can lead to an increase in the\namount of part defects. For a more comprehensive study, research should also include\nmore specific information in regards to this data.\n54\nReferences\nBeaumont, J.P., Nagel, R., & Sherman, R. (2001). Successful Injection Molding. Hanser\nGardner Publications, Inc.\nBrydson, J. A. (1999). Plastics Materials. Woburn, MA: Butterworth-Heinemann.\nJohannaber, F. (1994). Injection Molding Machines. Cincinnati, Ohio: Hanser/Gardner\npublications, Inc.\nKolz, A. R., McFarland, L. A., & Silverman, S. B. (1998). Cognitive Ability and Job\nExperience as Predictors of Work Performance: The Journal of Psychology, 132(5),\n539-548\nQui\u00f1ones, M. A., Ford, J. K., & Teachout, M. S. (2001). The Relationship Between\nWork Experience and Job Performance: A conceptual and Meta-Analytic Review.\nPersonnel Psychology.\nReinhold, V. N. (1986). Injection Molding Handbook. New York, NY: Van Nostrand\nReinhold Company, Inc.\nReinhold, V. N. (1987). Injection Molds and Molding. New York, NY: Van Nostrand\nReinhold Company Inc.\nReinhold, V. N. (1991). Plastics Engineering Handbook. New York, NY: Van Nostrand\nReinhold.\nWhelan, A., & Craft, J. L. (1978). Development in Injection Molding _1. Barking, Essex,\nEngland: Applied Science Publishers Ltd.\nAppendix A\nScrap Reduction Research Datasheet\nJob #:\nCollection Date:\nPress #:\nItem #:\nProduction Order #:\nDates of Production:\nContamination\nPartials\nVoids/Bubbles\nSink\nSplay\nGood\nScrap Total\nShift\nInitials\nSupervisor's Signature:\nProduction Manager's Signature:\nMaterial:\nMelt Temps:\nMold Temps:\nPressure:\n55\n555\nAppendix B\nPress Investigation\nFor\nScrap Reduction Research\n56\n56\nPress #\nTonnage\nEstimated Age\nPress #\nTonnage\nEstimated Age\nA-1\nB-5\nA-3\nB-6\nA-4\nB-8\nA-5\nB-9\nA-6\nB-10\nA-7\nC-1\nA-8\nC-2\nA-9\nC-3\nA-10\nC-4\nA-11\nC-5\nA-12\nC-6\nA-13\nC-7\nA-14\nC-8\nA-15\nA-16\nC-9\nC-10\nB-1\nB-2\nB-3\nB-4\nSupervisor's Signature:\nProduction Manager's Signature:\nAppendix C\nInitials\nOperator Data\nFor\nScrap Reduction Research\nYear(s) of Experience (estimated)\nUnder 1 yr\n1-2 yrs Over 2 yrs\nSupervisor's Signature:\nProduction Manager's Signature:\n57\n"}, "expected_output": {"claims": [{"unit": "%", "value": 0.8, "evidence": ["In the range\nof melt temperatures under 500\u00b0F, the scrap level was 0.8%.", "0.8%"]}, {"unit": "%", "value": 3.9, "evidence": ["The scrap jumped to a\nhigher level of 3.9% with the melt temperatures between 500-600\u00b0F.", "3.9%"]}, {"unit": "%", "value": 4.2, "evidence": ["4.20"]}, {"unit": "%", "value": 4.6, "evidence": ["The over all scrap level related with the use of this material was\n29\n4.6%."]}, {"unit": "%", "value": 5.3, "evidence": ["Ultem was another material used often by Short Run. The study shows that the\nover all scrap level related with using this material was 5.3%. Most scrap problems\noccurred due to splay. The next problem occurred due to contamination scrap level.\nVoid/Bubble and burn/streak scrap levels were much lower than splay and contamination\nscrap levels. For this material, scrap levels due to sink marks and partial fill were very\nlow. Therefore, splay and contamination should be closely monitored when using Ultem.\n", "Ultem was another material used often by Short Run. The study shows that the over all scrap level related with using this material was 5.3%.", "5.3%"]}, {"unit": "percent", "value": 2.5, "evidence": ["Table 4 shows that the over all scrap level associated with using Cycolac was\n2.5%. The most common scrap problem presented with using this material proved to be\n", "Table 4 shows that the over all scrap level associated with using Cycolac was 2.5%.", "39,056 72 15 3 2 3 420 2.5%"]}, {"unit": "percent", "value": 2.9, "evidence": ["At higher mold temperatures (200\u00b0F and up) the\nscrap level was 2.9%."]}, {"unit": "percent", "value": 3, "evidence": ["Table 1 show that the over all scrap level associated with using this material was 3%.", "3.0%", "As Figure 5 shows, at the lower range of pressures (under 10,000 PSI), the scrap\nlevel was very high at 9.3%. At the mid-range of pressures (10,000-20,000 PSI), the\nscrap was lower at 3.7%. The process at higher pressures, 20,000 PSI or higher, resulted\nin a scrap level of 4.8%. In general, the scrap level was lowest at pressures of 10,000PSI\nto 20,000PSI. The processes that require higher pressures produce higher level of scrap.\nHowever, at a lower pressure (under 10,000 PSI) the scrap level also becomes very high.\nThe five most commonly used materials for this research were Calibre, Radel,\nUltem, Cycolac, and RTP. Calibre was the most commonly used material in the data\ncollected for this research. Table 1 show that the over all scrap level associated with\nusing this material was 3%. The contamination level related with this material resulted in\nthe highest specific scrap level. The scrap level due to burning, void/bubble, and splay\nwere also monitored. Scrap levels due to partial fill were very low using this material,\n"]}, {"unit": "percent", "value": 3.7, "evidence": ["At the mid-range of pressures (10,000-20,000 PSI), the scrap was lower at 3.7%.", "3.7%", "As Figure 5 shows, at the lower range of pressures (under 10,000 PSI), the scrap\nlevel was very high at 9.3%. At the mid-range of pressures (10,000-20,000 PSI), the\nscrap was lower at 3.7%. The process at higher pressures, 20,000 PSI or higher, resulted\nin a scrap level of 4.8%. In general, the scrap level was lowest at pressures of 10,000PSI\nto 20,000PSI. The processes that require higher pressures produce higher level of scrap.\nHowever, at a lower pressure (under 10,000 PSI) the scrap level also becomes very high.\nThe five most commonly used materials for this research were Calibre, Radel,\nUltem, Cycolac, and RTP. Calibre was the most commonly used material in the data\ncollected for this research. Table 1 show that the over all scrap level associated with\nusing this material was 3%. The contamination level related with this material resulted in\nthe highest specific scrap level. The scrap level due to burning, void/bubble, and splay\nwere also monitored. Scrap levels due to partial fill were very low using this material,\n"]}, {"unit": "percent", "value": 4.7, "evidence": ["The scrap percentage reduces to about half (at 4.7%) in the mid-range\nof mold temperatures (100 \u00b0F -200\u00b0F)."]}, {"unit": "percent", "value": 4.8, "evidence": ["The process at higher pressures, 20,000 PSI or higher, resulted in a scrap level of 4.8%.", "4.8%", "As Figure 5 shows, at the lower range of pressures (under 10,000 PSI), the scrap\nlevel was very high at 9.3%. At the mid-range of pressures (10,000-20,000 PSI), the\nscrap was lower at 3.7%. The process at higher pressures, 20,000 PSI or higher, resulted\nin a scrap level of 4.8%. In general, the scrap level was lowest at pressures of 10,000PSI\nto 20,000PSI. The processes that require higher pressures produce higher level of scrap.\nHowever, at a lower pressure (under 10,000 PSI) the scrap level also becomes very high.\nThe five most commonly used materials for this research were Calibre, Radel,\nUltem, Cycolac, and RTP. Calibre was the most commonly used material in the data\ncollected for this research. Table 1 show that the over all scrap level associated with\nusing this material was 3%. The contamination level related with this material resulted in\nthe highest specific scrap level. The scrap level due to burning, void/bubble, and splay\nwere also monitored. Scrap levels due to partial fill were very low using this material,\n"]}, {"unit": "percent", "value": 5.8, "evidence": ["5.80%"]}, {"unit": "percent", "value": 5.9, "evidence": ["5.90%"]}, {"unit": "percent", "value": 7, "evidence": ["7.00%"]}, {"unit": "percent", "value": 8.6, "evidence": ["At lower mold temperatures (under 100\u00b0F), the scrap level reaches a very\nhigh level of 8.6%."]}, {"unit": "percent", "value": 9.3, "evidence": ["The over all scrap level related with using RTP is presented in Table 5, and was found to be 9.3%.", "9.3%", "The over all scrap level related with using RTP is presented in Table 5, and was\nfound to be 9.3%. The table also shows the biggest problem was scrap due to\ncontamination. Sink marks and partial fill scrap levels should also be considered. Splay,\nbubble, and burning scrap levels were rarely problematic when using this material.\nTherefore, most scrap reduction efforts should be focused on contamination during the\n", "at the lower range of pressures (under 10,000 PSI), the scrap level was very high at 9.3%.", "As Figure 5 shows, at the lower range of pressures (under 10,000 PSI), the scrap\nlevel was very high at 9.3%. At the mid-range of pressures (10,000-20,000 PSI), the\nscrap was lower at 3.7%. The process at higher pressures, 20,000 PSI or higher, resulted\nin a scrap level of 4.8%. In general, the scrap level was lowest at pressures of 10,000PSI\nto 20,000PSI. The processes that require higher pressures produce higher level of scrap.\nHowever, at a lower pressure (under 10,000 PSI) the scrap level also becomes very high.\nThe five most commonly used materials for this research were Calibre, Radel,\nUltem, Cycolac, and RTP. Calibre was the most commonly used material in the data\ncollected for this research. Table 1 show that the over all scrap level associated with\nusing this material was 3%. The contamination level related with this material resulted in\nthe highest specific scrap level. The scrap level due to burning, void/bubble, and splay\nwere also monitored. Scrap levels due to partial fill were very low using this material,\n"]}, {"unit": "percent", "value": 34, "evidence": ["34.0%"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_551cddaddf7359d6"}}
{"id": "c7ff0fb7d1e1609a0682e871", "input": {"query": "What is the recycled content or secondary material percentage in the Bisley cabinet steel? Please provide numerical percentages.", "source_url": "https://cdn3.evostore.io/documents/bisley_ireland/s_p_01607_epd_bisley_two_door_cupboards.pdf", "document_text": "BISLEY\nENVIRONMENTAL\nPRODUCT DECLARATION\nIn accordance with ISO 14025\nDesign museum Gent\n2\nN\nOffice Storage | TWO-DOOR CUPBOARDS\nThe environmental impacts of this product\nhave been assessed from cradle to grave.\nThis Environmental Product Declaration has\nbeen verified by an independent third party.\nAn EPD should provide current information,\nand may be updated if conditions change.\nThe stated validity is therefore subject to\nthe continued registration and publication\nat www.environdec.com.\nDeclaration Number: S-P-01607\nIssued on: 2019-10-16\nValid until: 2024-08-28\nProgramme Operator: EPD International AB\nEPD\nTHE INTERNATIONAL EPD\u24c7 SYSTEM\nR\n1.\nINTRODUCTION\nThis EPD provides environmental performance indicators for storage furniture manufactured by FC Brown\nunder the brand name \"Bisley\u201d. This is a cradle-to-grave EPD, based on a life cycle assessment (LCA)\nstudy which used production data for 2018 from Bisley's manufacturing facilities in Newport, Wales, UK.\nBackground data were taken from the ecoinvent database (v3.4).\nThe EPD presents details of the LCA, a description of the product life cycle it covers, values for relevant\nenvironmental indicators and a brief explanation of those results.\nTwo-door Cupboards EPD: Lateralfile\u2122, Systemfile\u2122, Essentials and Be\u2122\nEPD programme\nEPD programme operator\nEPD owner\nProduct codes\nCPC code\nDeclared unit\nSystem boundaries\nReference year for data\nDeclaration no.\nDate of publication/valid until\nProcedure for data follow-up\nduring EPD validity\nEPD geographical scope\nEPD based on Product\nCategory Rules (PCR)\nPCR review conducted by\nVerification\nThird party Verifier\nApproved by\nLCA conducted by\nThe International EPD\u00ae System\nEPD International AB, Box 210 60, SE-100 31 Stockholm, Sweden\ninfo@environdec.com | www.environdec.com\nBisley, Caswell Way, Reevesland Industrial Estate, Newport, Gwent\nNP19 4PW, UK | www.bisley.com\nSee Appendix\nCPC 3812 under the UN CPC classification system v2.1\n1 item of storage furniture in use for 15 years\nCradle to grave\n2018\nS-P-01607\nIssued on 2019-10-16 / valid until 2024-08-28\nInvolves third party Verifier:\nWorldwide\nyes\nno\nPCR Furniture, except Seats and Mattresses - 2012:19, V2.0\n(valid 2019-06-17 to 2023-06-17)\nProduct Category Classification: UN CPC 3812/3813/3814\nThe Technical Committee of the International EPD\u00ae System Chair:\nMr Filippo Sessa, Quantis; contact via info@environdec.com\nIndependent verification of this EPD and data, according to ISO\n14025/2006: \u2610 internal certification\nUgo Pretato,\nStudio Fieschi & Soci S.r.l., Italy:\nexternal verification\nUplott\nThe International EPD\u00ae System Technical Committee,\nsupported by the Secretariat\nEuGeos Limited, UK - +44 (0)1625 434423\nwww.eugeos.co.uk\nEuGeos\nThe EPD owner has the sole ownership, liability and responsibility of the EPD. EPDS within the same\nproduct category but from different programmes may not be comparable.\n2 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nCOMPANY PROFILE\nSince launching the iconic MultiDrawer in 1958, Bisley has pioneered innovative storage that meets\nthe challenges of the time and stays relevant, decade after decade. When it comes to creating working\nenvironments, Bisley is the name people in over 50 countries turn to for quality they can trust.\nBisley produces 15,000 items each week from its factory in Wales and makes over 4000 world-wide\ndeliveries every month via its 10 international offices. Through continued production investment and\nacquisition, Bisley can combine the strength and durability of steel with the pleasing aesthetic qualities of\nwood to provide a unique choice of solutions for an extensive range of markets.\nBisley products maximise workspaces by providing intelligent and often bespoke storage solutions which\ncreate organised environments. Whether that means furniture for focused and individual working or\nadaptable pieces for collaborative spaces, meeting rooms or break-out areas, the extensive portfolio helps\npeople across the world be comfortable and effective, wherever they are working.\nThe product is manufactured at Bisley's facility in Newport, Gwent, Wales.\nContact:\nBisley\nCaswell Way\nReevesland Industrial. Estate\nNewport\nGwent NP19 4PW\nUnited Kingdom\nT. 01633 637383\ninfo@bisley.com\nwww.bisley.com\n3 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nPRODUCT INFORMATION\nThis EPD applies to Bisley two-door cupboards in the LateralFileTM, System File\u2122, Essentials and Be\u2122\nranges.\nThe cupboards are available in 3 widths and up to 12 heights; every unit comes with levelling feet, double\nskin doors and soft close functionality as standard. They are made versatile by the large variety of internal\noptions available, including shelving, from fitted, roll-out and lateral rails, to pigeon hole sets and coat rails.\nThe ventilated door option is ideal for coat storage. All optional elements are to be ordered separately.\nAll Bisley two-door cupboards are classified CPC 3812 under the UN CPC classification system v2.1.\nIn this EPD, environmental performance is declared for the smallest, a medium-sized and the largest two-\ndoor cupboards in each of the 4 ranges, without optional accessories.\nFIGURE 1. LATERALFILE\u2122M FIGURE 2. SYSTEMFILETM\nFIGURE 3. ESSENTIALS\nFIGURE 4. BETM\nBISLEY\nTWO-DOOR\nCUPBOARDS\nRANGE\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nUNIT\nUNIT\nUNIT\nCODE\nCODE\nCODE\nCONFIGURATION\nCONFIGURATION\nCONFIGURATION\nLATERALFILETM\n08D20\nno internal fittings\n09D36\nno internal fittings\n10D64\nno internal fittings\nSYSTEMFILETM\nSYD08/2\nno internal fittings\nSYD10/4\nno internal fittings\nSYD10/6\nno internal fittings\nESSENTIALS\nYECB0807\n/1S\n1 shelf\nYECB1010\n1 shelf\n/1S\nYECB1019\n/4S\n4 shelves\nBETM\nBH08071\n004NP\nno internal fittings\nBH081411\n58NP\nno internal fittings\nBH101813\n18NP\nno internal fittings\nManufacturing\nSteel in sheet or coil form is cut to size, and formed using standard steelworking techniques\nof bending, pressing, stamping, welding to produce the main parts of Bisley products. These\nare coated with polymer-based, solvent-free epoxy-polyester powder coatings. Main parts are\ncombined, and small parts (for example handles, locks, adjustable feet) purchased from external\nsuppliers added, in final product assembly. Products are then packed prior to despatch.\n4| Environmental Product Declaration: Two-door Cupboards\nBISLEY\nPackaging\nEach unit is wrapped in plastic film for protection, with cardboard reinforcement at key points (e.g.\ncorners); this packaging remains in place until the product reaches the point of use. Packed products are\npalletised for transport to customers.\nTransportation\nBisley products are sent to Bisley's own distribution subsidiaries, or directly to large customers and major\nprojects, by road or by road and sea.\nProduct Use and Maintenance\nAll Bisley products carry a 10-year warranty. They require no energy or water inputs to function. Bisley\nguidance to customers is that products should be cleaned periodically with mild detergent and warm\nwater.\nEnd-of-life\nWhen the user has no further use for Bisley office furniture products, they may be reused by others,\nrecycled or disposed of as non-hazardous waste. Reuse is recommended, but if no route for reuse is\navailable, the product - which is more than 95% steel - should be recycled with other ferrous-metal goods.\nFurther Product Information\nDetailed product information and datasheets can be found on our website www.bisley.com or by\ncontacting Customer Service by calling 01633 637383 or emailing info@bisley.com.\nCONTENT DECLARATION\nThe material composition of Bisley's storage furniture products covered by this EPD is shown below:\nMATERIAL\n% OF MASS PER FUNCTIONAL UNIT\nLATERALFILE\u2122\nSYSTEMFILE\u2122\nESSENTIALS\nBETM\nsteel\n98-99\n>99\n>99\n95-96\nepoxy-polyester resin\n<0.1\n<0.1\n<0.1\n<0.1\nother metals: aluminium,\nnickel, zinc\n<0.5\n<0.2\n<0.2\n<0.2\nother polymers\n1-2\n<0.1\n<0.1\n4\nSteel used by Bisley has a recycled content typical of European steel production, quoted as 56% by steel\nsuppliers in 2018. This is assumed to be 32% pre-consumer, 24% post-consumer scrap based on the\ninformation in the background LCA database.\nProducts are shrink-wrapped in plastic film to protect them until they reach the user; cardboard\nreinforcement is used on the corners of units and other areas particularly susceptible to damage. Cardboard\nis assumed to have a recycled content of approx. 75% based on the information in the background LCA\ndatabase. Distribution packaging varies according to the destination, but products are always distributed on\nwooden pallets; approximately 20% of all pallets used by Bisley in 2018 had been used at least once before.\nNo substance included in the Candidate List of Substances of Very High Concern for authorisation under\nthe REACH Regulations is present in the furniture, either above the threshold for registration with the\nEuropean Chemicals Agency or above 0.1% (wt/wt).\n5 Environmental Product Declaration: Two-door Cupboards\nBISLEY\nTECHNICAL DATA\nBisley two-door cupboards are one element in a systems approach to storage which can evolve with your\nchanging needs.\nKey technical properties and certifications are shown in the table below; consult the relevant product\nTechnical Data Sheet for a comprehensive specification.\nTECHNICAL PROPERTIES (TESTS)\nBS EN 14073-2:2004 Office furniture. Storage furniture.\nSafety requirements\nVALUE\nALL SIZES\nUNIT\nPass\nN/A\nBS EN 14073-3:2004 Office furniture. Storage furniture. Test\nmethods for the determination of stability and strength of the\nstructure\nPass\nN/A\nBS EN 14074:2004 Office furniture. Tables and desks and\nstorage furniture. Test methods for the determination of\nstrength and durability of moving parts\nPass\nN/A\nANSI/BIFMA M7.1/X7.1-2011(R2016) Test method /\nacceptance criteria for VOC emissions from furniture used in\noffices to be classified as low-emitting product\nPass\nN/A\nVolatile organic chemical emission testing method for\nCalifornia Department of Public Health specification 01350\nPass\nN/A\nPHYSICAL DATA\n2-DOOR CUPBOARDS - LATERALFILE\u2122 RANGE\nUNIT\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nDimensions (height, length,\ndepth)\nmm\n693 x 800 x 470\n1225 x 900 x 470\n2099 x 1000 x 470\nVolume\nMass (approx.)\nStorage units according to\nBIFMA storage PCR\n(1 unit = 0.15m\u00b3)\nm\u00b3\n0.26\n0.52\n0.99\nkg\n27\n45\n75\nnumber\n1.74\n3.45\n6.58\n6 Environmental Product Declaration: Two-door Cupboards\nBISLEY\nPHYSICAL DATA\n2-DOOR CUPBOARDS - SYSTEMFILE\u2122 RANGE\nUNIT\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nDimensions (height, length,\nmm\n693 x 800 x 470\n1301 x 1000 x 470\n1947 x 1000 x 470\ndepth)\nVolume\n0.26\nMass (approx.)\nm\u00b3\nkg\n25\n44\n61\n0.61\n0.92\nStorage units according to\nBIFMA storage PCR\n(1 unit\n0.15m\u00b3)\nDimensions (height, length,\nnumber\n1.74\n4.08\n2-DOOR CUPBOARDS - ESSENTIALS RANGE\n6.10\nUNIT\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nmm\n693 x 800 x 470\n1000 x 1000 x 470\n1947 x 1000 x 470\ndepth)\nVolume\nm\u00b3\n0.26\n0.47\n0.92\nMass (approx.)\nkg\n26\n30\n48\nStorage units according to\nBIFMA storage PCR\nnumber\n1.74\n3.13\n6.10\n(1 unit = 0.15m\u00b3)\nDimensions (height, length,\n2-DOOR CUPBOARDS - BET RANGE\nUNIT\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nmm\n734 x 800 x 440\n1494 x 800 x 440\n1874 x 800 x 440\ndepth)\nVolume\nm\u00b3\n0.26\n0.52\n0.66\nMass (approx.)\nkg\n35\n49\n61\nStorage units according to\nBIFMA storage PCR\nnumber\n1.72\n3.51\n4.40\n(1 unit\n0.15m\u00b3)\nResidual Risks and Emergencies\nThere are no residual risks associated with the normal day-to-day use of Bisley's storage furniture.\nCare must be taken to follow the guidance for safe use in the product information documents for Bisley's\ntwo-door cupboards available from www.bisley.com/resources/product-information/.\n7 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nENVIRONMENTAL PERFORMANCE RELATED INFORMATION\nLCA INFORMATION\nThis section of the EPD records key features of the LCA on which it is based.\nScope\nThis cradle-to-grave EPD is applicable globally; end-of-life scenarios are based on European statistics for\nwaste management. For the presentation of results, and reflecting the different sources of data used, the\nlife cycle of products is divided into three different stages:\n\u2022\nUpstream processes (from cradle-to-gate)\n\u2022 Core processes (from gate-to-gate)\n\u2022\nDownstream processes (from gate-to-grave)\nSystem Boundaries\nThe system boundary of the EPD is defined using a modular approach reflecting the three life-cycle stages.\nStorage furniture is used in buildings, therefore the equivalence of the modules covered by this EPD to the\nmodules defined in EN 15804 is also presented (see table below).\nThose modules included in the LCA are denoted by v; those not declared by ND: those not relevant by NR.\nUPSTREAM PROCESSES\nRaw material extraction &\nproduction\nTransport\nManufacturing (main parts)\nElectricity and fuel use\nManufacturing (auxiliary products,\npackaging)\nWaste treatment\nCORE PROCESSES\nTransport\nManufacturing/Assembly\nMaintenance (equipment)\nWaste treatment\nElectricity and fuel use\nDistribution transport\nDOWNSTREAM PROCESSES\nProduct end-of-life\nvvvvvvvvvvvvvvv\nEN 15804 MODULES\nA1\nA1\nA1\nA1\nA1 A1\nA2 A3 A3\nA3\n13\n8 | Environmental Product Declaration: Two-door Cupboards\nB1 v\nUse\nB2 v\nC1-\nA3 A3\nA4\nB4 v\nA5\nC4\nB3 NR\nB5-B7 NR\nBISLEY\nPackaging end-of-life\nThe product life cycle covered by this EPD is illustrated below.\nMaterial recycling\nBasic material production, transportation,\ntransformation\nSteel &\nPolymers\nSmall\nparts\nCoating\nmaterials\nAncillary\nmaterials\nTransport to factory\nCarcass production\nCoating\nFuel extraction\nand energy\nproduction\nUPSTREAM\nFuel extraction\nand energy\nproduction\nAssembly\nPacking\nCORE\nDelivery to customer\nUse\nProduct removal after use\nRe-use\nRecycling\nLandfill\n14%\n85%\nPRODUCT LIFE CYCLE (CRADLE-TO-GRAVE)\n9| Environmental Product Declaration: Two-door Cupboards\nPackaging\nwaste management\nDOWNSTREAM\n1%\nSYSTEM BOUNDARY\nBISLEY\nFunctional Unit\nThe functional unit is 1 item of storage furniture in use for 15 years. Bisley products carry a 10-year\nwarranty; this is the minimum service life that can be expected. In this LCA, the default product lifetime\nspecified in the PCR - 15 years - is applied. One product unit therefore provides the functional unit.\nCut-off Criteria\nThe collected data covered all raw materials, consumables and packaging materials; associated transport to\nthe manufacturing site; process energy and water use; direct production wastes; emissions to air and water.\nAccording to the PCR, flows can be omitted (cut off) from a core process in the LCA up to a maximum\nof 1% of the total mass of material inputs or 1% of the total energy content of fuels and energy carriers;\nsmall components such as plastic washers, small screws and adjustable feet amounting, in combination,\nto <1% of total input materials were omitted from the LCA underpinning this EPD. This is consistent with\nrequirements of the PCR and General Programme Instructions.\nData Sources and Data Quality\nData characterising the core processes (cabinet manufacture, coating, assembly and packing) were\ncollected for the calendar year 2018. Therefore the producer-specific data used in LCA calculations are\nbased on 1 year averaged data, which matches the requirement of EN 15804. The data have been updated\nwithin the last 5 years. These data were checked to ensure that sufficient materials and water are included\nwithin the inputs to account for all products, wastes and emissions.\nBACKGROUND DATA:\nBackground (generic) data were taken from the ecoinvent database (v3.4); thus generic data used in the\nLCA have been updated within the last 10 years.\nData quality has been reviewed for processes that contribute significantly to the overall LCA. Other data\nwere judged fit for purpose. No environmental impact potential stemming from proxy data exceeds 10% for\nany impact category.\nBisley purchase electricity on a low-carbon tariff; the fuel mix notified by the supplier was used to model\nelectricity supply to the Newport factory.\nAllocation\nIn the background data, the ecoinvent default allocation is applied to all processes except those in which\nsecondary materials are used, where the \"cut-off\u201d allocation is applied. This ensures that secondary\nmaterials are free of upstream burdens that arise prior to their reaching the \"end of waste\u201d state; this is in\naccordance with the PCR and also Section 6.3.4.2 of EN 15804.\nFollowing ISO 14044, the overall process is subdivided as far as possible, so that flows dedicated to a\nparticular product type are fully assigned to that product type and the need for allocation is minimised.\nUtility, packaging and ancillary material inputs to the production facility have been allocated across all\nproducts manufactured at the facility, including those not covered by this EPD.\nAssumptions and Estimates\nInputs to and outputs from the system are accounted for over a 100-year time period; long-term emissions\nare therefore omitted from the impact assessment part of the LCA.\nThe \"primary energy used as material\u201d indicators (PERM; PENRM) are calculated using - as characterisation\nfactors-published values for constituent materials which can yield energy on combustion, where available,\nand from published calorific values where PEM values are not available. Calculations of PE(N)RM are based\non a feedstock energy content of 27MJ/kg for pvc (where present), 52MJ/kg for other polymers, 43MJ/kg\nfor coating materials, 16MJ/kg for wood and 14MJ/kg for cardboard.\n\"Primary energy as fuel\" indicators (PENRE, PERE) are calculated as the total primary energy demand\nminus primary energy used as material.\n10 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\n\n\nThe secondary material indicator counts scrap steel, recycled polymer and other recycled material inputs to\nthe product and its constituent components, re-used wooden packaging and recycled paper/board inputs\nto packaging manufacture.\nSCENARIOS:\nTransport to the customer, product maintenance, transport to waste management of packaging and\nproduct at end-of-life, and management of end-of-life product and waste packaging are characterised using\nscenarios.\nUnder normal use conditions, no replacement parts or maintenance are required during the warranty\nperiod of 10 years. Product maintenance, which comprises cleaning according to the manufacturer's\ninstructions, is assumed to consume 1l water and 5g detergent per year. No other inputs or outputs are\nrequired for use of the product. The default product lifetime specified in the PCR is applied, so that one\nproduct fulfils the functional unit.\nThe effective mass per unit volume of the product is obtained by dividing the reported mass by volume,\napplying the values in the \u201cTechnical Data\" section above. Other relevant parameters for transport are\nshown in the table below:\nSCENARIO PARAMETERS, TRANSPORT\nPARAMETER AND UNIT\nVALUE & UNIT\nTO CUSTOMER\nTO END-OF-LIFE MANAGEMENT\nFuel type and consumption\nroad: diesel - 0.21/km\nsea: fuel oil - 2.5 g/tkm\ndiesel: 0.2 l/km\nDistance\nroad: 560\nsea: 890\nCapacity utilisation (including\nempty returns) %\n58\nVolume capacity utilisation\nfactor\n1\nroad: 50\n28\n1\nThe proportions of each waste management method assumed for packaging materials and for products\n(cabinets) are shown in the table below, based on data for Bisley's major European markets (source:\neurostat).\nMATERIAL/PRODUCT\nRE-USE\nRECYCLING\nPERCENTAGE TO:\nINCINERATION\nWITH ENERGY\nRECOVERY\nLANDFILL\n9\n6\nCardboard packaging\n0\n85\nPlastic packaging\n0\n42\n34\n24\nWood (pallets)\n0\n47\n27\n26\nCabinet\n14\n85\n0\n1\n11 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nENVIRONMENTAL INDICATORS AND INTERPRETATION\nThis EPD contains environmental information about the specified products, in the form of quantitative\nindicator values for a number of parameters, which encompass calculated environmental impact potentials,\nresource and energy use, and waste generation.\nThe parameters are listed below along with the abbreviations used for them in the tables of indicator values\nthat follow.\nPARAMETER\nENVIRONMENTAL IMPACTS\nClimate change - GWP100\nABBREVIATION\nUNITS\nGWP\nkg CO2 eq\n(fossil, biogenic, land use and transformation)\nAcidification potential - fate excluded\nAP (fx)\nkg SO2 eq\nEutrophication - generic\nEP\nkg PO4\u00b3 eq\nPhotochemical oxidant creation potential\nPOCP\nKg NMVOC eq\nDepletion of abiotic resources - elements, ultimate reserves\nADPE\nKg Sb eq\nDepletion of abiotic resources - fossil fuels\nADPFF\nMJ\nWater scarcity potential\nWSP\nm\u00b3 eq\nRESOURCE USE\nRenewable primary energy as energy carrier\nPERE\nMJ\nRenewable primary energy resources as material utilisation\nPERM\nTotal renewable primary energy use\nPERT\n(sum of the two parameters above)\nNon-renewable primary energy as energy carrier\nPENRE\nNon-renewable primary energy resources as material\nutilisation\nPENRM\n\u03a3\u03a3\u03a3\u03a3\u03a3\nMJ\nMJ\nMJ\nMJ\nTotal non-renewable primary energy use\nPENRT\nMJ\n(sum of the two parameters above)\nUse of secondary material\nSM\nkg\nUse of renewable secondary fuels\nRSF\nUse of non-renewable secondary fuels\nNRSF\n\u03a3\u03a3\nMJ\nMJ\nNet use of fresh water\nFW\nm\u00b3\n12 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nPARAMETER\nWASTES\nHazardous waste disposed\nNon-hazardous waste disposed\nRadioactive waste disposed\nOUTPUT FLOWS\nABBREVIATION\nUNITS\nHWD\nkg\nNHWD\nkg\nTRWD\nkg\nComponents for re-use\nCFR\nkg\nMaterials for recycling\nMFR\nkg\nMaterials for energy recovery\nMER\nExported energy - electricity\nEEE\nExported energy - thermal\nEET\n\u03a9\u03a3 \u03a3\nkg\nMJ\nMJ\nOTHER ENVIRONMENTAL INDICATORS\nHuman toxicity - cancer impacts\nHTC\ncases\nHuman toxicity - non-cancer impacts\nHTNC\ncases\nFresh water ecotoxicity\nFWE\nkgPAF.m3.day\nLand use\nLU\nspecies.yr\nAcidification potential, fate included - average Europe*\nAP\nkg SO2 eq\nOzone layer depletion - ODP steady state*\nODP\nkg CFC-11 eq\n* Additional indicator specified by EN15804, for information\nEnvironmental indicator results are shown in the following tables for the declared unit of 1 item of storage\nfurniture (two-door cupboard) in use for 15 years, in Bisley's Lateral FileTM, System FileTM, Essentials and Be\u2122\nranges.\n13 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nMEDIUM UNIT\nTwo-door Cupboards - LateralFile\u2122\nEnvironmental indicator results are shown in the 5 following tables for the declared unit of 1 item of storage furniture in use for 15 years.\nTWO-DOOR CUPBOARDS - LATERALFILE\u2122 RANGE\nSMALLEST UNIT\n\u2610\nLARGEST UNIT\n11\n!!!!!!!!!!!\nIIIIIIIIII\nDOWN\nPROCESSES\n7.67E+00\n1.55E+00\n1.20E+01\n9.25E+00\n9.88E-03\n4.73E-03\n4.39E-03\n3.57E-03\n2.51E-02\n2.02E-02\n2.34E-05\n5.22E-05\n1.99E+02\n8.44E+01\n3.56E+00\n2.76E+00\n!!!!!!!!!!!\nIIIIIII\nENVIRONMENTAL\nUNIT\nIMPACT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nCORE\nPROCESSES\n7.30E+01\n1.06E+01\n4.91E+00\n1.18E+01\nfossil\nbiogenic\n2.21E+00\n1.53E-01\n1.53E+00\n1.89E-01\nGWP\nkg CO\u2082 eq\nland*\n1.10E-01\n2.45E-03\n4.21E-02\n3.38E-03\n4.27E-02\n7.53E+01\n1.08E+01\n6.48E+00\ntotal\nAP(fx)\nkg SO2 eq\n1.06E-01\n6.43E-03\n3.23E-03\nEP\nkg PO43-\n3.67E-02\n3.41E-03\n2.62E-03\neq\nKg NMVOC\nPOCP\n3.20E-01\n1.87E-02\n1.29E-02\neq\nADPE\nKg Sb eq\n5.50E-04\n1.65E-05\n3.30E-05\nADPFF\nMJ\n1.08E+03\n1.82E+02\n5.36E+01\nWSP\nm\u00b3 eq\n4.49E+01\n3.25E+00\n2.50E+00\n14 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\n15 | Environmental Product Declaration: Two-door Cupboards\nRESOURCE\nUNIT\nUSE\nPERE\nMJ\nPERM\nMJ\nPERT\nMJ\nPENRE\nMJ\nPENRM\nMJ\nPENRT\nBISLEY\nMJ\nSM\nkg\nRSF\nHlllll!!!!!\nTWO-DOOR CUPBOARDS - LATERALFILE\u2122 RANGE\nLARGEST UNIT\nMEDIUM UNIT\nSMALLEST UNIT\nHIIIIIIIIII\n////////\n{1111111111\nIIIIIIII\n!!!!!!!!!!\n||||||||111111\n!!!!!!!!!!!\nIll\n/////////ll\nMJ\nNRSF\nMJ\nFW\nm3\n16 | Environmental Product Declaration: Two-door Cupboards\nTWO-DOOR CUPBOARDS - LATERALFILE\u2122 RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nWASTES\nUNIT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nHWD\nkg\n5.35E-02\n2.24E-03\n1.70E-03\n5.74E-02\n8.13E-02\n3.13E-03\n2.68E-03\n8.71E-02\n1.52E-01\n4.63E-03\n4.31E-03\nNHWD\nkg\n5.23E+00\n1.01E+00\n4.09E+00\n1.03E+01\n8.33E+00\n1.45E+00\n5.92E+00\n1.57E+01\n1.49E+01\n2.18E+00\n8.95E+00\nTRWD\nkg\n2.72E-03\n3.44E-03\n3.40E-04\n6.50E-03\n4.36E-03\n3.56E-03\n5.60E-04\n8.48E-03\n7.88E-03\n3.77E-03\n9.20E-04\n1.26E-02\nBISLEY\nTWO-DOOR CUPBOARDS - LATERALFILE\u2122M RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nOUTPUT FLOWS\nUNIT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nTOTAL\nUP\nCORE\nPROCESSES PROCESSES\nCFR\nkg\n0.00E+00\n0.00E+00\n3.84E+00\n3.84E+00\n0.00E+00\n0.00E+00\nMFR\nkg\n7.93E-02\n2.39E-03\n2.34E+01\n2.35E+01\n8.07E-02\n2.40E-03\nMER\nkg\n6.10E-11\n1.69E-12\n3.53E-09\n3.60E-09\n7.03E-11\n2.14E-12\nTOTAL\n6.33E+00\n3.82E+01\n!!!!!!\nLARGEST UNIT\nJIIIII\nVIIIII\n3.61E-09\nEEE\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\nEET\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\nTOTAL\n1.04E+01\n6.30E+01\n3.66E-09\n0.00E+00\n0.00E+00\n\u2610 11\nTWO-DOOR CUPBOARDS - LATERALFILE\u2122 RANGE\nLARGEST UNIT\nMEDIUM UNIT\nSMALLEST UNIT\nOTHER\nUNIT\nJI I II II\n( ! ! ! ! ! !\n!!!!!!!\n7!!!!!!\n!!!!\n[1111]\nENVIRONMENTAL\nINDICATORS\nTOTAL\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nI\nHTC\ncases\n5.49E-05\n3.31E-07\n6.96E-07\n5.59E-05\n1.\nHTNC\ncases\n5.45E-05\n1.12E-06\n1.00E-06\n5.67E-05\nFWE\nkgPAF.\nm3.day\n4.53E+05\n2.66E+04\n7.23E+03\n4.87E+05\nLU\nspecies.yr\n1.90E-08\n9.74E-10\n2.84E-09\n2.28E-08\nAP\nkg SO2 eq\n3.42E-01\n1.63E-02\n1.49E-02\n3.73E-01\nODP\nkg CFC-11\neq\n4.95E-06\n3.20E-06\n5.99E-07\n8.75E-06\n17 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nTwo-door Cupboards - SystemFile\u2122\nEnvironmental indicator results are shown in the 5 following tables for the declared unit of 1 item of storage furniture in use for 15 years.\n18 | Environmental Product Declaration: Two-door Cupboards\nTWO-DOOR CUPBOARDS - SYSTEMFILE\u2122 RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nENVIRONMENTAL\nUNIT\nIMPACT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nCORE\nPROCESSES\nDOWN\nPROCESSES\nfossil\n6.35E+01\n1.05E+01\n4.56E+00\n1.18E+01\n7.47E+00\nbiogenic\n1.99E+00\n1.49E-01\n1.53E+00\n1.88E-01\n1.54E+00\nGWP\nkg CO\u2082 eq\nland*\n1.03E-01\n2.34E-03\n4.20E-02\n3.35E-03\n4.27E-02\nBISLEY\ntotal\n6.55E+01\n1.06E+01\n6.13E+00\nAP(fx)\nkg SO2 eq\n9.64E-02\n6.04E-03\n3.04E-03\nEP\nkg PO43-\n3.19E-02\n3.30E-03\n2.50E-03\neq\nPOCP\nKg NMVOC\n2.85E-01\n1.80E-02\n1.20E-02\neq\nLARGEST UNIT\n\u2610\n[!!!!!!!!\nTlllll\n!!!!!!!!!!I\n||||||||||||\n1.20E+01\n9.06E+00\n9.77E-03\n!!!!!!!1111\n4.36E-03\n2.49E-02\n4.63E-03\n3.51E-03\n1.97E-02\nADPE\nKg Sb eq\n5.30E-04\n1.57E-05\n3.06E-05\n2.31E-05\n5.08E-05\nADPFF\nMJ\n9.29E+02\n1.80E+02\n4.97E+01\n1.99E+02\n8.23E+01\n3.55E+00\n2.74E+00\nWSP\nm\u00b3 eq\n4.01E+01\n3.22E+00\n2.47E+00\n19 | Environmental Product Declaration: Two-door Cupboards\nRESOURCE\nUNIT\nUSE\nUP\nPROCESSES\n\u2610 cllll!!!!!!\nTWO-DOOR CUPBOARDS - SYSTEMFILE\u2122 RANGE\nLARGEST UNIT\nMEDIUM UNIT\nSMALLEST UNIT\n+IIIIIIIII\n///////// 111\n!!!!!!!1111\nIII\n!!!!!!!!!!!\n[!!!!!!!\n[!!!!!!!lII\nIl\nPERE\nMJ\n4.94E+01\nPERM\nMJ\n1.50E+01\nPERT\nMJ\n6.44E+01\nPENRE\nMJ\n1.02E+03\nPENRM\nMJ\n1.38E+01\nPENRT\nMJ\n1.04E+03\nBISLEY\nSM\nkg\n1.79E+01\nRSF\nMJ\n0.00E+00\nNRSF\nMJ\n0.00E+00\nFW\nm3\n9.38E-01\n20 | Environmental Product Declaration: Two-door Cupboards\nTWO-DOOR CUPBOARDS - SYSTEMFILE\u2122M RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nWASTES\nUNIT\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nHWD\nkg\n5.25E-02\n2.13E-03\n1.57E-03\n5.62E-02\n8.15E-02\n3.10E-03\n2.61E-03\n8.72E-02\n1.07E-01\n3.96E-03\n3.54E-03\nBISLEY\nNHWD\nkg\n4.68E+00\n9.60E-01\n3.85E+00\n9.50E+00\n7.96E+00\n1.43E+00\n5.79E+00\n1.52E+01\n1.09E+01\n1.85E+00\n7.51E+00\nTRWD\nkg\n2.37E-03\n3.42E-03\n3.20E-04\n6.11E-03\n4.10E-03\n3.56E-03\n5.50E-04\n8.20E-03\n5.64E-03\n3.68E-03\n7.50E-04\n1.01E-02\nTWO-DOOR CUPBOARDS - SYSTEMFILE\u2122M RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nOUTPUT FLOWS\nUNIT\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nPROCESSES\nTOTAL\nCFR\nkg\n0.00E+00\n0.00E+00\n3.53E+00\n3.53E+00\n0.00E+00\n0.00E+00\n6.15E+00\nMFR\nkg\n7.88E-02\n2.39E-03\n2.19E+01\n2.20E+01\n8.02E-02\n2.40E-03\n3.79E+01\nMER\nkg\n9.46E-11\n1.64E-12\n3.53E-09\n3.63E-09\n1.03E-10\n2.12E-12\n3.64E-09\nLARGEST UNIT\n$!!!!!\nEEE\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\nTOTAL\n8.49E+00\n5.21E+01\n3.65E-09\n0.00E+00\nEET\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n\n\n\u2610 llllll\n!!!!!\nTWO-DOOR CUPBOARDS - SYSTEMFILE\u2122 RANGE\nLARGEST UNIT\nMEDIUM UNIT\nSMALLEST UNIT\nOTHER\nUNIT\n! ! ! ! ! ! !\n!!!!!!!\n!!!!!!\n7IIIIII\n!!!!!!\nENVIRONMENTAL\nINDICATORS\nDOWN\nTOTAL\nUP\nCORE\nPROCESSES PROCESSES PROCESSES\nHTC\ncases\n5.09E-05\n3.18E-07\n6.42E-07\n5.19E-05\nHTNC\ncases\n5.00E-05\n1.10E-06\n9.41E-07\n5.20E-05\nFWE\nkgPAF.\nm3.day\n3.86E+05\n2.61E+04\n6.83E+03\n4.19E+05\nLU\nspecies.yr\n1.74E-08\n9.15E-10\n2.74E-09\n2.11E-08\nAP\nkg SO2 eq\n2.75E-01\n1.56E-02\n1.38E-02\n3.05E-01\nODP\nkg CFC-11\neq\n4.41E-06\n3.18E-06\n5.52E-07\n8.14E-06\n21 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\n22 | Environmental Product Declaration: Two-door Cupboards\nTwo-door Cupboards - Essentials\nEnvironmental indicator results are shown in the 5 following tables for the declared unit of 1 item of storage furniture in use for 15 years.\nTWO-DOOR CUPBOARDS - ESSENTIALS RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nENVIRONMENTAL\nUNIT\nIMPACT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nCORE\nPROCESSES\nBISLEY\nfossil\n6.57E+01\n1.06E+01\n4.72E+00\nbiogenic\n2.04E+00\n1.52E-01\n1.53E+00\nGWP\nkg CO\u2082 eq\nland*\n1.01E-01\n2.43E-03\n4.20E-02\ntotal\n6.78E+01\n1.08E+01\n6.29E+00\n!!!!! IIIII\nAP(fx)\nkg SO\u2082 eq\n1.00E-01\n6.35E-03\n3.12E-03\nEP\nkg PO43-\neq\n3.23E-02\n3.39E-03\n2.55E-03\nKg NMVOC\nPOCP\n2.96E-01\n1.86E-02\n1.24E-02\neq\nADPE\nKg Sb eq\n2.00E-04\n1.63E-05\n3.17E-05\nADPFF\nMJ\n9.62E+02\n1.81E+02\n5.15E+01\nWSP\nm\u00b3 eq\n4.16E+01\n3.24E+00\n2.48E+00\nDOWN\nPROCESSES\n1.08E+01\n5.32E+00\n1.60E-01\n2.62E-03\n1.53E+00\n\u2610 lllllllll\nLARGEST UNIT\n|||||||||||||\n||||||||||||||||\nIII\nTIIIII\n4.22E-02\n1.10E+01\n6.90E+00\n7.06E-03\n3.45E-03\n3.59E-03\n2.76E-03\n1.99E-02\n1.40E-02\n1.77E-05\n3.59E-05\n1.85E+02\n5.82E+01\n3.31E+00\n2.54E+00\n23 | Environmental Product Declaration: Two-door Cupboards\nRESOURCE\nUNIT\nUSE\nUP\nPROCESSES\n\u2610 !!\nTWO-DOOR CUPBOARDS - ESSENTIALS RANGE\nLARGEST UNIT\nMEDIUM UNIT\nSMALLEST UNIT\nIIIIIIIIIII\n[!!!!!!!!!!\n|||||||||||||||||\n1111111111\n!!!!!!!!!!\n!!!!!!!!111\nIll\n1111111111\nPERE\nMJ\n5.08E+01\nPERM\nMJ\n1.50E+01\nPERT\nMJ\n6.58E+01\nPENRE\nMJ\n1.06E+03\nPENRM\nMJ\n1.33E+01\nBISLEY\nPENRT\nMJ\n1.08E+03\nSM\nkg\n1.87E+01\nRSF\nMJ\nNRSF\nMJ\n0.00E+00\n0.00E+00\nFW\nm3\n9.73E-01\n24 Environmental Product Declaration: Two-door Cupboards\nTWO-DOOR CUPBOARDS - ESSENTIALS RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nWASTES\nUNIT\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nHWD\nkg\n4.21E-02\n2.21E-03\n1.63E-03\n4.60E-02\n5.49E-02\n2.40E-03\n1.85E-03\n5.91E-02\n8.32E-02\n3.34E-03\n2.86E-03\nNHWD\nkg\n4.82E+00\n9.99E-01\n3.96E+00\n9.78E+00\n5.53E+00\n1.09E+00\n4.36E+00\n1.10E+01\n8.72E+00\n1.55E+00\n6.24E+00\nTRWD\nkg\n2.47E-03\n3.43E-03\n3.30E-04\n6.24E-03\n2.83E-03\n3.46E-03\n3.80E-04\n6.66E-03\n4.51E-03\n3.59E-03\n6.00E-04\n8.70E-03\nBISLEY\nTWO-DOOR CUPBOARDS - ESSENTIALS RANGE\nLIGHTEST\nMEDIUM\nOUTPUT FLOWS\nUNIT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nTOTAL\nUP\nCORE\nPROCESSES PROCESSES\nTOTAL\nCFR\nkg\n0.00E+00\n0.00E+00\n3.67E+00\n3.67E+00\n0.00E+00\n0.00E+00\nMFR\nkg\n7.88E-02\n2.39E-03\n2.28E+01\n2.29E+01\n7.92E-02\n2.39E-03\nMER\nkg\n5.83E-11\n1.68E-12\n3.53E-09\n3.59E-09\n7.83E-11\n1.77E-12\n4.21E+00\n2.62E+01\nIIIIII\n3.62E-09\nHEAVIEST\nTOTAL\n6.76E+00\n4.16E+01\n|| ! ! ! ! !\n!!!!!!\n0.00E+00\nEEE\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\nEET\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n3.62E-09\n0.00E+00\n0.00E+00\n25 | Environmental Product Declaration: Two-door Cupboards\nOTHER\nENVIRONMENTAL\nINDICATORS\nUNIT\nSMALLEST UNIT\n\u2610 lllllll\nTWO-DOOR CUPBOARDS - ESSENTIALS RANGE\nLARGEST UNIT\nMEDIUM UNIT\n!!!!!!!\n!!!!!!!\n7!!!!!!\nHI!!!!!\n11\n!!!!!\nTOTAL\nI\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nHTC\ncases\n5.29E-05\n3.28E-07\n6.66E-07\n5.38E-05\nHTNC\ncases\n5.13E-05\n1.12E-06\n9.70E-07\n5.34E-05\nFWE\nkgPAF.\nm3.day\n3.86E+05\n2.65E+04\n7.01E+03\n4.19E+05\nLU\nspecies.yr\n1.78E-08\n9.61E-10\n2.78E-09\n2.15E-08\nAP\nkg SO2 eq\n2.85E-01\n1.61E-02\n1.43E-02\n3.15E-01\nODP\nkg CFC-11\neq\n4.59E-06\n3.19E-06\n5.73E-07\n8.36E-06\nBISLEY\nMEDIUM UNIT\nTwo-door Cupboards-Be\u2122\nEnvironmental indicator results are shown in the 5 following tables for the declared unit of 1 item of storage furniture in use for 15 years.\nTWO-DOOR CUPBOARDS - BETM RANGE\nSMALLEST UNIT\n\u2610\nLARGEST UNIT\n11/1/1111\n!!!!!!!111\n!!!!!!!!\nIIIIIIIIIII\n!!!!!!!!!II\nCORE\nPROCESSES\n1.95E-01\n3.53E-03\n1.22E+01\n1.04E-02\n4.54E-03\n2.61E-02\n2.44E-05\n2.02E+02\n3.61E+00\n!!!!!!!!!!!\nIIIIII\nENVIRONMENTAL\nUNIT\nIMPACT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\n9.15E+01\n1.11E+01\n6.07E+00\n1.20E+01\nfossil\nbiogenic\n2.58E+00\n1.67E-01\n1.54E+00\nGWP\nkg CO\u2082 eq\nland*\n1.34E-01\n2.81E-03\n4.23E-02\n9.43E+01\n1.13E+01\n7.65E+00\ntotal\nAP(fx)\nkg SO2 eq\n1.32E-01\n7.77E-03\n3.86E-03\nEP\nkg PO43-\n4.57E-02\n3.79E-03\n3.02E-03\neq\nPOCP\nKg NMVOC\neq\n4.01E-01\n2.12E-02\n1.60E-02\nADPE\nKg Sb eq\n5.80E-04\n1.92E-05\n4.11E-05\nADPFF\nMJ\n1.39E+03\n1.89E+02\n6.66E+01\nWSP\nm\u00b3 eq\n5.73E+01\n3.37E+00\n2.61E+00\n26 Environmental Product Declaration: Two-door Cupboards\nBISLEY\n27 | Environmental Product Declaration: Two-door Cupboards\nRESOURCE\nUNIT\nUSE\nUP\nPROCESSES\n\u2610 Kl\u2800\u2800\nTWO-DOOR CUPBOARDS - BETM RANGE\nMEDIUM UNIT\nLARGEST UNIT\nSMALLEST UNIT\n#{}}}\n///////I\n!!!!!!!!!!!\nIIIIII\n|| \\ / // /// III\n!!!!!!!!!!\n[!!!!!!!\n!!!!!!!1111\n!!!!!!!!!!!!\nIIIIIII\nPERE\nMJ\n7.46E+01\nPERM\nMJ\n1.50E+01\nPERT\nMJ\n8.96E+01\nPENRE\nMJ\n1.47E+03\nPENRM\nMJ\n7.98E+01\nPENRT\nMJ\n1.54E+03\nBISLEY\nSM\nkg\n2.40E+01\nRSF\nMJ\n0.00E+00\nNRSF\nMJ\n0.00E+00\nFW\nm3\n1.34E+00\n28 | Environmental Product Declaration: Two-door Cupboards\nTWO-DOOR CUPBOARDS - BET RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nLARGEST UNIT\nWASTES\nUNIT\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nDOWN\nPROCESSES PROCESSES\nHWD\nkg\n6.55E-02\n2.58E-03\n2.11E-03\n7.02E-02\n9.19E-02\n3.27E-03\n2.91E-03\n9.81E-02\n1.10E-01\n3.84E-03\n3.54E-03\nNHWD\nkg\n6.52E+00\n1.18E+00\n4.86E+00\n1.26E+01\n9.09E+00\n1.51E+00\n6.33E+00\n1.69E+01\n1.11E+01\n1.80E+00\n7.50E+00\nTRWD\nkg\n3.41E-03\n3.48E-03\n4.40E-04\n7.33E-03\n4.78E-03\n3.58E-03\n6.10E-04\n8.97E-03\n5.83E-03\n3.66E-03\n7.50E-04\n1.02E-02\nBISLEY\nTWO-DOOR CUPBOARDS - BET RANGE\nSMALLEST UNIT\nMEDIUM UNIT\nOUTPUT FLOWS\nUNIT\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nTOTAL\nUP\nPROCESSES\nCORE\nPROCESSES\nCFR\nkg\n0.00E+00\n0.00E+00\n4.89E+00\n4.89E+00\n0.00E+00\n0.00E+00\nMFR\nkg\n7.99E-02\n2.39E-03\n2.91E+01\n2.92E+01\n8.12E-02\n2.41E-03\nMER\nkg\n6.74E-11\n1.86E-12\n3.54E-09\n3.60E-09\n7.69E-11\n2.21E-12\nTOTAL\n6.89E+00\n4.06E+01\n!!!!!!\n3.61E-09\nEEE\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\nEET\nMJ\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\n0.00E+00\nLARGEST UNIT\nTOTAL\n8.48E+00\n!!!!!\n5.00E+01\n3.62E-09\n0.00E+00\n0.00E+00\n29 | Environmental Product Declaration: Two-door Cupboards\nOTHER\nENVIRONMENTAL\nINDICATORS\nUNIT\nSMALLEST UNIT\n\u2610 lllllll\nTWO-DOOR CUPBOARDS - BETM RANGE\nLARGEST UNIT\nMEDIUM UNIT\n[!!!!!!\nJIIIIII\nT!!!!!!\n!!!!!!!\n!!!!!\nTOTAL\nI\n1.\nUP\nCORE\nDOWN\nPROCESSES PROCESSES PROCESSES\nHTC\ncases\n6.84E-05\n3.72E-07\n8.77E-07\n6.97E-05\nHTNC\ncases\n6.78E-05\n1.21E-06\n1.21E-06\n7.02E-05\nFWE\nkgPAF.\nm3.day\n5.50E+05\n2.84E+04\n8.54E+03\n5.87E+05\nLU\nspecies.yr\n2.26E-08\n1.17E-09\n3.14E-09\n2.69E-08\nAP\nkg SO2 eq\n4.21E-01\n1.84E-02\n1.84E-02\n4.58E-01\nODP\nkg CFC-11\neq\n6.19E-06\n3.26E-06\n7.57E-07\n1.02E-05\nBISLEY\nInterpretation\nSteel accounts for more than 95% of the mass of the product. Steel production accounts for the largest\nfraction of the total indicator value across the life cycle for most environmental categories covered. Small\ncomponents, such as locks, are important for the resource depletion category - relative to the proportion of\nthe product's mass they represent -because they contain more specialised, scarcer materials than the main\nbody of the product.\nIndicator values obtained for human toxicity, ecotoxicity, land use, ODP and water scarcity should be used\nwith caution; all are subject to uncertainties in data or method which limit the scope for their use as the\nbasis for comparisons.\nADDITIONAL ENVIRONMENTAL INFORMATION\nAt Bisley, we take our impact on the environment very seriously. In 2015, The Furniture Makers' Guild\nawarded Bisley with the prestigious Manufacturing Guild Mark, which observes excellence in sustainability\nand production. We are members of the Furniture Industry Sustainability Program (FISP), Confederation of\nBritish Metal Formers (CBM) and the British Contract Furniture Association (BCFA).\nWe strive to ensure that the company adheres to and exceeds all environmental regulation, continuously\nevaluating the impact of our product and processes, in addition to guaranteeing that the physical products\nare safe and worthy of use. We favour steel as it is highly recyclable, durable and maintains its quality\nthrough the recycling process.\nAll our products are precision engineered and built to last at our facility in Newport which is certified to ISO\n9001, ISO 14001 and ISO 45001. When our products come to the end of their natural lifespan, we make\nsure that they are able to be recycled into other timber and metal based products.\n30 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\n\n\nREFERENCES\necoinvent database (v3.4) - www.ecoinvent.ch\nEurostat - European Packaging waste by waste management operations and waste flow, for France,\nGermany, the Netherlands, UK, Ireland and Spain. From www.eurostat.eu\nEN 15804:2012 + A1:2013 - Sustainability of construction works - Environmental Product Declarations -\nCore rules for the product category of construction products\nGeneral Program Instructions, Version 3.0, 2017-12-11 - The International EPD\u24c7 System - EPD\nInternational AB\nISO 14001:2015 - Environmental management systems - Requirements with guidance for use\nISO 14025:2009-11: Environmental labels and declarations - Type III environmental declarations - Principles\nand procedures\nPCR Furniture, except Seats and Mattresses (UN CPC 3812/3813/3814) 2012:19, Version 2.0-2019-06-17\nStorage Furniture LCA Report - EuGeos Limited, May 2019\nBOFA\nmember\nTUV NORD\nO\n0043\nR\nOF SYST\n0043\nUKAS\nMANAGEMENT\nISO 14001\nSYSTEMS\n0043\nREGISTRAR\nO\nCURS\nISO\n45001\nCERTIFIED\nOCCUPATIONAL\nHEALTH\nAND SAFETY\nSYSTEM\nManufacturing\nGuild Mark\nAwarded by The Furniture Makers' Company\nfisp\nFurniture Industry\nSustainability\nProgramme\nREGISTRAR\nURS\nGLOSSARY\nThe International EPD\u24c7 System: a programme for Type III environmental declarations, maintaining a\nsystem to verify and register EPDs as well as keeping a library of EPDS and PCRs in accordance with ISO\n14025. (www.environdec.com)\nLife cycle assessment (LCA): LCA studies the environmental aspects and quantifies the potential impacts\n(positive or negative) of a product (or service) throughout its entire life. ISO standards ISO 14040 and ISO\n14044 set out conventions for conducting LCA.\nREACH Regulation: REACH is the European Regulation on Registration, Evaluation, Authorisation and\nRestriction of Chemicals. It entered into force in 2007, replacing the former legislative framework for\nchemicals in the EU.\n31 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\nANNEX: APPLICABLE PRODUCT CODES\nThis EPD applies to Bisley's two-door cupboards with the codes listed below:\nLATERALFILE\u2122\nSYSTEMFILE\u2122\nESSENTIALS\nBETM\n08D20\nSYD08/2\nYECB0807/1S\nBH08071004NP\n08DD21\nSYD08/3\nYECBDH0807/1S\nBH10071004NP\n08D24\nSYD08/34\nYECB0810\nBH08111073NP\n08D30\nSYD08/4\nYECB0810/1S\nBH10111073NP\n08D32\nSYD08/43\nYECB0811\nBH08141158NP\n08D34\nSYD08/50\nYECB0811/1S\nBH10141158NP\n08D36\nSYD08/54\nYECB0812\nBH08181318NP\n08D40\nSYD08/6\nYECB0812/25\nBH10181318NP\n08D43\nSYD08/6/C\nYECB0815\n08D48\nSYD10/2\nYECB0815/3S\n08D50\nSYD10/3\n08D54\n08D60\nSYD10/4\n08D61\nSYD10/43\n08D64\nSYD10/50\n09D20\nSYD10/54\n09D30\nSYD10/6\n09D32\n09D34\n09D36\n09D60\nSYD10/34\nSYD10/6/C\nYECB0819\nYECB0819/4S\nYECB0910\nYECB0910/1S\nYECB0911\nYECB0911/1S\nYECB0912\nYECB0912/25\nYECB0915\nYECB0915/3S\nYECB0919\n09D64\n10D20\n10D24\n10D30\n10D32\n10D34\nYECB0919/4S\nYECBOF0910/1S\nYECBOF0911/1S\nYECBOF0912/2S\nYECBOF0915/3S\nYECBOF0919/4S\n10D36\n10D40\n10D43\n10D48\n10D50\n10D54\n10D60\n10D61\n10D64\nYECB1007/1S\nYECBDH1007/1S\nYECB1010\nYECB1010/1S\nYECB1011\nYECB1011/1S\nYECB1012\nYECB1012/25\nYECB1015\nYECB1015/3S\nYECB1019\nYECB1019/4S\n32 | Environmental Product Declaration: Two-door Cupboards\nBISLEY\n"}, "expected_output": {"claims": [{"unit": "%", "value": 56, "evidence": ["Steel used by Bisley has a recycled content typical of European steel production, quoted as 56% by steel\nsuppliers in 2018."]}]}, "metadata": {"product_category": "Furniture & other goods", "request_id": "req_662493a8619d9c43"}}
{"id": "8b5a7391af352e8cbbb0f103", "input": {"query": "What are the solution heat treatment and aging parameters (time in hours) for 7xxx series aluminum alloys like 7075, 7050, 7150?", "source_url": "https://materialsdata.nist.gov/bitstream/handle/11115/192/Heat%20Treating%20of%20Aluminum%20Alloys.pdf?sequence=3&isAllowed=y", "document_text": "ASM Handbook, Volume 4: Heat Treating\nASM Handbook Committee, p 841-879\nDOI: 10.1361/asmhba0001205\nCopyright \u00a9 1991 ASM International\u24c7\nAll rights reserved.\nwww.asminternational.org\nHeat Treating of Aluminum Alloys\nHEAT TREATING in its broadest sense,\nrefers to any of the heating and cooling\noperations that are performed for the pur-\npose of changing the mechanical properties,\nthe metallurgical structure, or the residual\nstress state of a metal product. When the\nterm is applied to aluminum alloys, howev-\ner, its use frequently is restricted to the\nspecific operations employed to increase\nstrength and hardness of the precipitation-\nhardenable wrought and cast alloys. These\nusually are referred to as the \"heat-treat-\nable\" alloys to distinguish them from those\nalloys in which no significant strengthening\ncan be achieved by heating and cooling. The\nlatter, generally referred to as \"non-heat-\ntreatable\" alloys, depend primarily on cold\nwork to increase strength. Heating to de-\ncrease strength and increase ductility (an-\nnealing) is used with alloys of both types;\nmetallurgical reactions may vary with type\nof alloy and with degree of softening desired.\nExcept for the low-temperature stabilization\ntreatment sometimes given for 5xxx series\nalloys (which is a mill treatment and not\ndiscussed in this article), complete or partial\nannealing treatments are the only ones used\nfor non-heat-treatable alloys. A general\noverview of these heat treatments is covered\nin the article \"Principles of Heat Treating of\nNonferrous Alloys\" in this Volume.\nPrecipitation from Solid Solution\nOne essential attribute of a precipitation-\nhardening alloy system is a temperature-\ndependent equilibrium solid solubility char-\nacterized by increasing solubility with\nincreasing temperature (see, for example,\nthe phase diagrams in Fig 1 and 2). Al-\nthough this condition is met by most of the\nbinary aluminum alloy systems, many ex-\nhibit very little precipitation hardening, and\nthese alloys ordinarily are not considered\nheat treatable. Alloys of the binary alumi-\nnum-silicon and aluminum-manganese sys-\ntems, for example, exhibit relatively insig-\nnificant changes in mechanical properties as\na result of heat treatments that produce\nconsiderable precipitation. The major alu-\nminum alloy systems with precipitation\nhardening include:\n\u2022\nsystems\nAluminum-copper-magnesium\n(magnesium intensifies precipitation)\nsystems\n\u2022 Aluminum-magnesium-silicon\nwith strengthening from Mg\u2082Si\nAluminum-zinc-magnesium systems with\nstrengthening from MgZn2\n\u2022\n\u2022 Aluminum-zinc-magnesium-copper\ntems\nsys-\nThe general requirement for precipitation\nstrengthening of supersaturated solid solu-\ntions involves the formation of finely dis-\npersed precipitates during aging heat treat-\nments (which may include either natural aging\nor artificial aging). The aging must be accom-\nplished not only below the equilibrium solvus\ntemperature, but below a metastable miscibil-\nity gap called the Guinier-Preston (GP) zone\nsolvus line. The supersaturation of vacancies\nallows diffusion, and thus zone formation, to\noccur much faster than expected from equi-\nlibrium diffusion coefficients. In the precipi-\ntation process, the saturated solid solution\nfirst develops solute clusters, which then be-\ncome involved in the formation of transitional\n(nonequilibrium) precipitates.\nTemperature, \u00b0C\n800\n600\n400\n200\n\u0391\u0399\nAl + CuAl2\nThe mechanism of strengthening from\nprecipitation involves the formation of co-\nherent clusters of solute atoms (that is, the\nsolute atoms have collected into a cluster\nbut still have the same crystal structure as\nthe solvent phase). This causes a great deal\nof strain because of mismatch in size be-\ntween the solvent and solute atoms. Conse-\nquently, the presence of the precipitate par-\nticles, and even more importantly the strain\nfields in the matrix surrounding the coher-\nent particles, provide higher strength by\nobstructing and retarding the movement of\ndislocations. The characteristic that deter-\nmines whether a precipitate phase is coher-\nent or noncoherent is the closeness of\nmatch or degree of disregistry between\natomic spacings on the lattice of the matrix\nand on that of the precipitate. These\nchanges in properties result from the forma-\ntion of solute-rich microstructural domains,\nor GP zones.\nThe exact size, shape, and distribution of\nGP zones depend on the alloy in which they\nform and on the thermal and mechanical\nhistory of the specimen. Their shape can\n(b)\nL\nA1 + L\nTemperature range for\nsolution heat treating\nTemperature range\nfor annealing\nTemperature range for\nprecipitation heat\ntreating\n2\n4\n6\n8\n10\n12\nCopper.%\n1400\n1000\n600\n200\nTemperature, \u00b0F\nAluminum-copper systems with strength- Fig 1 Portion of aluminum-copper binary phase diagram. Temperature ranges for annealing, precipitation heat\n\u2022\nening from CuAl2\ntreating, and solution heat treating are indicated. The range for solution treating is below the eutectic\nmelting point of 548 \u00b0C (1018 \u00b0F) at 5.65 wt% Cu.\n842 / Heat Treating of Nonferrous Alloys\nTemperature, \u00b0C\n(a)\nMagnesium, wt%\n(b)\n700\n600\nSolidus\n500\n400\nSolvus\n300\n200\n100\n1200\n1000\n595 \u00b0C\nat 1.85%\nMg2Si\n800\n600\n400\n200\nMg-Si ratio of 1.73:1\n1\n0\n0.5\n1.0\n1.5\n2.0\nMg2Si, %\n570\n1.0\nTemperature, \u00b0F\n660 750 840 930 1020 1110\n1.4\n1.2\n0.8\nSolvus with silicon\nand Mg2Si present\n1.0\n0.6\n0.8\n0.6\n0.4\n0.4\n0.2\n0\n300 350 400 450 500 550 600\nTemperature, \u00b0C\n0.2\n0\nTemperature, \u00b0F\nSilicon, wt%\nFig 2 Equilibrium solubility as function of tempera-\nture for (a) Mg2Si in aluminum with an Mg-Si\nratio of 1.73-to-1 and (b) magnesium and silicon in solid\naluminum when both Mg2Si and silicon are present\nsometimes be deduced by refined studies of\ndiffuse x-ray scattering. Under favorable\nconditions, GP zones can be seen in trans-\nmission electron micrographs. Spherical\nsolute-rich zones usually form when the\nsizes of the solvent and solute atoms are\nnearly equal, as in the aluminum-silver and\naluminum-zinc systems. If there is a large\ndifference in atom sizes, as in the alumi-\nnum-copper system, the GP zones usually\nform as disks whose planes lie parallel with\nsome low-index plane of the matrix lattice.\nSometimes, the solute atoms occupy pre-\nferred lattice sites within the GP zone, and\nthus form a small region of lattice order.\nThe GP zones are of the size range of tens\nof angstroms in diameter. They are essential-\nly distorted regions of the matrix lattice, rath-\ner than discrete particles of a new phase\nhaving a different lattice. As such, they are\ncompletely coherent with the matrix, impos-\ning local but often large strains on it. These\nmechanical strains, as well as the presence of\na locally solute-rich, sometimes ordered lat-\ntice, can account for large changes in me-\nchanical properties of the alloy before any\nlong-range microstructural changes occur.\nThe GP zones are characteristically meta-\nstable and thus dissolve in the presence of a\nmore stable precipitate. This dissolution\ncauses a precipitate-free, visibly denuded\nregion to form around the stable precipitate\nparticles. The final structure consists of\nequilibrium precipitates, which do not con-\ntribute as significantly to hardening. More\ndetailed information about preprecipitation\nphenomena can be found in the article\n\"Structures Resulting From Precipitation\nFrom Solid Solution\" in Volume 9 of the\n9th Edition of Metals Handbook.\nPrecipitation in Aluminum-Copper Alloys.\nFigure 1, which illustrates the required sol-\nubility-temperature relationship needed in\nprecipitation strengthening, shows the tem-\nperature ranges required for solution treat-\nment and subsequent precipitate hardening\nin the aluminum-copper system. The equi-\nlibrium solid solubility of copper in alumi-\nnum increases as temperature increases-\nfrom about 0.20% at 250 \u00b0C (480 \u00b0F) to a\nmaximum of 5.65% at the eutectic melting\ntemperature of 548 \u00b0C (1018 \u00b0F). (It is con-\nsiderably lower than 0.20% at temperatures\nbelow 250 \u00b0C.) For aluminum-copper alloys\ncontaining from 0.2 to 5.6% Cu, two distinct\nequilibrium solid states are possible. At\ntemperatures above the lower curve in Fig 1\n(solvus), the copper is completely soluble,\nand when the alloy is held at such temper-\natures for sufficient time to permit needed\ndiffusion, the copper will be taken com-\npletely into solid solution. At temperatures\nbelow the solvus, the equilibrium state con-\nsists of two solid phases: solid solution, a,\nplus an intermetallic-compound phase 0\n(Al\u2082Cu). When such an alloy is converted to\nall solid solution by holding above the sol-\nvus temperature and then the temperature is\ndecreased to below the solvus, the solid\nsolution becomes supersaturated and the\nalloy seeks the equilibrium two-phase con-\ndition; the second phase tends to form by\nsolid-state precipitation.\nThe preceding description is a gross over-\nsimplification of the actual changes that\noccur under different conditions even in\nsimple binary aluminum-copper alloys. A\nvariety of different nonequilibrium precipi-\ntate structures is formed at temperatures\nbelow solvus. In alloys of the aluminum-\ncopper system, a succession of precipitates\nis developed from a rapidly cooled super-\nsaturated solid solution (SSS). These pre-\ncipitates develop sequentially either with\nincreasing temperature or with increasing\ntime at temperature between room temper-\nature and the solvus. The several stages are\nidentified by the following notation:\nSSSGP zones \u21920\" \u2192\n0'\u21920 (Al\u2082Cu)\nAt temperatures in the natural aging\nrange (about 20 to 60 \u00b0C, or 0 to 140 \u00b0F),\nthe distribution of copper atoms changes\nwith time from random to the disklike pla-\nnar aggregates (GP zones), which form on\nparticular crystallographic planes of the alu-\nminum matrix. These aggregates create co-\nherency strain fields that increase resis-\ntance to deformation, and their formation is\nresponsible for the changes in mechanical\nproperties that occur during natural aging.\nAt higher temperatures, transition forms of\napproximate composition Al\u2082Cu develop\nand further increase strength. In the highest\nstrength condition, both the 0\" and 0' tran-\nsition precipitates may be present. When\ntime and temperature are increased suffi-\nciently to form high proportions of the equi-\nlibrium 0, the alloy softens and is said to be\n\"overaged.\"\nThe commercial heat-treatable aluminum\nalloys are, with few exceptions, based on\nternary or quaternary systems with respect\nto the solutes involved in developing\nstrength by precipitation. Commercial al-\nloys whose strength and hardness can be\nsignificantly increased by heat treatment\ninclude 2xxx, 6xxx, and 7xxx series wrought\nalloys (except 7072) and 2xx.0, 3xx.0, and\n7xx.0 series casting alloys. Some of these\ncontain only copper, or copper and silicon,\nas the primary strengthening alloy addi-\ntion(s). Most of the heat-treatable alloys,\nhowever, contain combinations of magne-\nsium with one or more of the elements\ncopper, silicon, and zinc. Characteristical-\nly, even small amounts of magnesium in\nconcert with these elements accelerate and\naccentuate precipitation hardening, while\nalloys in the 6xxx series contain silicon and\nmagnesium approximately in the propor-\ntions required for formulation of magnesium\nsilicide (Mg2Si). Although not as strong as\nmost 2xxx and 7xxx alloys, 6xxx series al-\nloys have good formability, weldability, ma-\nchinability, and corrosion resistance, with\nmedium strength.\nIn the heat-treatable wrought alloys, with\nsome notable exceptions (2024, 2219, and\n7178), such solute elements are present in\namounts that are within the limits of mutual\nsolid solubility at temperatures below the\neutectic temperature (lowest melting tem-\nperature). In contrast, some of the casting\nalloys of the 2xx.0 series and all of the 3xx.0\nseries alloys contain amounts of soluble\nelements that far exceed solid-solubility\nlimits. In these alloys, the phase formed by\ncombination of the excess soluble elements\nwith the aluminum will never be dissolved,\nalthough the shapes of the undissolved par-\nticles may be changed by partial solution.\nMost of the heat-treatable aluminum alloy\nsystems exhibit multistage precipitation and\nundergo accompanying strength changes\nanalogous to those of the aluminum-copper\nsystem. Multiple alloying additions of both\nmajor solute elements and supplementary\nelements employed in commercial alloys are\nstrictly functional and serve with different\nheat treatments to provide the many differ-\nent combinations of properties\u2014physical,\nmechanical, and electrochemical-that are\nrequired for different applications. Some al-\nloys, particularly those for foundry produc-\ntion of castings, contain amounts of silicon\nfar in excess of the amount that is soluble or\nneeded for strengthening alone. The function\nhere is chiefly to improve casting soundness\nand freedom from cracking, but the excess\nsilicon also serves to increase wear resis-\ntance, as do other microstructural constitu-\nents formed by manganese, nickel, and iron.\nParts made of such alloys are commonly\nused in gasoline and diesel engines (pistons,\ncylinder blocks, and so forth).\nAlloys containing the elements silver,\nlithium, and germanium are also capable of\nproviding high strength with heat treatment,\nand in the case of lithium, both increased\nelastic modulus and lower density, which\nare highly advantageous-particularly for\naerospace applications (see the following\nsection \"Aluminum-Lithium Alloys\" in this\narticle). Commercial use of alloys contain-\ning these elements has been restricted either\nby cost or by difficulties encountered in\nproducing them. Such alloys are used to\nsome extent, however, and research is be-\ning directed toward overcoming their disad-\nvantages.\nIn the case of alloys having copper as the\nprincipal alloying ingredient and no magne-\nsium, strengthening by precipitation can be\ngreatly increased by adding small fractional\npercentages of tin, cadmium, or indium, or\ncombinations of these elements. Alloys\nbased on these effects have been produced\ncommercially but not in large volumes be-\ncause of costly special practices and limita-\ntions required in processing, and in the case\nof cadmium, the need for special facilities to\navoid health hazards from formation and\nrelease of cadmium vapor during alloying.\nSuch alloys, as well as those containing\nsilver, lithium, or other particle-forming el-\nements, may be used on a selective basis in\nthe future.\nAluminum-Lithium Alloys. Like other age-\nhardened aluminum alloys, aluminum-lithi-\num alloys achieve precipitation strengthen-\ning by thermal aging after a solution heat\ntreatment. The precipitate structure is sen-\nsitive to a number of processing variables,\nincluding, but not limited to, the quenching\nrate following the solution heat treatment,\nthe degree of cold deformation prior to\naging, and the aging time and temperature.\nMinor alloying elements can also have a\nsignificant effect on the aging process by\nchanging the interface energy of the precip-\nitate, by increasing the vacancy concentra-\ntion, and/or by raising the critical tempera-\nture for homogeneous precipitation. Like\nsome other age-hardened 2xxx aluminum\nalloys, aluminum-lithium-base alloys also\ngain increased strength and toughness from\ndeformation prior to aging. This unusual\nphenomenon has given rise to a number of\nthermomechanical processing steps for alu-\nminum-lithium alloys aimed at optimizing\nmechanical properties after artificial aging.\nThe age hardening of aluminum-lithium\nalloys involves the continuous precipitation\nof d' (AlLi) from a supersaturated solid\nsolution. The aluminum and lithium in the 8'\nprecipitates are positioned at specific loca-\ntions. The eight shared corner sites are\noccupied by lithium, and the six shared\nfaces are occupied by aluminum. This gives\nrise to the aluminum-lithium composition of\n8' precipitates. The geometrical similarity\nbetween the lattice of the precipitates and\nthe face-centered cubic lattices of the solid\nsolution facilitates the observed cube/cube\norientation. The lattice parameters of the\nprecipitate are also closely matched to\nthose of the matrix. Consequently, the mi-\ncrostructure of an aluminum-lithium alloy\nsolution heat treated and aged for short\ntimes below the d' solvus is characterized\nby a homogeneous distribution of coherent,\nspherical d' precipitates.\nAluminum-lithium-base alloys are micro-\nstructurally unique. They differ from most\nof the aluminum alloys in that once the\nmajor strengthening precipitate (8') is ho-\nmogeneously precipitated, it remains coher-\nent even after extensive aging. In addition,\nextensive aging at high temperatures (>190\n\u00b0C, or 375 \u00b0F) can result in the precipitation\nof icosahedral grain-boundary precipitates\nwith five-fold symmetry. Although the qua-\nsi-crystalline structure and the composition\nof these grain-boundary precipitates are not\nyet exactly known, it has been suggested\nthat both the precipitates and the precipi-\ntate-free zones (PFZs) near the grain bound-\naries might play a major role in the fracture\nprocess.\nThe low ductility and toughness of binary\naluminum-lithium alloys can be traced, at\nleast in part, to the inhomogeneous nature\nof their slip, resulting from coherent-parti-\ncle hardening of spherical d' precipitates.\nThe presence of equilibrium 6 (aluminum-\nlithium) precipitates at grain boundaries can\nalso cause PFZs, which can induce further\nstrain localization and promote intergranu-\nlar failure. Consequently, for the develop-\nment of commercial alloys, slip has been\nhomogenized by introducing dispersoids\n(manganese, zirconium) and semicoherent/\nincoherent precipitates, such as\n(Al\u2082CuLi), 0' (Al\u2082Cu), or S (Al\u2082LiMg),\nthrough copper or magnesium additions.\nMagnesium and copper improve the\nstrength of aluminum-lithium alloys through\nsolid-solution and precipitate strengthening,\nand they can minimize the formation of\nPFZs near grain boundaries. Zirconium,\nwhich forms the cubic Al3Zr coherent dis-\npersoid, stabilizes the subgrain structure\nand suppresses recrystallization.\nT\u2081\nDevelopment of commercially available\naluminum-lithium-base alloys was started\nby adding lithium to aluminum-copper, alu-\nminum-magnesium, and aluminum-copper-\nHeat Treating of Aluminum Alloys / 843\nmagnesium alloys. These alloys were cho-\nsen to superimpose the precipitation-\nhardening characteristics of aluminum-\ncopper-, aluminum-copper-magnesium-,\nand aluminum-magnesium-base precipitates\nto the hardening of lithium-containing pre-\ncipitates. Proceeding in this manner, alloys\n2020 (Al-Cu-Li-Cd), 01429 (Al-Mg-Li), 2090\n(Al-Cu-Li), and 2091 and 8090 (Al-Cu-Mg-\nLi) evolved. Besides these registered al-\nloys, other commercial aluminum-lithium\nalloys include Weldalite 049 and CP276.\nProperties and applications of these alloys\nare discussed in the article \"Aluminum-\nLithium Alloys\" in Volume 2 of the 10th\nEdition of Metals Handbook.\nIn terms of d' precipitation, the only\neffect of magnesium appears to be a reduc-\ntion in the solubility of lithium. The micro-\nstructure of an aluminum-magnesium-lithi-\num alloy in the early stages of aging is\nsimilar to that of an aluminum-lithium alloy.\nPrecipitation in the aluminum-copper-lithi-\num system is more complicated than that in\neither the aluminum-lithium or aluminum-\nmagnesium-lithium systems.\nEffects on Physical and Electrochemical\nProperties. The above description of the\nprecipitation processes in commercial heat-\ntreatable aluminum alloys (as well as the\nheat-treatable binary alloys, none of which\nis used commercially in wrought form) af-\nfect not only mechanical properties but also\nphysical properties (density and electrical\nand thermal conductivities) and electro-\nchemical properties (solution potential). On\nthe microstructural and submicroscopic\nscales, the electrochemical properties de-\nvelop point-to-point nonuniformities that\naccount for changes in corrosion resistance.\nMeasurements of changes in physical and\nelectrochemical properties have played an\nimportant role in completely describing pre-\ncipitation reactions and are very useful in\nanalyzing or diagnosing whether heat-treat-\nable products have been properly or im-\nproperly heat treated. Although they may\nbe indicative of the strength levels of prod-\nucts, they cannot be relied upon to deter-\nmine whether or not the product meets\nspecified mechanical-property limits. Since\nelements in solid solution are always more\nharmful to electrical conductivity than the\nsame elements combined with others as\nintermetallic compounds, thermal treat-\nments are applied to ingots used for fabri-\ncation of electrical conductor parts. These\nthermal treatments are intended to precipi-\ntate as much as possible of the dissolved\nimpurities. Iron is the principal element\ninvolved, and although the amount precip-\nitated is only a few hundredths of a percent,\nthe effect on electrical conductivity of the\nwire, cable, or other product made from the\ningot is of considerable practical impor-\ntance. These alloys may or may not be heat\ntreatable with respect to mechanical prop-\nerties. Electrical conductor alloys 6101 and\n844 / Heat Treating of Nonferrous Alloys\n6201 are heat treatable. These alloys are\nused in tempers in which their strengthening\nprecipitate, the transition form of Mg\u2082Si, is\nlargely out of solid solution to optimize both\nstrength and conductivity.\nStrengthening by Heat Treatment\n\u2022\n\u2022\nHeat treatment to increase strength of\naluminum alloys is a three-step process:\nSolution heat treatment: dissolution of\nsoluble phases\nQuenching: development of supersatura-\ntion\nAge hardening: precipitation of solute at-\noms either at room temperature (natural\naging) or elevated temperature (artificial\naging or precipitation heat treatment)\nEach of these steps and the use of quench-\nfactor analysis are described in the follow-\ning four sections. Typical solution and pre-\ncipitation heat treatments for mill products\nare given in Tables 1(a, b, and c) and 2, and\ntreatments for castings are given in Table 3.\nTemper designations are defined at the end\nof this article.\nSolution Heat Treating\nTo take advantage of the precipitation-\nhardening reaction, it is necessary first to\nproduce a solid solution. The process by\nwhich this is accomplished is called solution\nheat treating, and its objective is to take into\nsolid solution the maximum practical\namounts of the soluble hardening elements\nin the alloy. The process consists of soaking\nthe alloy at a temperature sufficiently high\nand for a time long enough to achieve a\nnearly homogeneous solid solution.\nNominal commercial solution heat-treat-\ning temperature is determined by the com-\nposition limits of the alloy and an allowance\nfor unintentional temperature variations.\nAlthough ranges normally listed allow vari-\nations of \u00b16 \u00b0C (\u00b110 \u00b0F) from the nominal,\nsome highly alloyed, controlled-toughness,\nhigh-strength alloys require that tempera-\nture be controlled within more restrictive\nlimits. Broader ranges may be allowable for\nalloys with greater intervals of temperature\nbetween their solvus and eutectic melting\ntemperatures.\nOverheating. Care must be exercised to\navoid exceeding the initial eutectic melting\ntemperature. If appreciable eutectic melting\noccurs as a result of overheating, properties\nsuch as tensile strength, ductility, and frac-\nture toughness may be degraded. Materials\nthat exhibit microstructural evidence of\noverheating are generally categorized as\nunacceptable by specification. Evidence of\ngrain-boundary melting that occurs above\nthe eutectic melting temperature of the alloy\nusually is not detectable by either visual\nexamination or nondestructive testing.\nAlthough maximum temperature must be\nrestricted to avoid melting, the lower limit\nshould, when possible, be above the temper-\nature at which complete solution occurs (sol-\nvus). In the alloy represented by line (a) in Fig\n1, these temperatures would be about 575 and\n515 \u00b0C (1065 and 960 \u00b0F), respectively. How-\never, under production conditions, the tem-\nperature interval for solution treatment\n(shown in Fig 1 for typical 2xxx or 2xx.x)\nalloys provides a margin to safeguard against\neutectic melting and a cushion on the low side\nfor increased solution and diffusion rates.\nFor alloys containing more than 5.65%\nCu, complete solution can never occur. For\nthese alloys, such as alloy 2219 (which has\n5.8 to 6.8% Cu), the minimum solution heat-\ntreating temperature is established so that it\nis as close as practical to the eutectic tem-\nperature while providing a margin of safety\ncommensurate with the capability of the\nequipment. Line (b) in Fig 1 is another\nexample of a composition above 5.65% Cu\nthat does not allow complete dissolution of\naluminum-copper precipitates.\nFor more complex ternary and quaterna-\nry systems, solution treatments are modi-\nfied according to the effect of new elements\non the solid solubility and/or the eutectic\nmelting points of the basic binary system. In\naluminum-lithium alloys, for example, mag-\nnesium reduces the solubility of lithium in\naluminum. In the aluminum-copper system,\nmagnesium also lowers the eutectic melting\npoint. The proximity of typical solution-\ntreating temperature ranges to eutectic\nmelting temperatures for three common alu-\nminum-copper-magnesium alloys is shown\nin the following table:\nSolution-treating\ntemperature\nEutectic melting\ntemperature\n\u00b0F\nAlloy\n\u00b0C\n\u00b0F\n\u00b0C\n2014\n2017\n2024\n496-507\n496-507\n488-499\n925-945\n925-945\n910-930\n510\n950\n513\n502\n955\n935\nSimilar considerations apply to other age-\nhardenable alloy systems such as aluminum-\nmagnesium-silicon alloys. For example, ac-\ncording to Fig 2(a), a 1.08% Mg\u2082Si alloy\nwould be soaked at a temperature in excess of\n500 \u00b0C (930 \u00b0F) but below the solidus of 595 \u00b0C\n(1100 \u00b0F) to avoid incipient melting. Howev-\ner, because some alloy constituents may form\ncomplex eutectics that melt at temperatures\nbelow the equilibrium eutectic temperature,\nthe upper limit for solution treatment of alu-\nminum-magnesium-silicon alloys is in the\nrange of 515 to 540 \u00b0C (960 to 1000 \u00b0F). At 540\n\u00b0C (1000 \u00b0F), about 0.6% Mg can be placed in\nsolution (Fig 2b).\nNonequilibrium Melting. When high heat-\ning rates are employed, the phenomenon of\nnonequilibrium melting must be considered.\nThis phenomenon can also be explained with\nthe help of the aluminum-copper phase dia-\ngram (Fig 1). The room-temperature micro-\nstructure of an F-temper product containing\n4% Cu consists of a solid solution of copper in\naluminum and particles of Al\u2082Cu. When this\nproduct is heated slowly, the Al\u2082Cu begins to\ndissolve, and if heating is slow enough, all of\nthe Al,Cu is dissolved when temperatures\nabove the solvus (500 \u00b0C, or 932 \u00b0F) are\nreached. When the heating rate is high, how-\never, much of the Al\u2082Cu remains undis-\nsolved. If a material with this microstructure\nis heated at or above the eutectic temperature\nof 548 \u00b0C (1018 \u00b0F), melting will begin at the\ninterface between the Al\u2082Cu and the matrix.\nWith sufficient time above the eutectic tem-\nperature, this metastable liquid will dissolve\nto form a solid solution and will leave no trace\nprovided that hydrogen gas has not con-\ndensed at the interface to form a void. If the\nproduct is quenched before the liquid has\ntime to equilibrate, however, it will solidify\nand form fine eutectic rosettes. This nonequi-\nlibrium melting should not be confused with\ntrue equilibrium melting, which would occur\nin any alloy containing more than 5.65% Cu.\nIn such an alloy, eutectic melting is equilibri-\num melting. No matter how long such an\nalloy is held above the eutectic temperature,\nthe liquid will never solidify. In commercial\nalloys, which usually are ternaries or quater-\nnaries of the major alloying elements, the\nsituation is more complex. Different phases\nhave different solvus temperatures, and non-\nequilibrium melting may occur at different\ntemperatures depending on composition, size\nof precipitates, and rate of heating. When\nnew solution heat-treating equipment (which\nprovides higher heating rates) is employed,\ncareful examination of alloy microstructures\nshould be included as part of the certification\nprocess.\nUnderheating. When the temperatures at-\ntained by the parts or pieces being heat treat-\ned are appreciably below the normal range,\nsolution is incomplete, and strength some-\nwhat lower than normal is expected. In the\naluminum-copper system (Fig 1), the shallow\nslope of the solvus at its intersection with the\ncomposition line indicates that a slight de-\ncrease in temperature will result in a large\nreduction in the concentration of the solid\nsolution and a correspondingly significant de-\ncrease in final strength. The effect of solution-\ntreating temperature on the strength of two\naluminum alloys is illustrated by the following\ndata:\n\u00b0C\nSolution-\ntreating\ntemperature\n\u00b0F\nMPa\nTensile strength\nksi\nYield strength\nMPa\nksi\n6061-T6 sheet 1.6 mm (0.064 in.) thick\n493\n920\n301\n43.7\n272\n39.4\n504\n940\n316\n45.8\n288\n41.7\n516\n960\n333\n48.3\n305\n44.3\n527\n980\n348\n50.5\n315\n45.7\n2024-T4 sheet 0.8 mm (0.032 in.) thick\n488\n910\n419\n60.8\n255\n37.0\n491\n915\n422\n61.2\n259\n37.5\n493\n920\n433\n62.8\n269\n39.0\n496\n925\n441\n63.9\n271\n39.3\nTable 1(a)\nalloying\nHeat Treating of Aluminum Alloys / 845\nTypical solution and precipitation heat treatments for commercial heat-treatable aluminum alloy mill products with copper\nAlloy\nProduct form\nAl-Cu alloys without magnesium alloying\nSolution heat treatment(a)\nPrecipitation heat treatment\nMetal temperature(b)\nMetal temperature(b)\nTemper\nTime(c),\n\u00b0C\n\u00b0F\ndesignation\n\u00b0C\n\u00b0F\nh\nTemper\ndesignation\n2011\nRolled or cold finished rod and bar\n525\n975\nT3(d)\n160\n320\n14\nT8(d)\nT4\nT451(e)\n2025\nDie forgings\n515\n960\nT4\n170\n340\n2219(f)\nFlat sheet\n535\n995\nT31(d)\n175\n350\nT37(d)\n165\n325\nT42\n190\n375\nPlate\n535\n995\nT31(d)\n175\n350\nT37(d)\n175\n350\nT351(e)\n175\n350\nT42\n190\n375\n2219(f)\nRolled or cold finished wire, rod, and bar\nExtruded rod, bar, shapes, and tube\n535\n995\nT351(e)\n190\n375\n535\n995\nT31(d)\n190\n375\nT3510(e)\n190\n375\nT3511(e)\n190\n375\nT42\n190\n375\nDie forgings and rolled rings\n535\n995\nT4\n190\n375\nHand forgings\nAl-Cu-Mg alloys\n535\n995\nT4\n190\n375\nT352(f)\n175\n350\n00000000000000\n10\nT6\n18\nT81(d)\n24\nT87(d)\n36\nT62\n18\nT81(d)\n18\nT87(d)\n18\nT851(e)\n36\nT62\n18\nT851(e)\n18\nT81(d)\n18\nT8510(e)\n18\n36\n26\n26\n18\n2018\n2024(h)\nDie forgings\nFlat sheet\n510(g)\n950(g)\nT4\n170\n340\n495\n920\nT3(d)\n190\n375\nT361(d)\n190\n375\nT42\n190\n375\n190\n375\n2024(h)\nCoiled sheet\n495\n920\nT4\nT42\n190\n375\n190\n375\nPlate\n495\n920\nRolled or cold finished wire, rod, and bar\n495\n920\n22\nT351(e)\n190\n375\nT361(d)\n190\n375\nT42\n190\n375\nT4\n190\n375\nT351(e)\n190\n375\nT36(d)\n190\n375\nT42\n190\n375\nExtruded rod, bar, shapes, and tube\n495\n920\nT3\n190\n375\nT3510(e)\n190\n375\nT3511(e)\n190\n375\nT42\n190\n375\nDrawn tube\n2036\nSheet\n2038\nSheet\n2218\nDie forgings\n495\n920\nT3(d)\nT42\n500\n930\nT4\n540\n1000\nT4\n205\n400\n510(g)\n950(g)\nT4\n170\n340\n10\n510(i)\n950(i)\nT41\n240\n460\n02826:26282228622261 206\nT8511(e)\nT62\nT6\nT6\nT852f)\nT61\nT81(d)\nT861(d)\n9\nT62\nT72\n9\nT62\nT72\nT851(e)\nT861(d)\n9\nT62\nT6\nT851(e)\nT86(d)\nT62\nT81\nT8510(e)\nT8511(e)\nT62\nT6\nT61\nT72\nAl-Cu-Mg-Si alloys\n2008\nSheet\n510\n950\nT4(d)(j)\n205\n400\n1\nT62(e)\n2014(h)\nFlat sheet\n500\n935\nT3(d)\n160\n320\n18\nT62\nT42\n160\n320\n18\nT6\nCoiled sheet\n500\n935\nT4\n160\n320\n18\nT6\nT42\n160\n320\n18\nT62\nPlate\n500\n935\nT42\n160\n320\n18\nT62\nT451(e)\n160\n320\n18\nT651(e)\nRolled or cold finished wire, rod, and bar\n500\n935\nT4\n160(k)\n320(k)\n18\nT6\nT42\n160(k)\n320(k)\n18\nT451(e)\n160(k)\n320(k)\n18\nT62\nT651(e)\nExtruded rod, bar, shapes, and tube\n500\n935\nT4\n160(k)\n320(k)\n18\nT6\nT42\n160(k)\n320(k)\n18\nT4510(e)\n160(k)\n320(k)\n18\nT62\nT6510(e)\nDrawn tube\n500\n935\nT4\n160(k)\n320(k)\n18\nT6\nT42\n160(k)\n320(k)\n18\nT62\nDie forgings\n500(1)\n935(1)\nT4\n170\n340\n10\nT6\n2017\nRolled or cold finished wire, rod, and bar\n500\n935\nT4\nT42\n2117\nRolled or cold finished wire and rod\n500\n935\nT4\nT42\n2618\nForgings and rolled rings\n530\n985\nT4\n200\n390\n20\nT61\n4032\nDie forgings\n510(h)\n950(h)\nT4\n170\n340\n10\nT6\nAl-Cu-Li alloys\n2090\n2091\n8090\nCP276\nSheet\nSheet\nExtruded bar\nExtruded bar\n540\n1000\nT3(d)\n165\n325\n24\nT83(d)\n530\n990\nT3(d)\n120\n250\n24\nT84(d)\n530\n990\nT3(d)\n190\n375\n12\nPeak aged(d)\n530\n990\nT3(d)\n190\nExtruded bar\n540\n1000\nT3(d)\n190\n375\n375\n12\nPeak aged(d)\n12-15\nPeak aged(d)\n(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water.\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) Cold working subsequent to solution heat treatment and prior to any precipitation heat treatment is necessary\nto attain the specified properties for this temper. (e) Stress relieved by stretching to produce a specified amount of permanent set subsequent to solution heat treatment and prior to any precipitation heat\ntreatment. (f) Stress relieved by 1 to 5% cold reduction after solution treatment and prior to precipitation heat treatment. (g) Quenched in water at 100 \u00b0C (212 \u00b0F). (h) These heat treatments also apply to\nalclad sheet and plate of these alloys. (i) Quenched with room-temperature air blast. (j) See U.S. Patent 4,840,852. (k) An alternative heat treatment of 8 h at 177 \u00b0C (350 \u00b0F) may also be used. (1) Quenched\nin water at 60 to 80 \u00b0C (140 to 180 \u00b0F).\n846 / Heat Treating of Nonferrous Alloys\nTable 1(b) Typical solution and precipitation heat treatments for Mg-Si aluminum alloys (6xxx series alloys)\nSolution heat treatment(a)\nPrecipitation heat treatment\nMetal temperature(b)\nMetal temperature(b)\nAlloy\n6005\nProduct form\n\u00b0C\n\u00b0F\nTemper\ndesignation\n\u00b0C\n\u00b0F\nTime(c), h\nTemper\ndesignation\nExtruded rod, bar, shapes,\n530(d)\n985(d)\nTI\n175\n350\n8\nT5\nand tube\n6009(e)\nSheet\n555\n1030\nT4\n205\n400\n1\nT6(e)\n6010\nSheet\n565\n1050\nT4\n205\n400\n1\nT6(e)\n6053\nDie forgings\n520\n970\nT4\n170\n340\n10\nT6\n6061(f)\nSheet\n530\n985\nT4\n160\n320\n18\nT6\nT42\n160\n320\n18\nT62\nPlate\n530\n985\nT4(g)\n160\n320\n18\nT6(g)\nT42\n160\n320\n18\nT62\nT451(h)\n160\n320\n18\nT651(h)\nRolled or cold finished\n530\n985\nT4\n160(i)\n320(i)\n18\nT6\nwire, rod, and bar\n160(i)\n320(i)\n18\nT89(j)\n160(i)\n320(i)\n18\nT93(k)\n160(i)\n320(i)\n18\nT913(k)\n160(i)\n320(i)\n18\nT94(k)\nT42\n160(i)\n320(i)\n18\nT62\nT451(h)\n160(i)\n320(i)\n18\nT651(h)\nExtruded rod, bar, shapes,\nand tube\n530(d)\n985(d)\nT4\n175\n350\n8\nT6\nT4510(h)\n175\n350\n8\nT4511(h)\n175\n350\n8\nT6510(h)\nT6511(h)\n530\n985\nT42\n175\n350\n8\nT62\n6061(f)\nDrawn tube\n530\n985\nT4\n160(i)\n320(i)\n18\nT6\nT42\n160(i)\n320(i)\n18\nT62\nDie and hand forgings\n530\n985\nT4\n175\n350\n8\nT6\nRolled rings\n530\n985\nT4\n175\n350\n8\nT452(1)\n175\n350\n8\nT6\nT652(1)\n6063\nExtruded rod, bar, shapes,\n(d)\n(d)\nTI\n205(m)\n400(m)\n1\nT5\nand tube\n520(d)\n970(d)\nT4\n175(n)\n350(n)\n520\n970\nT42\n175(n)\n350(n)\nDrawn tube\n520\n970\nT4\n175\n350\n175\n350\n175\n350\n175\n350\nT42\n175\n350\n6013(0)\nSheet\n570\n1055\nW(p)\n190\n375\nPlate\n570\n1055\nW(p)\n190\n375\n\u221e \u221e \u221e \u221e \u221e \u221e \u221e \u221e\n8\nT6\n8\nT62\n8\nT6\n8\nT83(j)(d)\n8\nT831(j)(d)\n8\nT832(j)(d)\n8\nT62\n4\nT6\n4\nT651\n6066\nExtruded rod, bar, shapes,\n530\n990\nT4\n175\n350\n8\nT6\nand tube\nT42\nDrawn tube\n530\n990\nT4\nT42\nDie forgings\n530\n990\nT4\n6070\nExtruded rod, bar, shapes,\n545(d)\n1015(d)\nT4\nand tube\nT42\n6111\nSheet\n560\n1040\nT4\n6151\nDie forgings\n515\n960\nT4\nRolled rings\n515\n960\nT4\n6262\nRolled or cold finished\n540\n1000\nT4\nwire, rod, and bar\nT42\n6262\nExtruded rod, bar, shapes,\nand tube\n540(d)\n1000(d)\nT4\n540\nDrawn tube\n540\n1000\n1000\nT42\nT4\nT42\n6463\nExtruded rod, bar, shapes,\nand tube\n(d)\n(d)\nTI\n520(d)\n6951\nSheet\n520\n530\n970(d)\n970\n985\nT4\nT42\nE = FEE FEE FERRE PREFECE\n175\n350\nT4510(h)\n175\n350\nT4511(h)\n175\n350\n175\n350\n175\n350\n175\n350\nao ao ao ao \u221e \u221e \u221e\n8\nT62\n8\nT6510(h)\n8\nT6511(h)\n8\nT6\n8\nT62\n8\nT6\n160\n320\n18\nT6\n160\n320\n18\nT62\n175\n350\n8\nT6(q)\n170\n340\n10\nT6\n170\n340\n10\nT6\nT452(1)\n170\n340\n10\nT652(1)\n170\n340\n8\nT6\n170\n340\n12\nT451\n170\n340\n170\n340\n175\n350\n12\n2882\nT9(k)\nT651(h)\nT62\nT6\nT4510(h)\n175\n350\n12\n175\n350\n12\nT6510(h)\nT62\n170\n340\n8\nT6\n170\n340\n8\nT9(k)\n170\n340\n8\nT62\n205(m)\n400(m)\n1\nT5\n175(n)\n350(n)\n8\nT6\n175(n)\n350(n)\n8\nT62\nT4\n160\n320\n18\nT6\nT42\n160\n320\n18\nT62\n(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water,\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) By suitable control of extrusion temperature, product may be quenched directly from extrusion press to\nprovide specified properties for this temper. Some products may be adequately quenched in room-temperature air blast. (e) Alternate heat treatments of 4 h at 190 \u00b0C (375 \u00b0F) or 8 h at 175 \u00b0C (350 \u00b0F) may\nalso be used. See U.S. Patent 4,082,578. (f) These heat treatments also apply to alclad sheet and plate in these alloys. (g) Applicable to tread plate only. (h) Stress relieved by stretching to produce a specified\namount of permanent set prior to precipitation heat treatment. (i) An alternative heat treatment of 8 h at 170 \u00b0C (340 \u00b0F) also may be used. (j) Cold working after solution treatment is necessary to attain\nspecified properties during precipitation heat treatments. (k) Cold working after precipitation heat treatment is necessary to attain specified properties. (1) Stress relieved by 1 to 5% cold reduction subsequent\nto solution heat treatment and prior to precipitation heat treatment. (m) An alternative treatment of 3 h at 182 \u00b0C (360 \u00b0F) also may be used. (n) An alternative treatment of 6 h at 182 \u00b0C (360 \u00b0F) also may\nbe used. (o) See U.S. Patent 4,589,932. (p) Two weeks of natural aging to a T4 condition. (q) Artificially aged in laboratory from T4 to T6.\nTable 1(c)\nHeat Treating of Aluminum Alloys / 847\nTypical solution and precipitation heat treatments for heat-treatable Zn-Mg aluminum alloys from the 7xxx series\nSolution heat treatment(a)\nPrecipitation heat treatment\nMetal temperature(b)\nMetal temperature(b)\nAlloy\n7001\nProduct form\n\u00b0C\n\u00b0F\nTemper\ndesignation\n\u00b0C\n\u00b0F\nTime(c),\nh\nTemper\ndesignation\nExtruded rod, bar, shapes,\nand tube\n465\n870\nW\n120\n250\n24\nT6\n120\n250\n24\nT62\nW510(d)\n120\n250\n24\nT6510(d)\nW511(d)\n120\n250\n24\nT6511(d)\n7005\nExtruded rod, bar, and\n:\nT53(e)\nshapes\n7050\nPlate\n475\n890\nW51(d)\n(f)\n(f)\n(f)\nT7651(g)\n(h)\n(h)\n(h)\nT7451(g)\nExtrusions\n475\n890\nW510(d)\n(f)\n(f)\n(f)\nT76510(g)\nW511(d)\n(f)\n(f)\n(f)\nT76511(g)\nDie and hand forgings\n475\n890\nW\n(h)\n(h)\n(h)\nW52(d)\n(h)\n(h)\n(h)\nT74(g)\nT7452(g)\n7075(i)\nSheet\n480\n900\nW\n120(j)\n250(j)\n24\nT6\n120(j)\n250(j)\n24\nT62\n(f)\n(f)\n(f)\nT76(g)\n(h)(k)\n(h)(k)\n(h)(k)\nT73(g)\nPlate\n480\n900\nW\n\u0791\u0791\n120(j)\n250(j)\n24\nT62\nW51(d)\n(h)(k)\n(h)(k)\n(h)(k)\n120(j)\n250(j)\n24\n(f)\n(f)\n(f)\nT7351(d)(g)\nT651(d)\nT7651(g)\n7075(i)\nRolled or cold finished\nwire, rod, and bar\n490\n915\nW\n120\n250\n24\nT6\n120\n250\n24\n(h)(k)\n(h)(k)\n(h)(k)\nW51(d)\n120\n250\n(h)(k)\n(h)(k)\n24\n(h)(k)\nT62\nT73(g)\nT651(d)\nT7351(d)(g)\nExtruded rod, bar, shapes,\nand tube\n465\n870\nW\n120(1)\n250(1)\n24\nT6\n120(1)\n250(1)\n24\nT62\n(h)(k)\n(h)(k)\n(h)(k)\nT73(g)\n(f)\n(f)\n(f)\nW510(d)\n120(1)\n250(1)\n24\n(h)(k)\n(h)(k)\n(h)(k)\nT76(g)\nT6510(d)\nT73510(d)(g)\n(f)\n(f)\n(f)\nW511(d)\n120(1)\n250(1)\n24\n(h)(k)\n(h)(k)\n(h)(k)\n(f)\n(f)\n(f)\nT76510(g)\nT6511(d)\nT73511(d)(g)\nT76511(g)\nDrawn tube\n465\n870\n=\nW\n120\n250\n24\nT6\n120\n250\n24\n(h)(k)\n(h)(k)\n(h)(k)\nT62\nT73(g)\nDie forgings\n470(m)\n880(h)\n3\n120\n250\n24\n(h)\n(h)\n(h)\nW52(n)\n(h)\n(h)\n(h)\nHand forgings\n470(m)\n880(h)\nW\n120\n250\n24\n(h)\n(h)\n(h)\nW52(n)\n120\n250\n24\n(h)\n(h)\n(h)\nRolled rings\n470\n880\n7175\nDie forgings\n(o)\n(o)\n(o)\n(o)\n(0)\nHand forgings\n(o)\n(0)\n(0)\n(o)\nBBBBBB\nW\n120\n250\n24\nT6\nW\n(o)\n(0)\n(o)\nW\n(o)\n(o)\n(o)\nW52(n)\n(o)\n(o)\nW\n(0)\nW52(n)\n(0)\n(0)\n(o)\nT6\nT73(g)\nT7352(n)(g)\nT6\nT73(g)\nT652(n)\nT7352(n)(g)\nT66(0)\nT74(g)(0)\nT7452(n)(g)(0)\nT74(g)(0)\nT7452(n)(g)(0)\n7475\nSheet\n515(p)\n960(p)\nW\n120\n250\n3\nplus 155\n315\n3\n(f)\n(f)\n(f)\nT61(p)\nT761(g)(p)\nPlate\n510(p)\n950(p)\nW51(d)\n120\n250\n24\n(f)\n(f)\nT651(p)\nT7651(g)(p)\n(h)\n(h)\n(h)\nT7351(g)(p)\nAlclad\n7475\nSheet\n495\n920\nW\n120\n250\n3\nplus 155\n(f)\n315\n3\n(f)\n(f)\nT61(p)\nT761(g)(p)\n(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water,\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) Stress relieved by stretching to produce a specified amount of permanent set after solution treatment and\nprior to precipitation heat treatment. (e) No solution heat treatment; 72 h at room temperature following press quench, followed by two-stage precipitation heat treatment comprised of 8 h at 107 \u00b0C (225 \u00b0F)\nplus 16 h at 149 \u00b0C (300 \u00b0F). (f) Aging practice varies with product, size, nature of equipment, loading procedures, and furnace-control capabilities. The optimum practice for a specific item can be ascertained\nonly by actual trial treatment of the item under specific conditions. Typical procedures involve a two-stage treatment comprised of 3 to 30 h at 121 \u00b0C (250 \u00b0F) followed by 15 to 18 h at 163 \u00b0C (325 \u00b0F) for\nextrusions. An alternative two-stage treatment of 8 h at 99 \u00b0C (210 \u00b0F) followed by 24 to 28 h at 163 \u00b0C (325 \u00b0F) also may be used. (g) Aging of aluminum alloys 7050, 7075, 7175, and 7475 from any temper\nto the T73 or T76 temper series requires closer-than-normal controls on aging variables such as time, temperature, heatup rate, and so forth, for any given item. In addition, when material in a T6-type temper\nis reaged to a T73- or T76-type temper, the specific condition of the T6 material (such as property levels and other effects of processing variables) is extremely important and will affect the capability of the\nreaged material to conform to the requirements specified for the applicable T73- or T76-type temper. (h) Two-stage treatment comprised of 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by: 24 to 30 h at 163 \u00b0C (325 \u00b0F)\nfor sheet and plate: 8 to 10 h at 177 \u00b0C (350 \u00b0F) for rolled or cold finished rod and bar; 6 to 8 h at 177 \u00b0C (350 \u00b0F) for extrusions and tube; 8 to 10 h at 177 \u00b0C (350 \u00b0F) for forgings in the T73 temper; and 6\nto 8 h at 177 \u00b0C (350 \u00b0F) for forgings in the T7352 temper. (i) These heat treatments also apply to alclad sheet and plate of these alloys. (j) An alternative two-stage treatment comprised of 4 h at 96 \u00b0C (205 \u00b0F)\nfollowed by 8 h at 157 \u00b0C (315 \u00b0F) also may be used. (k) For sheet, plate, tube, and extrusions, an alternative two-stage treatment comprised for 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by 14 to 18 h at 168 \u00b0C\n(335 \u00b0F) may be used, provided that a heatup rate of approximately 14 \u00b0C/h (25 \u00b0F/h) is employed. For rolled or cold finished rod and bar, the alternative treatment is 10 h at 177 \u00b0C (350 \u00b0F). (I) An alternative\nthree-stage treatment comprised of 5 h at 99 \u00b0C (210 \u00b0F), 4 h at 121 \u00b0C (250 \u00b0F), and then 4 h at 149 \u00b0C (300 \u00b0F) may also be used. (m) Quenched in water at 60 to 80 \u00b0C (140 to 180 \u00b0F). (n) Stress relieved\nby 1 to 5% cold reduction after solution treatment and prior to precipitation heat treatments. (o) 7175-T74 and -T7452 heat treatments are directed to specific results, may vary from supplier to supplier and\nare either proprietary or patented. (p) Must be preceded by soak at 466 to 477 \u00b0C (870 to 890 \u00b0F). See U.S. Patent 3,791,880.\n848 / Heat Treating of Nonferrous Alloys\nTable 2 Soak times and maximum quench delays for solution treatment of wrought\naluminum alloys\nSee Table 1 for solution-treating temperatures.\nMaximum\nquench\ndelay, s\n7\nSoak time, minutes\nAir furnace(b)\nSalt bath(c)\nThickness(a), mm (in.)\nmin\nmax(d)\nmin\nmax(d)\n\u22640.41 (0.016)\n20\n25\n10\n15\n5\n0.51 (0.020)\n20\n30\n10\n20\n7\n0.64 (0.025)\n25\n35\n15\n7\n0.81 (0.032)\n25\n35\n15\n1.02 (0.040)\n30\n40\n20\n30\n10\n1.27 (0.050)\n30\n40\n20\n30\n10\n1.35 (0.053)\n30\n40\n20\n30\n10\n1.80 (0.071)\n35\n45\n25\n10\n2.03 (0.080)\n35\n45\n25\n10\n35\n45\n25\n10\n40\n55\n30\n15\n40\n55\n30\n15\n50\n60\n35\n15\n50\n60\n35\n15\n55\n65\n35\n15\n65\n75\n45\n55\n15\n+30\n+20\n30\n+20\n(e)\n2.29 (0.090)\n2.54 (0.100)\n3.18 (0.125)\n4.06 (0.160)\n4.57 (0.180)\n6.35 (0.250)\nFor each additional 12.7 (12) or fraction\nRivets (all)\n>6.35 (0.250)-12.7 (0.500)\n222222222222222328\n+30\n60\n28339999972288678\n22222222222222292\n2822288333444445\n5\n(a) Minimum dimension of thickest section. (b) Soak time begins when all pyrometer instruments recover to original operating\ntemperature. (c) Soak time begins at time of immersion except when a heavy charge causes bath temperature to drop below specified\nminimum, in which case soak time begins when bath regains minimum temperature. (d) Applicable to alclad materials only. (e) Increases\nin thickness above 12.7 mm (12 in.) do not affect maximum quench delay, which remains constant at 15 s.\nIn the tabulation above, note especially the\neffects of small increments of temperature,\nwithin the normal range, on the properties\nof 0.8 mm (0.032 in.) 2024-T4 sheet.\nSolution-Treating Time. The time at the\nnominal solution heat-treating temperature\n(soak time) required to effect a satisfactory\ndegree of solution of the undissolved or\nprecipitated soluble phase constituents and\nto achieve good homogeneity of the solid\nsolution is a function of microstructure be-\nfore heat treatment. This time requirement\ncan vary from less than a minute for thin\nsheet to as much as 20 h for large sand or\nplaster-mold castings. Guideline informa-\ntion for soak times required for wrought\nproducts of various section thicknesses is\ngiven in Table 2. Similar guidelines for\ncastings are presented in Table 3. The time\nrequired to heat a load to the treatment\ntemperature in furnace heat treatment also\nincreases with section thickness and fur-\nnace loading, and thus total cycle time\nincreases with these factors.\nSoak time for alclad sheet and for parts\nmade from alclad sheet must be held to a\nminimum, because excessive diffusion of\nalloying elements from the core into the\ncladding reduces corrosion protection. For\nthe same reason, reheat treatment of alclad\nsheet less than 0.75 mm (0.030 in.) thick\ngenerally is prohibited, and the number of\nreheat treatments permitted for thicker al-\nclad sheet is limited.\nThe soak times for wrought alloys take\ninto account the normal thermal lag be-\ntween furnace and part and the difference\nbetween surface and center temperatures\nfor commercial equipment qualified to the\nstandards of MIL-H-6088. The rapid heat-\ning rates of salt baths permit all immersion\n\u2022\ntime to be counted as soak time unless the\nbath temperature drops below the minimum\nof the range. Even then, soak time begins as\nsoon as the bath temperature returns to the\nminimum. In air furnaces, soak time does\nnot begin until all furnace instruments re-\nturn to their original set temperature-that\nis, the temperature reading before insertion\nof the load.\nIn air furnaces, thermocouples may also\nbe attracted to, or buried in, parts located in\nthe load in such a manner as to represent\nthe hottest and coldest temperatures in each\nzone. In this way, it is possible to ensure\nthat adequate soaking is obtained.\nSpecial consideration is given also to es-\ntablishing soak times for hand and die forg-\nings; soak time in some specifications is\nextended to complete solution and homog-\nenization in areas that received marginal\nreduction during forging. Considerable vari-\nation exists in the amount of soak time\nadded; some specifications call for an arbi-\ntrary addition, such as one hour, and others\nrequire one hour per inch of thickness of the\noriginal forging.\nIn air furnaces, careful attention should\nbe given to arrangement of the load. Air\nflow and natural temperature distribution\nwithin the furnace should be arranged to:\n\u2022\nOffer minimum resistance to air flow\nProduce the least disturbance in the natu-\nral temperature distribution\n\u2022 Afford constant replenishment of the en-\nvelope of air around each part\nIt is common practice to specify a minimum\nspacing of 50 mm (2 in.) between parts, but\nlarge complex shapes may require consider-\nably greater spacing. Many operators have\nfound conservative loading practices to be\nmore economical in the long run than heavi-\ner loading, because with lighter loads heat-\ning rates are higher and fewer rejections and\nservice failures are encountered.\nHigh-Temperature Oxidation. There is a\ncondition, commonly but erroneously\nknown as HTO or high-temperature oxida-\ntion, which can lead to deterioration of\nproperties in aluminum alloys. High-tem-\nperature oxidation is a misnamed condition\nof hydrogen diffusion that affects surface\nlayers during elevated-temperature treat-\nment. This condition can result from mois-\nture contamination in the furnace atmo-\nsphere and is sometimes aggravated by\nsulfur (as in heat-treatment furnaces also\nused for magnesium alloy castings) or other\nfurnace refractory contamination.\nMoisture in contact with aluminum at\nhigh temperatures serves as a source of\nnascent hydrogen, which diffuses into the\nmetal. Foreign materials, such as sulfur\ncompounds, function as decomposers of the\nnatural oxide surface film, eliminating it as a\nbarrier either between the moisture and the\naluminum or between the nascent hydrogen\nand the aluminum. The most common man-\nifestation of high-temperature oxidation is\nsurface blistering, but occasionally the only\nmanifestations are internal discontinuities\nor voids, which can be detected only by\ncareful ultrasonic inspection or by metallo-\ngraphic techniques.\nIt is important to recognize that the symp-\ntoms of high-temperature oxidation are\nidentical to those of unsoundness or high\ngas content in the original ingot or of other\nimproper mill practice. Blisters resulting\nfrom ingot defects, improper extrusion or\nimproper rolling may be lined up in the\ndirection of working. However, it usually is\nimpossible to distinguish among defect\nsources, and therefore the possibility that a\ncontaminated atmosphere is the cause of\nthe defects must be checked.\nNot all alloys and product forms are\nequally vulnerable to this type of attack.\nThe 7xxx series alloys are most susceptible,\nfollowed by the 2xxx alloys. Extrusions\nundoubtedly are the most susceptible form;\nforgings are probably second. Low-strength\nalloys and alclad sheet and plate are rela-\ntively immune to high-temperature oxida-\ntion. (Blistering of alclad material as a result\nof inadequate bonding is not the same as the\nblistering caused by high-temperature oxi-\ndation.)\nIf the protective oxide film formed during\nmill operations is removed from the mill\nproduct by a subsequent mechanical condi-\ntioning operation such as sanding, the con-\nditioned surface will be more susceptible to\nhigh-temperature oxidation than those from\nwhich the film was not removed.\nMoisture can be minimized by thoroughly\ndrying parts and racks before they are\ncharged. Drain holes often are needed in\nracks of tubular construction to avoid en-\nTable 3 Typical heat treatments for aluminum alloy sand and permanent mold castings\nSolution heat treatment(b)\nHeat Treating of Aluminum Alloys / 849\nAging treatment\nAlloy\nTemper\nType of\ncasting(a)\nTemperature(c)\nTemperature(c)\n\u00b0C\n\u00b0F\nTime, h\n\u00b0C\n\u00b0F\nTime, h\n201.0(d)\nT4\nS or P\n490-500(e)\n+525-530\n910-930(e)\n+980-990\n2\n14-20\nMinimum of 5 days at room temperature\nT6\nS\nT7\nS\n510-515(e)\n+525-530\n510-515(e)\n950-960(e)\n+980-990\n2\n14-20\n155\n+525-530\n950-960(e)\n+980-990\n2\n14-20\n190\n310\n370\n20\n5\nT43(f)\n525\n980\n20\nT71\n490-500(e)\n910-930(e)\n2\n+525-530\n+980-990\n14-20\n200\n204.0(d)\nT4\nS or P\n530\n985\n12\n24 h at room temperature + \u00bd to 1 h at 160 \u00b0C\n390\nMinimum of 5 days at room temperature\n4\nT4\nS or P\n520\n970\n10\nT6(g)\nS or P\n530\n985\n12\n(g)\n(g)\n206.0(d)\nT4\nS or P\n490-500(e)\n+525-530\n910-930(e)\n+980-990\n2\n14-20\nMinimum of 5 days at room temperature\nT6\nS or P\n490-500(e)\n+525-530\n910-930(e)\n+980-990\n2\n14-20\n155\n310\n12-24\nT7\nS or P\n490-500(e)\n910-930(e)\n2\n+525-530\n+980-990\n14-20\n200\n390\n4\nT72\nS or P\n490-500(e)\n910-930(e)\n2\n+525-530\n+980-990\n14-20\n243-248\n470-480\n208.0\nT55\n222.0\nO(h)\nT61\nT551\nSSSP\n155\n310\n16\n315\n600\n3\n510\n950\n12\n155\n310\n11\n170\n340\n16-22\nT65\n510\n950\n4-12\n170\n340\n7-9\n242.0\nO(i)\n345\n650\n3\nT571\n205\n400\n8\n165-170\n330-340\n22-26\nT77\nS\n515\n960\n5(j)\n330-355\n625-675\n2 (minimum)\nT61\nS or P\n515\n960\n4-12(j)\n205-230\n400-450\n3-5\n295.0\nT4\n515\n960\nT6\n515\n960\nT62\n515\n960\nT7\n515\n960\n296.0\nT4\n510\n950\nT6\n510\n950\nT7\n510\n950\n319.0\nT5\nT6\n505\n940\n505\n940\n328.0\nT6\n515\n960\n332.0\nT5\n333.0\nT5\nP\nT6\nP\n505\n950\nT7\nP\n505\n940\n336.0\nT551\nP\nT65\nP\n515\n960\n354.0\n(k)\n525-535\n980-995\n355.0\nT51\nS or P\nT6\nS\n525\n980\nP\n525\n980\nT62\n525\n980\nT7\n525\n980\n525\n980\nT71\n525\n980\n525\n980\nC355.0\nT6\n525\n980\nT61\n525\n980\n\u00a3sts\u00a3stts:E;g;sEs:\u0f68EsES\n12\n12\n155\n310\n3-6\n12\n155\n310\n12-24\n12\n260\n500\n46\n8\n...\n8\n155\n310\n1-8\n8\n260\n500\n46\n205\n400\n8\n12\n155\n310\n2-5\n4-12\n155\n310\n2-5\n12\n155\n310\n2-5\n205\n400\n7-9\n205\n400\n7-9\n6-12\n155\n310\n2-5\n6-12\n260\n500\n46\n205\n400\n7-9\n8\n205\n400\n7-9\n10-12\n(h)\n(h)\n(1)\n225\n440\n7-9\n12\n155\n310\n3-5\n4-12\n155\n310\n2-5\n4-12\n170\n340\n14-18\n12\n225\n440\n3-5\n4-12\n225\n440\n3-9\n12\n245\n475\n46\n4-12\n245\n475\n3-6\n12\n155\n310\n3-5\n6-12\nRoom temperature\n8 (minimum)\n155\n310\n10-12\n356.0\nT51\nT6\nT7\nT71\nA356.0\nT6\nT61\nSSPSPSesa\nS or P\n225\n440\n7-9\n540\n1000\n12\n155\n310\n3-5\n540\n1000\n4-12\n155\n310\n2-5\n540\n1000\n12\n205\n400\n3-5\n540\n1000\n4-12\n225\n440\n7-9\n540\n1000\n10-12\n245\n475\n3\n540\n1000\n4-12\n245\n475\n3-6\n540\n1000\n12\n155\n310\n3-5\n540\n1000\n6-12\nRoom temperature\n155\n310\n8 (minimum)\n6-12\n(continued)\n(a) S, sand; P, permanent mold. (b) Unless otherwise indicated, solution treating is followed by quenching in water at 65-100 \u00b0C (150-212 \u00b0F). (c) Except where ranges are given, listed temperatures are \u00b16 \u00b0C\nor \u00b110 \u00b0F. (d) Casting wall thickness, solidification rate, and grain refinement affect the solution heat-treatment cycle in alloys 201.0, 204.0, and 206.0, and care must be taken in approaching the final solution\ntemperature. Too rapid an approach can result in the occurrence of incipient melting. (e) For castings with thick or other slowly solidified sections, a pre-solution heat treatment ranging from about 490 to\n515 \u00b0C (910 to 960 \u00b0F) may be needed to avoid too rapid a temperature rise to the solution temperature and the melting of CuAl2. (f) Temper T43 for 201.0 was developed for improved impact resistance with\nsome decrease in other mechanical properties. Typical Charpy value is 20 J (15 ft lb). (g) The French precipitation treatment technology for the heat treatment of 204.0 alloy requires 12 h at temperature.\nThe aging temperatures of 140, 160, or 180 \u00b0C (285, 320, or 355 \u00b0F) are selected to meet the required combination of properties. (h) Stress relieve for dimensional stability as follows: hold 5 h at 413 \u00b1 14 \u00b0C\n(775 \u00b1 25 \u00b0F): furnace cool to 345 \u00b0C (650 \u00b0F) over a period of 2 h or more: furnace cool to 230 \u00b0C (450 \u00b0F) over a period of not more than 2 h; furnace cool to 120 \u00b0C (250 \u00b0F) over a period of approximately\n2 h; cool to room temperature in still air outside the furnace. (i) No quench required: cool in still air outside the furnace. (j) Air-blast quench from solution-treating temperature. (k) Casting process varies\n(sand, permanent mold, or composite) depending on desired mechanical properties. (1) Solution heat treat as indicated, then artificially age by heating uniformly at the temperature and for the time necessary\nto develop the desired mechanical properties. (m) Quench in water at 65-100 \u00b0C (150-212 \u00b0F) for 10-20 s only. (n) Cool to room temperature in still air outside the furnace.\n850 / Heat Treating of Nonferrous Alloys\nTable 3 (continued)\nSolution heat treatment(b)\nAging treatment\nTemperature(c)\nTemperature(c)\nAlloy\nTemper\nType of\ncasting(a)\n\u00b0C\n\u00b0F\nTime, h\n\u00b0C\n\u00b0F\nTime, h\n357.0\nT6\nP\nT61\nS\nA357.0\n359.0\nA444.0\nT4\nP\n520.0\nT4\nS\n535.0\nT5(h)\n705.0\nT5\nS\nandzannn\n540\n1000\n8\n175\n350\n6\n540\n1000\n10-12\n155\n310\n10-12\n(k)\n540\n1000\n8-12\n(h)\n(h)\n(h)\n(k)\n540\n1000\n10-14\n(h)\n(h)\n(h)\n540\n1000\n8-12\n430\n810\n18(m)\n400\n750\n5\nRoom temperature\n21 days\n100\n210\n8\nP\n:\nRoom temperature\n21 days\n100\n210\n10\n707.0\nT5\nSP\n155\n310\n3-5\n\u0420\nRoom temperature, or\n21 days\n100\n210\n8\nT7\n710.0\n711.0\nRE\nT5\nT1\n712.0\nT5\nSPSPS\n530\n530\n990\n990\n8-16\n175\n350\n4-10\n4-8\n175\n350\n4-10\nRoom temperature\n21 days\nRoom temperature\n21 days\n713.0\nT5\nS or P\nRoom temperature, or\n155\nRoom temperature, or\n120\n21 days\n315\n6-8\n21 days\n250\n16\n771.0\nT53(h)\nT5\nT51\nT52\nT6\nT71\n850.0\nT5\n851.0\nT5\nSSSSSSSS\n415(n)\n775(n)\n5(n)\n180(n)\n360(n)\n4(n)\n180(n)\n355(n)\n3-5(n)\n205\n405\n6\n(h)\n(h)\n(h)\n590(n)\n1090(n)\n6(n)\n130\n265\n3\n590(i)\n1090(i)\n6(i)\n140\n285\n15\nS or P\n220\n430\n7-9\nT6\nS or P\nP\n220\n430\n7-9\n480\n900\n6\n220\n430\n4\n852.0\nT5\nS or P\n220\n430\n7-9\n(a) S, sand; P, permanent mold. (b) Unless otherwise indicated, solution treating is followed by quenching in water at 65-100 \u00b0C (150-212 \u00b0F). (c) Except where ranges are given, listed temperatures are \u00b16 \u00b0C\nor \u00b110 \u00b0F. (d) Casting wall thickness, solidification rate, and grain refinement affect the solution heat-treatment cycle in alloys 201.0, 204.0, and 206.0, and care must be taken in approaching the final solution\ntemperature. Too rapid an approach can result in the occurrence of incipient melting. (e) For castings with thick or other slowly solidified sections, a pre-solution heat treatment ranging from about 490 to\n515 \u00b0C (910 to 960 \u00b0F) may be needed to avoid too rapid a temperature rise to the solution temperature and the melting of CuAl2. (f) Temper T43 for 201.0 was developed for improved impact resistance with\nsome decrease in other mechanical properties. Typical Charpy value is 20 J (15 ft lb). (g) The French precipitation treatment technology for the heat treatment of 204.0 alloy requires 12 h at temperature.\nThe aging temperatures of 140, 160, or 180 \u00b0C (285, 320, or 355 \u00b0F) are selected to meet the required combination of properties. (h) Stress relieve for dimensional stability as follows: hold 5 h at 413 \u00b1 14 \u00b0C\n(775 +25 \u00b0F); furnace cool to 345 \u00b0C (650 \u00b0F) over a period of 2 h or more; furnace cool to 230 \u00b0C (450 \u00b0F) over a period of not more than 2 h; furnace cool to 120 \u00b0C (250 \u00b0F) over a period of approximately\n2 h; cool to room temperature in still air outside the furnace. (i) No quench required; cool in still air outside the furnace. (j) Air-blast quench from solution-treating temperature. (k) Casting process varies\n(sand, permanent mold, or composite) depending on desired mechanical properties. (j) Solution heat treat as indicated, then artificially age by heating uniformly at the temperature and for the time necessary\nto develop the desired mechanical properties. (m) Quench in water at 65-100 \u00b0C (150-212 \u00b0F) for 10-20 s only. (n) Cool to room temperature in still air outside the furnace.\ntrapment of water. Another common re-\nquirement is adjustment of the position of\nthe quench tank with respect to furnace\ndoors and air intake. Because it is unlikely\nthat all moisture can be eliminated from the\natmosphere in a production heat-treating\nfurnace, it is extremely important to elimi-\nnate all traces of other contaminants from\nboth the parts and the furnace atmosphere.\nThe most virulent contaminants in attack-\ning aluminum are sulfur compounds. Resi-\ndues from forming or machining lubricants, or\nfrom a sulfur dioxide protective atmosphere\nused in prior heat treatment of magnesium,\nare potential sources of sulfur contamination.\nIn one plant, surface contamination resulted\nfrom sulfur-containing materials in tote boxes\nused to transport parts. In another, an epi-\ndemic of blistering was cured by rectifying a\n\"sour\" degreaser. In a third instance, it was\nfound that a vapor-degreasing operation was\nnot completely removing a thin, hard waxy\nresidue, and an alkaline cleaning operation\nwas added.\nVery often, the source of contamination\nis obscure and difficult to detect, and the\nproblem must be combated in another way.\nThe most common of the alternative meth-\nods is use of a protective fluoborate com-\npound in the furnace. Such a compound\nusually is effective in minimizing the harm-\nful effects of moisture and other undesirable\ncontaminants because it forms a barrier\nlayer or film on the aluminum surface. The\nadditive is not a universal solution; in some\napplications, high-temperature oxidation\nhas occurred even though a fluoborate com-\npound was employed. Also, the use of such\ncompounds, particularly ammonium fluo-\nborate, may present a hazard to personnel if\nused in poorly sealed furnaces or in furnac-\nes that discharge their atmospheres into\nenclosed areas.\nProtective fluoborate compounds accen-\ntuate staining or darkening of the parts\nbeing treated. (At times, this attack, partic-\nularly on parts located near the protective-\ncompound container during heat treatment,\nhas been severe enough to be termed \"cor-\nrosion.\") Although this minor nuisance\nmight be considered a small price to pay for\nsolution of a problem of high-temperature\noxidation, the residual compound in the\nfurnace dissipates slowly. Therefore, subse-\nquent loads of alloys and product forms\nwhose end uses require bright surfaces, and\nthat are not susceptible to high-temperature\noxidation, may be detrimentally affected.\nSuccessful use of fluoborate protective\ncompounds appears to depend on specify-\ning the right amount for each furnace; this\nmust be established on a trial-and-error\nbasis. One aircraft manufacturer adds 4\ng/m\u00b3 (0.004 oz/ft\u00b3) of furnace chamber to\neach load. Another adds 0.45 kg (1 lb) per\nshift to a metal container hung on the fur-\nnace chamber wall, thus avoiding loss of the\ncompound during quenching.\nA second method of combating high-tem-\nperature oxidation is to anodize the work\nbefore it is heat treated. The resultant alu-\nminum oxide film prevents attack by con-\ntaminants in the furnace atmosphere. The\nonly deterrents to the use of anodizing are\nits cost (in money and time) and the slight\nsurface frostiness which results from the\nsubsequent stripping operation.\nThe usual objection to the blistered sur-\nface produced by high-temperature oxida-\ntion is its unsightly appearance. This often\ncan be improved (for salvage purposes) by\napplying local pressure to flatten each blis-\nter and then finishing by a mechanical pro-\ncess such as polishing, buffing, sanding, or\n\n\nabrasive blasting. In general, the effect of\nHTO on static properties and fatigue\nstrength is slight. However, if a void result-\ning from HTO is located close to another\nstress concentration, such as a hole, much\ngreater degradation of fatigue strength is\nlikely. In critical aluminum alloy forgings,\nany blistering must be evaluated carefully\nfor its effect on the integrity of the part. Any\n\"cosmetic\" salvage should be performed\nonly after it has been established that the\nblisters are superficial and will not remain in\nthe finished product.\nPrecipitation Heat Treating without Prior\nSolution Heat Treatment. Certain alloys that\nare relatively insensitive to cooling rate\nduring quenching can be either air cooled or\nwater quenched directly from a final hot-\nworking operation. In either condition,\nthese alloys respond strongly to precipita-\ntion heat treatment. This practice is widely\nused in producing thin extruded shapes of\nalloys 6061, 6063, 6463, and 7005. Upon\nprecipitation heat treating after quenching\nat the extrusion press, these alloys develop\nstrengths nearly equal to those obtained by\nadding a separate solution heat treating op-\neration. Changes in properties occurring\nduring the precipitation treatment follow\nthe principles outlined in the discussion of\nsolution heat-treated alloys.\nQuenching\nQuenching is in many ways the most\ncritical step in the sequence of heat-treating\noperations. The objective of quenching is to\npreserve the solid solution formed at the\nsolution heat-treating temperature, by rap-\nidly cooling to some lower temperature,\nusually near room temperature. From the\npreceding general discussion, this statement\napplies not only to retaining solute atoms in\nsolution, but also to maintaining a certain\nminimum number of vacant lattice sites to\nassist in promoting the low-temperature dif-\nfusion required for zone formation. The\nsolute atoms that precipitate either on grain\nboundaries, dispersoids, or other particles,\nas well as the vacancies that migrate (with\nextreme rapidity) to disordered regions, are\nirretrievably lost for practical purposes and\nfail to contribute to the subsequent\nstrengthening.\nIn most instances, to avoid those types of\nprecipitation that are detrimental to me-\nchanical properties or to corrosion resis-\ntance, the solid solution formed during so-\nlution heat treatment must be quenched\nrapidly enough (and without interruption) to\nproduce a supersaturated solution at room\ntemperature the optimum condition for\nprecipitation hardening. The resistance to\nstress-corrosion cracking of certain copper-\nfree aluminum-zinc-magnesium alloys,\nhowever, is improved by slow quenching.\nMost frequently, parts are quenched by\nimmersion in cold water or, in continuous\nheat treating of sheet, plate, or extrusions in\nprimary fabricating mills, by progressive\nflooding or high-velocity spraying with cold\nwater. However, parts of complex shape,\noften with both thin and thick sections (such\nas die forgings, most castings, impact extru-\nsions, and components formed from sheet)\nare commonly quenched in a medium that\nprovides somewhat slower cooling. This\nmedium may be water at 65 to 80 \u00b0C (150 to\n180 \u00b0F), boiling water, an aqueous solution\nof polyalkylene glycol, or some other fluid\nmedium such as forced air or mist.\nIf appreciable precipitation during cool-\ning is to be avoided, two requirements must\nbe satisfied. First, the time required for\ntransfer of the load from the furnace to the\nquenching medium must be short enough to\npreclude slow precooling into the tempera-\nture range where very rapid precipitation\ntakes place. For alloy 7075, this range was\ndetermined to be 400 to 290 \u00b0C (750 to 550\n\u00b0F), and some sources quote this range (or a\nslightly different range) as the most critical\nrange for quenching of any aluminum alloy.\nLater work has shown that the most critical\nrange is alloy-dependent, and as will be\ndiscussed in detail under \"Quench-Factor\nAnalysis,\u201d significant errors can result from\nthe assumption that precipitation is negligi-\nble outside of a so-called \"critical range.'\nThe second requirement for avoidance of\nappreciable precipitation during quenching\nis that the volume, heat-absorption capaci-\nty, and rate of flow of the quenching medi-\num be such that little or no precipitation\noccurs during cooling. Any interruption of\nthe quench that might allow reheating into a\ntemperature range where rapid precipitation\ncan occur must be prohibited.\nFor maximum dimensional stability,\nsome forgings and castings are fan cooled or\nstill-air cooled. In such instances, precipita-\ntion-hardening response is limited, but sat-\nisfactory values of strength and hardness\nare obtained. Extrusions produced without\nseparate solution heat treatment can be air\nor mist quenched, but thicker sections may\nrequire water quenching by immersion or\nspraying. Alloys that are relatively dilute,\nsuch as 6063 and 7005, are particularly well\nsuited to air quenching, and their mechani-\ncal properties are not greatly affected by its\nlow cooling rate. Lower quenching rates are\nalso employed for forgings, castings, and\ncomplex shapes to minimize warpage or\nother distortion and the magnitude of resid-\nual stresses developed as a consequence of\ntemperature nonuniformity from surface to\ninterior.\nEffect of Quench Rate on Properties. As a\nbroad generalization, the highest strengths\nattainable and the best combinations of\nstrength and toughness are those associated\nwith the most rapid quenching rates. Resis-\ntance to corrosion and stress-corrosion\ncracking are other characteristics that are\ngenerally improved by maximum rapidity of\nquenching. Some of the alloys used in arti-\nHeat Treating of Aluminum Alloys / 851\nficially aged tempers, and in particular the\ncopper-free 7xxx alloys, are exceptions to\nthis rule. The effect of quench rate on\nmechanical properties may also depend on\nthe desired temper. In the underaged con-\ndition, for example, a slow quench rate is\nmore detrimental on ductility and fracture\ntoughness. Strength would be more affected\nafter near-to-peak aging.\nBecause of these effects, much work has\nbeen done over the years to understand and\npredict how quenching conditions and prod-\nuct form influence properties. The relative\neffects of quench methods can be compared\nin terms of average quench rates. In Fig 3,\nfor example, the effects of quenching on the\nyield strength of four alloys are compared in\nterms of average quenching rates through\nthe range from 400 to 290 \u00b0C (750 to 550 \u00b0F).\nFor alloys relatively high in sensitivity to\nquenching rate, such as 7075, rates of about\n300 \u00b0C/s (540 \u00b0F/s) or higher are required in\norder to obtain near-maximum strength af-\nter precipitation heat treatment. The other\nalloys in Fig 3 maintain their strengths at\ncooling rates as low as about 100 \u00b0C/s (180\n\u00b0F/s). Similar comparisons in terms of aver-\nage quench rates are shown in Tables 4 and\n5.\nAverage quench rates are useful in com-\nparing experimental results from various\nquench methods. In Table 4, for example, a\nsevere reduction in strength occurred at the\naverage quench rate of 36 \u00b0C/s (65 \u00b0F/s).\nHowever, average quench rates only com-\npare results in a \"critical\" temperature\nrange, where precipitation is most likely to\noccur. This method is not entirely accurate,\nbecause significant precipitation can also\noccur outside the specified critical temper-\nature range of average quench rates. More-\nover, for high-strength alloys, toughness\nand corrosion resistance may be impaired\nwithout significant loss of tensile strength.\nTherefore, a more sophisticated compar-\nison, known as quench-factor analysis, is\nneeded for quantitative property prediction\nor property optimization. Quench-factor\nanalysis, as discussed in a later section, is\nuseful when cooling rates are nonuniform.\nDelay in Quenching. Whether the transfer\nof parts from the furnace to the quench is\nperformed manually or mechanically, it\nmust be completed in less than the specified\nmaximum time. The maximum allowable\ntransfer time or \u201cquench delay\" varies with\nthe temperature and velocity of the ambient\nair and the mass and emissivity of the parts.\nFrom cooling curves such as those illustrat-\ned in Fig 4, maximum quench delays (see\ntable accompanying Fig 4) can be deter-\nmined that will ensure complete immersion\nbefore the parts cool below 400 \u00b0C (750 \u00b0F).\nMIL-H-6088 specifies maximum quench de-\nlays for high-strength alloys of 5, 7, 10, and\n15 s for thickness ranges of up to 0.016 in.\n(0.41 mm), 0.017 to 0.031 in. (0.43 to 0.79\nmm), 0.032 to 0.090 in. (0.81 to 2.29 mm),\n852 / Heat Treating of Nonferrous Alloys\nYield strength, MPa\n(a)\nFraction of WQ yield strength, %\n(c)\nAverage quenching rate\nfrom 750-550 \u00b0F, \u00b0F/s\n10\n600\n500 -7050-T736\n400\n300\n200\n100\n102\n103\n104\n7050-T6\n80\n60\n7050-T73\n\u30ad\u30fc\u30ad\n40\n6061-T6\n-2024-T4\n20\n1\n10\n102\n103\n104\nAverage quenching rate\nfrom 400-290 \u00b0C, \u00b0C/s\n100\nFurnace cooling\n80\n60\n50\n40\n20\n20\n20\nYield strength, ksi\nTensile strength, MPa\n(b)\nAverage cooling rate\nfrom 750-550 \u00b0F, \u00b0F/s\n10\n102 103 104 105\n700\n100\n600\n500\n7178-T6\n7075-T6\n7050-T73\n80\n7075-T73\n2014-T6\n2024-T4\n6070-T6\n1\n6061-T6\n60\n40\n400\n300\n200\n1\n10\nAverage cooling rate\nfrom 400-290 \u00b0C, \u00b0C/s\n102 103 104 105\nAir cooled\nForced air cooling\nAlloy and condition (Source: Ref 1)\nO 8090, peak aged\n\u26ab 2090, peak aged\n\u25b2 7150, aged 24 h at 120 \u00b0C\n\u25b2 7475, aged 24 h at 120 \u00b0C\n\u043d\u043d\n0\n0.01\n0.1\n1\n10\nAverage cooling rate, \u00b0C/s\nWQ\nTensile strength, ksi\nFig 3 Quench sensitivity of various aluminum alloys as a function of average quench rates. (a) Yield strength\nafter aging of four wrought alloys. (b) Tensile strength after aging of eight wrought alloys. (c) Relative\nquench sensitivity of two aluminum-lithium alloys (2090 and 8090, both solution treated for 1 h at 520 \u00b0C, or 970\n\u00b0F) and two Zn-Mg-Cu aluminum alloys (7150 and 7475, both solution treated for 40 min at 480 \u00b0C, or 895 \u00b0F)\nand over 0.090 in., respectively. Quench de-\nlay is conservatively defined as commencing\n\"when the furnace door begins to open or the\nfirst corner of a load emerges from a salt\nbath\" and ending \"when the last corner of the\nload is immersed in the water quench tank.\"\nRecommended maximum quench-delay times\nare listed in Table 2. However, exceeding the\nmaximum delay time is permitted if tempera-\nture measurements of the load prove that all\nparts are above 415 \u00b0C (775 \u00b0F) when\nquenched. The C-curves used in quench-fac-\ntor analysis can also assist in determining a\nmaximum allowable delay.\nIt is relatively easy to control quench\ndelay in day-to-day operations by using a\nTable 4 Effect of average quench rate on tensile properties of aluminum-lithium alloy\n2090\nAverage quench rate at center of plate\n0.5 \u00b0C/s (13 mm plate, air\ncooled)\n36 \u00b0C/s (38 mm plate, quenched\nin room-temperature water)\n46 \u00b0C/s (13 mm plate, quenched\nin boiling water)\n48 \u00b0C/s (13 mm plate, quenched\nin room-temperature water)\n85 \u00b0C/s (13 mm plate, quenched\nin ice brine)\n(a) Data are averages from 4 specimens.\nElongation(a), %\nYield\nstrength(a)\nTensile\nstrength(a)\nCondition\nMPa\nksi\nMPa\nksi\nAs-quenched\n162\n334\n2\n6% stretch + aged\n448\n513\n5\n8 h at 190 \u00b0C\nAs-quenched\n128\n312\n12\n6% stretch + aged\n338\n476\n6\n8 h at 190 \u00b0C\nAs-quenched\n138\n331\n16\n6% stretch aged\n530\n570\n9\n8 h at 190 \u00b0C\nAs-quenched\n139\n331\n17\n6% stretch aged\n526\n570\n7\n8 h at 190 \u00b0C\nAs-quenched\n135\n349\n19\n6% stretch aged\n8 h at 190 \u00b0C\n535\n575\n7\nstopwatch or, if necessary, by attaching\nthermocouples to parts. However, although\nthe cooling rate between 400 and 260 \u00b0C (750\nand 500 \u00b0F) is most critical and must be\nextremely high for many high-strength al-\nloys, it cannot be directly measured in pro-\nduction operations. It is usual to rely on\nstandardized practices, augmented by re-\nsults of tension tests and tests of suscepti-\nbility to intergranular corrosion.\nWater-immersion quenching normally is\ncontrolled in practice by stipulating maxi-\nmum quench-delay time and maximum wa-\nter temperature. The first requirement con-\ntrols the cooling rate during transfer and,\nfor high-strength alloys, often is based on\nthe criterion of complete immersion before\nthe metal cools below 415 \u00b0C (775 \u00b0F). This\nspecification of 415 \u00b0C (775 \u00b0F) is based on a\ncritical temperature for alloy 7075, which\nhas one of the more severe C-curves (Fig 5).\nTherefore, the criterion for complete im-\nmersion of other alloys might be based on a\ntemperature lower than the 415 \u00b0C (775 \u00b0F)\nspecification, depending on the characteris-\ntics of the particular C-curve.\nThe second requirement controls the\ncooling rate during immersion. MIL-H-6088\nspecifies that for water-immersion quench-\ning, except quenching of forgings and cast-\nings, the temperature of the water shall not\nexceed 38 \u00b0C (100 \u00b0F) upon completion of\nquenching. This requirement controls both\nthe temperature of the quench water prior\nto immersion and the ratio of the combined\nmass of load and rack to the volume of\nwater. However, to ensure adequate\nquenching effectiveness, it is necessary also\nthat the cooling fluid flow past all surfaces\nof each part during the first few seconds\nafter immersion. Before parts enter the fur-\nnace, their placement in racks or baskets\nshould be compatible with this requirement.\nDuring the first few seconds of quenching,\nagitation of the parts or the water should be\nsufficient to prevent local increases in tem-\nperature due to the formation of steam\npockets.\nIn one application, it was found that\n2024-T4 plates 13 by 760 by 760 mm (1/2 by\n30 by 30 in.), quenched singly into a large\nvolume of still water, were quite susceptible\nto intergranular corrosion. This susceptibil-\nity disappeared completely when the\nquenching practice was modified by adding\nsufficient agitation to break up the insulat-\ning blanket of steam that formed on the\nsurface of the hot metal. Quenching prac-\ntices for small parts such as fasteners and\nhydraulic fittings have been modified for the\nsame reason. Dumping in bulk from baskets\nhas been replaced by methods, such as the\nuse of shaker hearth furnaces or special\nracking, which permit parts to be quenched\nsingly.\nSpray Quenching. For spray quenching,\nthe quench rate is controlled by the velocity\nof the water and by volume of water per unit\nHeat Treating of Aluminum Alloys / 853\nTable 5 The effect of quench rate on the mechanical properties of age-hardened aluminum-lithium alloy 8090\nAlloy composition\nAl-2.28Li-0.86Cu-\n0.90Mg-0.13Zr-\n0.13Fe-0.06Si\nCooling from\nsolution treatment(a)\nYield strength(b)\nUltimate tensile\nstrength(b)\nStretch, %\nAging treatment\nMPa\nksi\nMPa\nksi\nElongation in 50 mm\n(2 in.)(b), %\nAir cool\n2\n190 \u00b0C for 16 h\n380\n55\n446\n64.5\n7.7\n(-0.25 \u00b0C/s)\n4\n170 \u00b0C for 24 h\n401\nPolymer quench\n(-18 \u00b0C/s)\n2\n190 \u00b0C for 16 h\n415\n4\n170 \u00b0C for 24 h\n415\nWater quench\n2\n190 \u00b0C for 16 h\n428\n(~120 \u00b0C)\n4\n170 \u00b0C for 24 h\n417\nAl-2.58Li-1.36Cu-\n0.89Mg-0.13Zr-\n0.17Fe-0.04Si\nAir cool\n2\n190 \u00b0C for 16 h\n417\n88 28 80\n58\n465\n67.5\n6.0\n60\n481\n70\n8.0\n60\n481\n70\n7.2\n62\n492\n71.4\n8.1\n60\n483\n70\n7.5\n60\n485\n70.3\n6.5\n(-0.25 \u00b0C/s)\n4\n170 \u00b0C for 24 h\n442\nPolymer quench\n(-18 \u00b0C/s)\n2\n190 \u00b0C for 16 h\n448\n170 \u00b0C for 24 h\n448\nWater quench\n(~120 \u00b0C/s)\n2\n190 \u00b0C for 16 h\n464\n170 \u00b0C for 24 h\n448\n2 22 28\n64\n503\n73\n4.5\n65\n524\n76\n65\n519\n75\n5.0\n67\n535\n77.5\n23 24\n6.8\n8.2\n65\n517\n75\n6.3\n(a) Solution treatment of 550 \u00b0C (1020 \u00b0F) for 1 h. (b) Data are averages from two specimens.\narea per unit time of impingement of the\nwater on the workpiece. Rate of travel of\nthe workpiece through the sprays is an\nimportant variable.\nLocal increases in temperature that occur\nwithin the first few seconds of quenching,\ncaused by a phenomenon such as plugged\nspray nozzles, are particularly deleterious.\nThe remaining \u201cinternal heat\u201d may be suf-\nficient to reheat the surface region. When\nthis happens, a large loss in strength occurs\nat the previously quenched surface. The\nloss of strength in the affected area of a\nheavy part is much more severe than that\ncaused by an inadequate quenching rate\nalone. This is illustrated for 75 mm (3 in.)\nthick 7075-T62 plate in Fig 6, which com-\npares, at various depths, the properties of a\nplate for which quenching was interrupted\non one side after 3 s with those of a plate\nthat was quenched from one side only.\nQuench Severity and Quenchant Selec-\ntion. Quench severity is commonly ex-\npressed in terms of an H-value (or Gross-\nmann number), where the H-value is related\nto the thermal conductivity (k) of the part(s)\nand the coefficient of heat transfer (C) be-\ntween the quenchant and the part. These\nquantities are related by the equation H =\nC/2k, where the coefficient of heat transfer\n(C) is affected by the quenchant velocity at\nthe surface of the part and several inherent\ncharacteristics of the quenchant (such as\nquenchant boiling point, viscosity, density,\nthermal conductivity, and specific heat).\nWater, which is the most widely used and\neffective quenching medium, can obtain\ncooling rates up to about 200 \u00b0C/s (400 \u00b0F/s)\nat the midplane of 25 mm (1 in.) thick\naluminum alloy plate (see the dashed line in\nFig 7). No rates higher than those defined\nby this line have been observed, although\nrates approaching them were measured with\nimpinging spray quenches. Lower cooling\nrates are achieved by immersion in heated\nwater (Fig 7) or by reducing the velocity of\nthe quenchant around the part (Table 6).\nCooling rates can also be reduced by low-\nering surface tension or by increasing the\nstability of the vapor film around the part.\nPolymer quenchants, which retard cool-\ning rates by the formation of films around\nthe part, are compared with water in Table\n6. The effective film coefficient is essential-\nly the heat transfer coefficient (C), which is\nrelated to the Grossmann number (H). The\napplication of polymer quenchants is cov-\nered in AMS specifications 3025 and 2770,\nalthough many aluminum and aerospace\ncompanies have developed internal specifi-\ncations that differ from AMS-2770. Typical\nparameters for quenching wrought products\n(other than forgings) in glycol-water solu-\ntions are presented in Table 7.\nOther Factors Affecting Quench Rate.\nQuenching rates are very sensitive to the\nsurface condition of the parts. Lowest rates\nare observed with products having freshly\nmachined or bright-etched, clean surfaces,\nor products that have been coated with\nmaterials that decrease heat transfer. The\npresence of oxide films or stains increases\ncooling rates. Further marked changes can\nbe effected through the application of non-\nreflective coatings, which also accelerate\nheating (Fig 8). Surface roughness exerts a\nsimilar effect; this appears related to vapor\nfilm stability. The manner in which complex\nproducts, such as engineered castings and\ndie forgings, enter the quenching medium\ncan significantly alter the relative cooling\nTemperature, \u00b0C\nTime per unit thickness, s/in\n0\n600\n1000\n2000\n3000\n1000\n800\n400\nAlclad\n600\n200\nNonclad\n400\nTemperature, \u00b0F\nmm\n0.41\n0.51\n0.64\n0.81\n200\n1.02\n0\n0\n50\n100\n150\nThickness\nMaximum quench delay, s\nin.\nAlclad\nNonclad\n0.016\n6.4\n4.4\n0.020\n8.0\n5.5\n0.025\n10.0\n6.8\n0.032\n12.8\n8.8\n0.040\n20.0\n11.0\nTime per unit thickness, s/mm\nFig 4 Cooling curves for alclad and nonclad aluminum products cooled from 495 \u00b0C (920 \u00b0F) in forced air. Air temperature, 25 \u00b0C (80 \u00b0F); air velocity, 2.3 m/s (450 ft/min).\nshown.\nTabulated values of quench delay (maximum delay before the material being quenched has cooled below 400 \u00b0C, or 750 \u00b0F) were determined from cooling curves\n854 / Heat Treating of Nonferrous Alloys\nTemperature, \u00b0C\nFig 5\n600\n500\n400\nA\n300\n200\nA: 7075\nB: 2017\n100\nC: 6061\nD: 6063\n0\n1\n10\n100\nTime, s\n103\n1110\n930\n750\n570\n390\n212\n32\nTemperature, \u00b0F\nTime-temperature-property curves at 95% of maximum tensile stress for various alloys. See the section\n\"Quench-Factor Analysis\" for discussion. Source: Ref 2\nrates at various points, thereby affecting\nmechanical properties and residual stresses\nestablished during quenching. Similarly,\nquenching complex extruded shapes whose\nwall thicknesses differ widely poses special\nproblems if distortion and stresses are to be\nminimized. In batch heat-treating opera-\ntions, placement and spacing of parts on the\nracks can be a major factor in determining\nthe quenching rates. In immersion quench-\ning, adequate volumes of the quenching\nmedium must be provided to prevent an\nexcessive temperature rise in the medium.\nWhen jet agitation is used to induce water\nTable 6 Grossmann numbers and heat transfer coefficients (C) of quenchant-to-part films\nQuenchant\nEffective film heat transfer\ncoefficient (C)\nBtu/ft2 h \u00b0F\n2460\n3105\nTemperature\nVelocity\n\u0422\u0443\u0440\u0435\n\u00b0C\n\u00b0F\nm/s\nft/min\nGrossmann\nNumber\n(H = C/2k)\nW/cm\u00b2. K\nWater\n27\n80\n0.00\n0\n1.07\n3.55\n0.25\n50\n1.35\n4.78\n0.50\n100\n1.55\n5.14\n3565\nWater\n38\n100\n0.00\n0\n0.99\n3.28\n2275\n0.25\n50\n1.21\n4.01\n2785\n0.50\n100\n1.48\n4.91\n3400\nWater\n49\n120\n0.00\n0\n1.10\n3.65\n2530\n0.25\n50\n1.29\n4.29\n2970\n0.50\n100\n1.60\n5.31\n3680\nWater\n60\n140\n0.00\n0\n0.86\n2.85\n1980\n0.25\n50\n1.09\n3.62\n2510\n0.50\n100\n1.33\n4.41\n3060\nWater\n71\n160\n0.00\n0\n0.21\n0.70\n485\n0.25\n50\n0.57\n1.89\n1310\n0.50\n100\n0.79\n2.62\n1815\n0.00\n0\n0.11\n0.36\n255\n0.25\n50\n0.21\n0.69\n485\n0.50\n100\n0.27\n0.89\n620\nWater\n93\n200\n0.00\n0\n0.06\n0.20\n138\n0.25\n50\n0.08\n0.27\n0.50\n100\n0.09\n0.30\n207\n100\n212\n0.00\n0.04\n0.13\n0.25\n50\n0.04\n0.13\n92\n0.50\n100\n0.04\n0.13\n92\nPolyalkylene glycol\n(UCON A)(a)\n30\n85\n0.00\n0.19\n0.63\n429\n0.25\n50\n0.21\n0.70\n475\n0.50\n100\n0.23\n0.77\n529\nPolyvinyl\npyrrolidone\n(PVP90)(a)\n30\n85\n0.00\n0\n0.44\n1.49\n0.25\n50\n0.40\n1.34\n0.50\n100\n0.42\n1.41\nWater\n82\n180\nWater\n80\n88088\n088\n184\n92\n1012\n912\n966\n(a) Polymer quenchants with concentrations of 25%. K is equal to the thermal conductivity of the aluminum alloy (7075). Source: Ref 4\nflow between parts, jets should not impinge\ndirectly and cause rapid localized cooling.\nQuenching to Minimize Residual Stress\nand Warpage. Although cold-water immer-\nsion or flushing is most common, because it\nproduces the most effective quench (and\nhas been required by MIL-H-6088 for 2014,\n2017, 2024, 2117, 7075, and 7178 alloys\nexcept forgings), it presents problems in-\nvolving residual stress and warpage.\nResidual stresses in heavy sections of\naluminum alloys originate from differential\nthermal expansion during quenching \u2015that\nis, the still-warm central material contracts,\npulling in the already cooled outer shell.\nThe magnitude of stresses increases with\nsection size, as shown in Fig 9.\nThe distribution pattern of residual\nstresses in as-quenched parts (compression\nin the outer layers and tension in the central\nportion) is usually desirable in service.\nCompressive stresses inhibit failure by fa-\ntigue and stress corrosion-two mecha-\nnisms that initiate in the outer fibers. Un-\nfortunately, metal-removal operations\nrequired after heat treating often expose\nmaterial that is stressed in tension. Also,\nmetal-removal operations that are asym-\nmetrical (with respect to residual stresses)\ncause distortion by redistributing residual\nstresses. When close-tolerance parts are\nbeing fabricated, the resulting warpage can\nbe costly and difficult to correct.\nAlthough service performance is some-\ntimes a factor, the major incentive for re-\nducing residual stress differentials has been\na reduction in warpage during machining or\nan improvement in shape before machining.\nOne approach to reducing the cooling-\nrate differential between surface and center\nis the use of a milder quenching medium-\nwater that is hotter than that normally used\nor water-glycol solutions. Boiling water,\nwhich is the slowest quenching medium\nused for thick sections, is sometimes em-\nployed for quenching wrought products\neven though it lowers mechanical properties\nand corrosion resistance. Quenching of\ncastings in boiling water, however, is stan-\ndard practice, and is reflected in design\nallowables.\nAnother approach to the minimization of\nresidual stresses that is generally successful\nconsists of rough machining to within 3.2\nmm (0.125 in.) or less of finish dimensions,\nheat treating, and then finish machining.\nThis procedure is intended to reduce the\ncooling-rate differential between surface\nand center by reducing thickness; other\nbenefits that accrue if this technique is used\nto reduce or reverse surface tension stress-\nes in finished parts are improvements in\nstrength, fatigue life, corrosion resistance,\nand reduced probability of stress-corrosion\ncracking.\nSeveral factors (especially quenching\nwarpage) sometimes preclude general use of\nthis procedure. The thinner and less sym-\nFig 6\nLongitudinal tensile strength, MPa\nLongitudinal yield strength, MPa\nHardness, HRB\n90\n70\nDepth, in.\n0\n0.5\n1.0\n1.5\n2.0\n2.5\n3.0\n60 O Control specimen\nQuenched from side A only\nA Quenched from side B, interrupted after 3 s\n50\n0\n10\n20\n30\nSide A\n0\n0.5\n600\n500\n400\n300\n0\nSide A\nb\n40\nDepth, mm\nDepth, in.\n1.0\n1.5\n10\n20\n30\n40\nDepth, mm\nDepth, in.\n0\n0.5\n1.0\n1.5\n550\nT\n450\n350\n250\n\u043e\n50\n60\n70\n80\nSide B\n2.0\n2.5\n3.0\n18\n50\n60\n70\nB\n80\n70\n60\n-50\n80\n788\nSide B\n2.0\n2.5\n3.0\nT\n70\n60\n50\n50\nLongitudinal tensile strength, ksi\n30\nLongitudinal yield strength, ksi\n40\n40\n30\n150\n0\n10\n20\n30\n40\n50\n60\n70\n80\nSide A\nDepth, mm\nSide B\nThrough-thickness property variations due to quench rate and temperature-rise effects in 7075-T62 plate\n75 mm (3 in.) thick\nmetrical a section, the more it will warp\nduring quenching, and the residual stresses\nresulting from straightening of warped parts\n(plus straightening costs) often are less de-\nsirable than the quenching stresses. Holding\nfixtures and die quenching may be helpful,\nbut precautions must be taken to ensure\nthat they do not retard quenching rates\nexcessively. Other factors that must be con-\nsidered are the availability of heat-treating\nfacilities and whether or not the advantages\nof such a manufacturing sequence offset the\ndelay and cost entailed in a double-machin-\ning setup.\nWarpage of thin sections during quench-\ning is also a problem. Even in the same\nload, symmetry of cooling usually varies\nsignificantly among identical parts and the\nresultant inconsistent warpage usually re-\nquires costly hand straightening. Conse-\nquently, a significant amount of effort has\nbeen devoted to reducing or eliminating\nwarpage by changing racking positions to\nachieve symmetry of cooling.\nFor sheet-metal parts, one manufacturer\nuses a double screen floor in the quenching\nrack to reduce the force of initial contact\nbetween water and parts. Others allow parts\nto \"free fall\" from rack to quench tank.\nSpacing and positioning on the rack are\ncarefully controlled so that parts will enter\nthe water with minimum impact. With this\ntechnique, water turbulences must be\navoided, because it will often cause parts to\nfloat for a few seconds, greatly reducing\ntheir cooling rate.\nHeat Treating of Aluminum Alloys / 855\nBecause of the difficulties encountered\nwith quenching in cold water, milder quen-\nchants have been employed. Indiscriminate\nuse of milder quenchants can have cata-\nstrophic effects; however, when their use is\nbased on sound engineering judgment and a\nmetallurgical knowledge of the effects on\nthe specific alloy, significant cost savings or\nperformance improvements can be realized.\nThe most frequent advantage is the re-\nduction in costly straightening operations\nand in resultant uncontrolled residual\nstresses. For example, one aircraft manu-\nfacturer utilizes water-spray and air-blast\nquenching for weldments and complex\nformed parts made from 6061, an alloy\nwhose corrosion resistance is insensitive to\nquenching rate. Straightening requirements\nare negligible and, through careful control\nof racking and coolant flow, the decrease in\nmechanical properties is minimized, as\nshown by the data in Fig 10.\nAnother development for reducing straight-\nening costs is quenching in water-polymer\nsolutions. Quenching of formed sheet-metal\nparts in aqueous solutions of polyalkylene\nglycol or in similar inversely soluble media\nhas significantly reduced the cost of straight-\nening these parts after quenching. The SAE\nheat-treatment specification AMS-2770 rec-\nommends, for several alloys, maximum thick-\nnesses that can be quenched in solutions of\nspecific concentrations while maintaining ac-\nceptable property levels. Typical parameters\nfor quenching wrought products (other than\nforgings) in glycol-water solutions are pre-\nsented in Table 7. Additional information on\npolymer quenchants for aluminum alloys can\nbe found in Ref 5.\nForming and Straightening after Quench-\ning. Immediately after being quenched,\nmost aluminum alloys are nearly as ductile\nas they are in the annealed condition. Con-\nsequently, it is often advantageous to form\nor straighten parts in this temper. More-\nover, at the mill level, controlled mechani-\ncal deformation is the most common meth-\nod of reducing residual quenching stresses.\nBecause precipitation hardening will occur\nat room temperature, forming or straighten-\ning usually follows as soon after quenching\nas possible. In addition, maximum effec-\ntiveness in stress relief is obtained by work-\ning the metal immediately after quenching.\nForming and straightening operations\nvary in degree from minor corrections of\nwarpage to complete forming of complex\nparts from solution-treated flat blanks. Par-\nticular value is gained when enough forming\ncan be done at this stage of processing to\neliminate the distortion caused by quench-\ning. However, production operations must\nbe adjusted so that most of the plastic\ndeformation is accomplished before an ap-\npreciable amount of precipitation hardening\ntakes place.\nAlthough the most severe forming opera-\ntions may have to be arranged to avoid\n856 / Heat Treating of Nonferrous Alloys\nAverage cooling rate at 400-290 \u00b0C, \u00b0C/s\nThickness, mm\n0.1\n1\n250\n75\n25\n25\n7.5\nAir cool\n2.5\n0.75\n0.25\n0.1\n1\nFig 7\n10\n10\n100\n103\n10\nComputed maximum\n(assumes instantaneous\ncooling of surface from\n875-210 \u00b0F)\n212 \u00b0F\n180 \u00b0F\n75 \u00b0F\n150 \u00b0F\n200 \u00b0F\nImmersion in water at\nindicated temperature\n100\nAverage cooling rate at 750-550 \u00b0F, \u00b0F/s\n0.1\n0.01\n103\n104\nThickness, in.\nEffects of thickness and quenching medium on average cooling rates at midplane of aluminum alloy\nsheet and plate quenched from solution temperatures. The dashed line delineates the maximum cooling\nrates theoretically obtainable at the midplane of plate, assuming an infinite heat transfer coefficient (C) and a\ndiffusivity factor of 1400 cm\u00b2/s. Source: Ref 3\nnatural aging, it often is desirable to allow\nsome natural aging to occur and thus avoid\nformation of L\u00fcders lines. This condition of\nnonuniform deformation is most likely to\noccur shortly after quenching and diminish-\nes significantly after a few hours of natural\naging. Complete freedom from L\u00fcders\nlines, however, may require one or two\ndays of natural aging prior to forming. Thus,\nthe forming operation may have to be timed\nso as to obtain the most appropriate trade-\noff of these characteristics for the specific\nparts involved. L\u00fcders lines also can be\nreduced by employing low strain rates or by\nforming at temperatures of 150 to 175 \u00b0C\n(300 to 350 \u00b0F).\nResidual stresses in sheet-metal parts\nformed in the quenched condition are higher\nthan those in parts formed in the annealed\ncondition. Consequently, forming in the\nquenched condition should be selected judi-\nprecipitation rate is low despite the high\ndegree of supersaturation. At intermediate\ntemperatures, precipitation rate is highest.\nConsequently, times to produce equal\namounts of precipitation follow a C-shape\npattern.\nUsing isothermal quenching techniques,\nFink and Willey pioneered the attempts to\ndescribe the effects of quench rates with the\nuse of C-curves (Ref 6). The C-curves plot\nthe time required at different temperatures\nto precipitate a sufficient amount of solute\nto: reduce strength by a certain amount (Fig\n5); cause a change in the corrosion behavior\nfrom pitting to intergranular (Fig 12); pro-\nduce a given electrical conductivity (Fig\n13); or relate other properties, such as frac-\nture toughness, to isothermal quench con-\nditions. The nose of the C-curves identifies\nthe critical temperature range (the region of\nhighest precipitation rates). Investigators\nuse critical temperature ranges in conjunc-\ntion with properties of samples quenched\ncontinuously from the solution temperature\nto compare relative sensitivities of alloys to\nquenching condition.\nAlthough average quench rates through a\ncritical temperature range can provide rea-\nsonable property predictions if cooling rates\nare fairly uniform, average quench rates can-\nnot provide quantitative predictions when\ncooling rates vary considerably during the\nciously for parts that are critical in fatigue quench. For such instances, a procedure\n(Fig 11) or stress corrosion.\nof parts\nRe-solution heat treatment\nformed after quenching often causes exces-\nsive grain growth in critically strained re-\ngions and thus is not recommended.\nQuench-Factor Analysis\nDuring the quenching of alloys from a\nsolid-solution temperature condition, the\nrate of precipitation during quenching is\nmaximized in a so-called \"critical\" temper-\nature range, because the diffusion of dis-\nsolved species and the subsequent nucle-\nation of precipitates exhibit opposite\nbehavior as a function of temperature. At\nhigh temperatures, nucleation rates are\nsmall because of the low degree of super-\nsaturation, and so precipitation rates are\nlow despite the high diffusion rates. At low\ntemperatures, diffusion rate is low, and thus\nTable 7 Limits for quenching in glycol-water solutions\nData are for wrought aluminum alloy products other than forgings.\nGlycol\nMaximum thickness\nconcentration,\nvol%\nAlloys\nmm\n12-16\n2014, 2017, 2117, 2024, 2219\n7075, 7175\n2.03\n25.4\n17-22\n2014, 2017, 2117, 2024, 2219\n1.80\nin.\n0.080\n1.000\n0.071\n7075, 7079, 7175, 7178, 6061\n12.7\n0.500\n23-28\n2014, 2017, 2117, 2024, 2219\n1.60\n0.063\n7075, 7079, 7175, 7178, 6061\n9.53\n0.375\n29-34\n2014, 2017, 2117, 2024, 2219\n1.02\n7075, 7079, 7175, 7178, 6061\n6.35\n35-40\n7075, 7079, 7175, 7178, 6061\n2.03\n0.040\n0.250\n0.080\nknown as \"quench-factor analysis\" uses in-\nformation from the entire C-curve to predict\nhow any quench curve affects properties.\nQuench-factor analysis is useful in designing\nsuitable limits for quench delays, or when it is\nnot sufficient just to ensure that the cooling\ncurve misses the nose of the C-curve.\nThe method of quench-factor analysis, as\noutlined by Evancho and Staley (Ref 7), is\nbased on the determination of a quench\nfactor (T), which is the major variable in the\nfollowing equation for precipitation kinetics\nduring continuous cooling:\n(=1- exp (KT)\n(Eq 1)\nwhere is the fraction transformed and k is\na constant related to the transformation\nfraction of a given C-curve. The quench\nfactor (T) is defined as:\nT=\n- \u1042\ndt\n(Eq 2)\nwhere is time and C, is critical time as a\nfunction of temperature to transform a spec-\nified fraction (x). The locus of critical times\nfor a given transformation fraction x (or a\npercentage of mechanical properties from\nprecipitation) is the C-curve, and the value\nof k is related to x as follows: k = ln(1 - x),\nor e\u2b51\n= 1 x. Therefore, when T = 1, the\nfraction transformed, 5, equals the fraction\nvalue designated by the C-curve. Equation\n2 is based on the assumption that the reac-\ntion rate is a function only of the amount\ntransformed and temperature.\nTemperature, \u00b0C\n(a)\nTemperature, \u00b0C\n(b)\nFig 8\n500\n400\n300\n200\n100\n500\n400\n300\n200\nBlack oxide\netched coating\n1\n2\n3\nTime, s\n4\n100\nPowdered oxide\nsprayed coating\n0\n20 \u00b0C water quench\nAs-rolled\nsurface\nSanded\nsurface\n5\nBoiling water quench\nBlack oxide\netched coating\nAs-rolled\nsurface\n6\n800\n600\n400\n200\n800\n600\nSurface\n400\nSanded\nsurface\n10\n20\n30\n40\n50\n60\n70\n80\nTime, s\n200\nTemperature, \u00b0F\nTemperature. \u00b0F\nEffect of surface conditions on the midplane cooling of a 13 mm (0.5 in.) thick plate of 7075 from\nquenching in (a) 20 \u00b0C (70 \u00b0F) water and (b) boiling water. Source: Ref 5\nTension\nLongitudinal stress, MPa\nCompression\nCross section of solid cylindrical\n2\nspecimen, in.\na\n10\n20 30\n40\n100\n50\n50\n100\n0\n10\n1\nQuenched\nin boiling\nwater\n10\n20\n30\nCross section of solid cylindrical\nspecimen, 103 mm\u00b2\n5\n5\nTension \u2192\n10\n\u2713 Compression\nTension\u2192\u2192\nLongitudinal stress, ksi\nLongitudinal stress, MPa\n- Compression\nCross section of solid cylindrical\nspecimen, in.2\n0\n10\n100\n20 30\nT\n40\n0\n50\n100\n0\n10\nQuenched\n10\nin cold\nwater\n20\nT\n30\nCross section of solid cylindrical\nspecimen, 103 mm\u00b2\n10\n- Compression\nTension\nLongitudinal stress, ksi\nFig 9 Effect of quenching from 540 \u00b0C (1000 \u00b0F) on residual stresses in solid cylinders of alloy 6151\nHeat Treating of Aluminum Alloys / 857\nThe numerical evaluation of the quench\nfactor involves the integration of Eq 2. This\nintegral can be graphically integrated using\nthe method illustrated in Fig 14. Examples\nof the way to use the quench factor (T) in the\nanalysis of quench methods are described\nbelow. Neither the average quenching rate\nthrough a critical temperature range nor\nquench-factor analysis can predict strength\nwhen the temperature increases during\nquenching after it is cooled below some\ncritical temperature. Under this condition,\nstrength in the affected areas can be signif-\nicantly lower than in other areas of the\nmaterial. The most likely way for this phe-\nnomenon to occur is during spray quench-\ning, when the surface cools rapidly by the\nimpinging spray, but reheats by heat flow\nfrom the hotter interior when the spray is\ninterrupted.\nPredicting Strengths of Thick Products.\nEffects of the quenching rate on alloy\nstrengths can be represented on a general-\nized graph of the type shown in Fig 3, and\nthe expected quenching rates of products\nhaving various dimensions can be deter-\nmined from Fig 7. Nevertheless, combining\nthese two kinds of information to predict\nmechanical properties must be done with\ncaution. Inconsistencies were encountered,\nfor example, in correlating properties of\nthick sections quenched in high-cooling-rate\nmedia with properties of thinner sections\nquenched in media affording milder quench-\ning action. One of the reasons for the incon-\nsistencies is believed to be the different\nshapes of the cooling curves. This difficulty\ncan be overcome by using quench-factor\nanalysis. The other reason is that the degree\nof recrystallization and texture of the thick\nand thin sections may be different.\nPredicting Corrosion Behavior. Alloy\n2024-T4, for example, is susceptible to in-\ntergranular corrosion when a critical\namount of solute is precipitated during\nquenching, but will corrode in the less se-\nvere pitting mode when lesser amounts are\nprecipitated. For predicting the effects of\nproposed quenching conditions on the cor-\nrosion characteristics of 2024-T4, the postu-\nlated quench curve is drawn and the quench\nfactor is calculated using the C-curve in Fig\n12. Corrosion characteristics are predicted\nfrom the plot in Fig 15. When the quench\nfactor (T) is less than 1.0, continuously\nquenched 2024-T4 will corrode by pitting.\nThese relationships are applied to studies\nof effects of proposed changes in quench\npractice on design of new quenching sys-\ntems. For example, consider that the goal of\na proposed quenching system for 2024-T4\nsheet products is to minimize warpage while\npreventing susceptibility to intergranular\ncorrosion. Warpage occurs when the stress-\nes imposed by temperature differences\nacross the parts exceed the flow stress. As\nquenching rate decreases, the tendency for\nlarge differences in temperature to occur\n% of maximum yield strength\n858 / Heat Treating of Nonferrous Alloys\nThickness, 0.001 in.\n100\n150\n100\n70\n50\n60\n50\n90\n6061-T6 sheet\n1250\n2500\n3750\n200\nWater\nspray\n250\nAir blast\nStress, MPa\n400\n3\nCurve Bend radius\n123\nNot bent\n3.2 mm\nCondition during flattening\nNot applicable\n50\nAnnealed\n3.2 mm\n4\n3.2 mm\n300\n5\n1.6 mm\nAs-quenched + 3 days storage\nAs-quenched + 14 days storage\nAs-quenched + 3 days storage\n40\n200\n5000\n6250\n7500\n100\n0.01\nThickness, um\nThickness, 0.001 in.\nStress ratio, 0.1\n0.1\n4\n5\n1\n30\n1\n2\n20\n3\n10\nStress, ksi\n% of maximum tensile strength\n100\n150\n100\n90\n80\n200\n250\nWater spray\nAir blast\n70\n6061-T6 sheet\n601250\n2500\n3750\n5000\n6250\nThickness, um\nFig 10\n7500\nEffect of quenching medium on strength of\n6061-T6 sheet. Water-immersion quench\nequals 100%. Control of coolant flow will minimize\ndecrease in mechanical properties.\ndecreases but the tendency for intergranular\ncorrosion to occur increases.\nThe C-curve in Fig 12 indicates that\nquenching rate can be decreased near the\nsolution heat-treating temperature and near\nroom temperature without greatly sacrific-\ning corrosion characteristics, but this infor-\nmation does not provide a quantitative an-\nswer. Simple calculations, however, can\nreveal a multitude of hypothetical cooling\ncurves that provide slow quenching during a\nlarge portion of the quench cycle but suffi-\nciently rapid quenching where critical times\nare short so that desirable corrosion char-\nacteristics are obtained.\nAs an example, one-, two-, and three-step\nquench curves that would ensure accept-\nMillions of cycles to failure\nFig 11 Fatigue characteristics of 1 mm (0.04 in.) alclad 2024-T4 sheet after 90\u00b0 bending in the annealed\ncondition and subsequent flattening as indicated. Flattening (unbending) was done either in the\nannealed condition (curve 2), or in the solution-treated and quenched condition (curves 3, 4, 5) with indicated\nstorage times at -18 to -12 \u00b0C (0 to 10 \u00b0F).\nable corrosion behavior in 2024-T4 sheet\n(quench factor, 0.99) were calculated. Some\nof these curves are plotted in Fig 16. This\nillustration shows that 2024 can be\nquenched at a rate of 470 \u00b0C/s (850 \u00b0F/s) or\nhigher and still develop acceptable corro-\nsion characteristics if the quenching rate is\nlinear from the solution temperature to 150\n\u00b0C (300 \u00b0F). If sheet 3.2 mm (0.125 in.) thick\nis air-blast quenched (rate of heat removal,\n5.68 W/m\u00b2 \u00b0C) to 395 \u00b0C (740 \u00b0F), however,\nthe quenching rate from 395 to 150 \u00b0C must\nbe at least 945 \u00b0C/s (1700 \u00b0F/s) to maintain\nthe acceptable corrosion behavior. It may\nalso be air-blast quenched to 395 \u00b0C (740\n\u00b0F), spray quenched at 3300 \u00b0C/s (6000 \u00b0F/s)\nto 250 \u00b0C (480 \u00b0F), then air-blast quenched\nto 150 \u00b0C (300 \u00b0F).\nHolding temperature, K\n700\n650\n600\n550\nOther curves could be drawn, of course,\nbut the important points are that air-blast\nquenching cannot be continued to more\nthan a few degrees below 395 \u00b0C (740 \u00b0F)\nand cannot be initiated at more than a few\ndegrees above 270 \u00b0C (520 \u00b0F) even if infi-\nnite quenching rates are attained from 395\nto 270 \u00b0C (740 to 520 \u00b0F).\nPredicting yield strength is more complex\nthan predicting corrosion behavior and re-\nquires some knowledge of the relationship\nbetween extent of precipitation and loss in\nability to develop property. Because attain-\nable strength of precipitation-hardening alu-\nminum alloys is a function of the amount of\nsolute remaining in solid solution after\nquenching, relationships between strength\n(\u03c3) attainable after continuous cooling and\nIACS\n17%\n18%\n425\n375\n19%\n20%\n325\n21%\n275\nHolding temperature, \u00b0C\nTemperature, \u00b0C\n500\n900\n400\n300\n200\n100\n0.1\nPredominantly\npitting\n1\nPredominantly\nintergranular\ncorrosion\n700\n500\n300\nTemperature, \u00b0F\n10\n100\n103\nCritical time, s\nFig 12\nC-curve indicating type of corrosion attack\non 2024-T4 sheet\n500\n17%\n450\n10\n100\n225\n19%\n20%\n21%\n18%\n22%\n175\n103\n104\n105\n106\nHolding time, s\nFig 13 Change in electrical conductivity of an Al-2.5% Li binary alloy after the following: solution treated at 540\n\u00b0C (1000 \u00b0F) for 12 h, immersed into an adjacent salt or oil bath for the appropriate isothermal holding\ntemperature and time, then quenched into water. Source: Ref 1\nTemperature\n\u30c9\u30c9\u30c9\nQuench curve\nTemperature\n(T\u2081 + T\u2082)/2\n\u3051\n+ T3)/2\nC-curve\nTemperature, \u00b0C\nHeat Treating of Aluminum Alloys / 859\n500\n900\nAir-blast quench\n3.2-mm (0.125-in.) sheet\non both sides\n800\n400\n470 \u00b0C/s (850 \u00b0F/s)\n700\n2780 \u00b0C/s (5000 \u00b0F/s)\nAir-blast quench\n600\n300\n3.2-mm (0.125-in.)\nsheet\n500\nTemperature, 'F\nTEITI\n1++AG\n|\n1+AtF-1\nTF-1 tF\nElapsed time\n(TF + TF)/2\n\u04212\nC\u2081\u2081 CF-1\nCritical time\nx =\n+\n\u0394\u03b9\u03b1 \u0394\u03b9\u03b1\nC\u2082\n+\nAtF-1\nCF-1\nFig 14 Method of determining quench factor, &, using a cooling curve and a C-curve\nAverage depth, mm\n0.20\nQuench\no Air-water-air\n\u2022 Water-air\n0.15\nA Air-water\n\u25b2 Water\n0.10\n\u25a1 Air\n0.05\nType of corrosion\nT\nQA\n0\n0.1\n1\n0.2\n\u0567\n1\nL\n1\n0.4\n0.6 0.8 1.0\n2.0\n4.0\n6.0 8.0 10.0\n20.0\nQuench factor, T\nP + SI\nP +\n0.1\nQuench\nAir-water-air\n\u2022 Water-air\n\u25b2 Air-water\n\u25b2 Water\n\u25a1 Air\nP-Pitting\n0.006\n0.004\n0.002\n2 P+SI - Pitting and slight intergranular\nP+1 Pitting and intergranular\nIntergranular\n0800\n0.2\n0.4\n0.6 0.8 1.0\n2.0\nQuench factor, 1\n4.0 6.0 8.0 10.0\n20.0\nFig 15 Type and depth of attack on 2024-T4 sheet versus quench factor\nquench factor (T) can be expressed as fol-\nlows:\n(Eq 3)\n\u03c3x=\u03c3max exp (k\u2081T)\nwhere \u03c3 max\nis the strength attainable with\nan infinite quenching rate and:\nT =\nSa\ndt\n(Eq 4)\nwhere t is time and Cx is the C-curve for\n\u03c3 that is, critical time as a function of\ntemperature to reduce attainable strength to\nx of \u03c3max. The constant k\u2081 is related to the\nnatural logarithm of x. For example, if 7 is\nbased on the C-curve for 99.5% of maxi-\nmum yield strength, then k\u2081 = -0.005013 =\nIn (0.995).\nThe advantage of predicting yield\nstrength from quench factor instead of from\naverage quenching rate is illustrated by the\nAverage depth, in.\nfollowing comparison. Four specimens of\nalloy 7075-T6 quenched by various means\n(see Fig 17) were selected. Yield strengths\nwere predicted both from average quench-\ning rate between 400 and 290 \u00b0C (750 and\n550 \u00b0F) and from quench factor. Quench\nfactor was calculated using the C-curve for\n99.5% maximum yield strength for 7075-T6\n(Fig 18), and yield strength was estimated\nfrom the above equation defining the\nquench factor (7) (see Fig 19).\nA comparison of predicted yield strength\nwith actual yield strength is given in Table\n8. Yield strengths predicted from quench\nfactor agree very well with measured yield\nstrengths for all specimens, the maximum\nerror being 19.3 MPa (2.8 ksi). Yield\nstrengths predicted from average quenching\nrates, however, differ from measured val-\nues by as much as 226 MPa (32.8 ksi).\n200\n400\n3330 \u00b0C/s\n(6000 F/s)\n945 \u00b0C/s\n300\n(1700 F/s)\n100\n1\n0\n2\n4\n6\nTime, s\nFig 16 Quench curves for 2024-T4 sheet, to elimi-\nnate susceptibility to intergranular corrosion\nThe advantage of using the quench factor\nfor predicting yield strength from cooling\ncurves is apparent. Cooling curves that\nhave long holding times either above or\nbelow the critical temperature range from\n400 to 290 \u00b0C (750 to 550 \u00b0F) cannot be used\nto predict yield strength from average\nquenching rate. In such instances, predic-\ntion of yield strength on the basis of quench\nfactor is particularly advantageous.\nWith the use of finite-element analysis,\nquench factors can also be plotted as a\nfunction of Grossmann quench severity val-\nues (H) or the heat transfer coefficients (C)\nbetween the quenchant and a particular part\n(Fig 20). However, an underlying assump-\ntion of both quench-factor analysis and av-\nerage-cooling-rate estimation is that the\nonly effect of temperature is on the kinetics\nof precipitation. This assumption is not val-\nid, however, when portions of the metal are\nquenched locally but reheated significantly\nbefore quenching is complete.\nAge Hardening\nAfter solution treatment and quenching,\nhardening is achieved either at room tem-\nperature (natural aging) or with a precipita-\ntion heat treatment (artificial aging). In\nsome alloys, sufficient precipitation occurs\nin a few days at room temperature to yield\nstable products with properties that are\nadequate for many applications. These al-\nloys sometimes are precipitation heat treat-\ned to provide increased strength and hard-\nness in wrought or cast products. Other\nalloys with slow precipitation reactions at\nroom temperature are always precipitation\nheat treated before being used.\nIn some alloys, notably those of the 2xxx\nseries, cold working of freshly quenched\nmaterial greatly increases its response to\nlater precipitation heat treatment. Mills take\nadvantage of this phenomenon by applying\na controlled amount of rolling (sheet and\nplate) or stretching (extrusion, bar, and\nplate) to produce higher mechanical proper-\nties. However, if the higher properties are\nused in design, reheat treatment must be\navoided.\nTemperature, \u00b0C\n860 / Heat Treating of Nonferrous Alloys\n500\nQuenched in still air to 370 \u00b0C (700 \u00b0F) + cold-water quench\n900\n500\n800\n400\n300\nQuenched in boiling water to\n315 \u00b0C (600 \u00b0F) + cold-water quench\n700\n-600\n00\nTemperature, \u00b0F\n- 400\nTemperature, \"C\n300\n\" 500\n200\nF6 Quenched in denatured alcohol to 290 \u00b0C\n\u2020(550 \u00b0F) cold-water quench\nCold-water quench\n200\n100\n0\n5\n10\n15\n20\n25\n30\n35\nTime, s\n400\n300\n300\n100\n0.1\n1\n10\n100\nCritical time, s\n1000\n10 000\n900\n700\n500\nFig 17 Cooling curves for 7075-T6 sheet\nNatural Aging. The more highly alloyed\nmembers of the 6xxx wrought series, the\ncopper-containing alloys of the 7xxx group,\nand all of the 2xxx alloys are almost always\nsolution heat treated and quenched. For\nsome of these alloys-particularly the 2xxx\nalloys the precipitation hardening that re-\nsults from natural aging alone produces\nuseful tempers (T3 and T4 types) that are\ncharacterized by high ratios of tensile to\nyield strength and high fracture toughness\nand resistance to fatigue. For the alloys that\nare used in these tempers, the relatively\nhigh supersaturation of atoms and vacan-\ncies retained by rapid quenching causes\nrapid formation of GP zones, and strength\nincreases rapidly, attaining nearly maxi-\nmum stable values in four or five days.\nTensile-property specifications for products\nin T3- and T4-type tempers are based on a\nnominal natural aging time of four days. In\nalloys for which T3- or T4-type tempers are\nstandard, the changes that occur on further\nnatural aging are of relatively minor magni-\ntude, and products of these combinations of\nalloy and temper are regarded as essentially\nstable after about one week.\nIn contrast to the relatively stable condi-\ntion reached in a few days by 2xxx alloys\nthat are used in T3- or T4-type tempers, the\n6xxx alloys and to an even greater degree\nthe 7xxx alloys are considerably less stable\nat room temperature and continue to exhibit\nsignificant changes in mechanical properties\nfor many years. The differences in rate and\nFig 18 C-curve for 99.5% maximum yield strength of 7075-T6 sheet\nduration of changes in tensile yield strength\nof representative alloys of the three types\nare illustrated in Fig 21. Because of the\nrelative instability of the 7xxx alloys, the\nnaturally aged temper (after solution heat\ntreatment and quenching) is designated by\nthe suffix letter W. For a specific descrip-\ntion of this condition, the time of natural\naging should be included (example: 7075-W,\n1 month).\nAging characteristics vary from alloy to\nalloy with respect to both time to initial\nchange in mechanical properties and rate of\nchange, but aging effects always are less-\nened by reductions in aging temperature\n(see Fig 21). With some alloys, aging can be\nsuppressed or delayed for several days by\nholding at a temperature of -18 \u00b0C (0 \u00b0F) or\nlower. It is usual practice to complete form-\ning and straightening before aging changes\nmechanical properties appreciably. When\nscheduling makes this impractical, aging\nmay be avoided in some alloys by refriger-\nating prior to forming. It is conventional\npractice to refrigerate alloy 2024-T4 rivets\nto maintain good driving characteristics.\nFull-size wing plates for current-generation\njet aircraft have been solution heat treated\nand quenched at the primary fabricating\nmill, packed in dry ice in specially designed\ninsulated shipping containers and transport-\ned by rail about 2000 miles to the aircraft\nmanufacturer's plant for forming.\nUnanticipated difficulties may arise as a\nresult of failure to control refrigerator or\npart temperature closely enough. If opening\nof the cold box to insert or remove parts is\ndone too frequently, the cooling capacity of\nthe refrigerator may be exceeded. At times,\nthe rate at which heavy-gage parts can be\ncooled in a still-air cold box has been found\nto be insufficient. This problem has been\nsolved in one plant by immersing parts in a\nsolvent at -40 \u00b0C (-40 \u00b0F) before placing\nthem in the refrigerator.\nThe T3-type tempers are distinguished\nfrom T4-type tempers by significant me-\nchanical-property differences resulting from\ncold work strain hardening associated with\ncertain mechanical operations performed\nafter quenching. Roller or stretcher leveling\nto achieve flatness or straightness introduc-\nes modest strains (on the order of 1 to 4%)\nthat cause changes in mechanical properties\n(primarily, increases in strength). Further\nincreases in strength can be obtained by\ncold rolling, additional stretching, combina-\ntions of these operations, or for products\nsuch as hand forgings, compressive defor-\nmation. The tempers produced by these\noperations followed by natural aging alone\n(no precipitation heat treatment) are classi-\nfied as T3-type tempers, and an additional\ndigit is used to indicate a variation in strain\nhardening that results in significant changes\nin properties. In the most recently intro-\nTable 8 Yield-strength values for 7075-T6 sheet predicted from cooling curves using\naverage quench rate and quench factor\n% of maximum yield strength\nAverage quench\nrate from 400 to\n290 \u00b0C (750 to\n100\n550 \u00b0F)\nQuench\nfactor,\nMeasured yield\nstrength\nYield strength\npredicted from\naverage quench\nrate\nYield strength\npredicted from\nquench factor\n80\nQuench\n\u00b0C/s\n\u00b0F/s\nT\nMPa\nksi\nMPa\nksi\nMPa\nksi\n60\n100 o/max exp(-0.005013 t)\nCold water\n935\n1680\n0.464\n506\n73.4\n499\n72.4\n498\n72.3\nDenatured alcohol to\n40\n290 \u00b0C (550 \u00b0F),\nthen cold water\n50\n90\n8.539\n476\n69.1\n463\n67.2\n478\n69.4\n20\nBoiling water to\n0\n315 \u00b0C (600 \u00b0F),\n0.01\n0.1\n1\n10\nQuench factor, t\n100\n103\nthen cold water\n30\n55\n15.327\n458\n66.4\n443\n64.2\n463\n67.1\nStill air to 370 \u00b0C\n(700 \u00b0F), then cold\nwater\n5\n9\n21.334\n468\n67.9\n242\n35.1\n449\n65.1\nFig 19 Yield strength versus quench factor\nTemperature, \"F\n\n\nQuench factor for 99.5% of\nattainable yield strength for alloy 7075-T73\n10\n5\n1\n40\n45\nSheet thickness, in.\nPlate thickness, in.\n0\n0.05\n0.1\n0.15\n0.2\n0.25\n0.3\n0\n0.5\n1\n1.5\n2\n2.5\n3\n50\n550\n50\n0\n1.25\n2.5\n(a)\nC=0.036 W/cm2.K\n3.75\nSheet thickness, mm\nC = 0.072\nC=0.108\nC = 0.144\nC = 0.216\nC = 0.288\nC = 0.432\n5\nC = 0.720\nC = 1.44\n6.25\n7.5\nQuench factor for 99.5% of\nattainable yield strength for alloy 7075-T73\n10\n5\n(b)\n1\n45\nC = 0.216\nHeat Treating of Aluminum Alloys / 861\nC = 2.16\nC = 2.88\nC = 4.32\nC=7.20\nPlot of quench factors derived from finite element analysis with given product sizes and film (heat transfer) coefficients (C). Heat transfer coefficients between\nFig 20\nthe quenchant and part are expressed in W/cm\u00b2 K. Source: Ref 4\n12.5\n25\n37.5\n50\n62.5\n75\nPlate thickness, mm\nduced 2xxx aircraft alloy, 2324, high\nstrength is achieved by cold rolling plate to\na T39 temper.\nPrecipitation heat treatments generally\nare low-temperature, long-term processes.\nTemperatures range from 115 to 190 \u00b0C (240\nto 375 \u00b0F); times vary from 5 to 48 h.\nChoice of time-temperature cycles for\nprecipitation heat treatment should receive\ncareful consideration. Larger particles of\nprecipitate result from longer times and\nhigher temperatures; however, the larger\nparticles must, of necessity, be fewer in\nnumber with greater distances between\nthem. The objective is to select the cycle\nthat produces optimum precipitate size and\ndistribution pattern. Unfortunately, the cy-\ncle required to maximize one property, such\nas tensile strength, is usually different from\nthat required to maximize others, such as\nyield strength and corrosion resistance.\nConsequently, the cycles used represent\ncompromises that provide the best combi-\nnations of properties.\nProduction of material in T5- through\nT10-type tempers (see the section on tem-\nper designations near the end of this article)\nnecessitates precipitation heat treating at\nelevated temperatures (artificial aging). Al-\nthough the hardening precipitate developed\nby this operation is submicroscopic, struc-\ntures before and after precipitation heat\ntreatment often can be distinguished by\netching metallographic specimens. In alumi-\nnum alloys in the solution heat treated and\nquenched condition, coloration contrast be-\ntween grains of differing orientation is rela-\ntively high, particularly in 2xxx series\nwrought alloys and 2xx.0 series casting al-\nloys. This contrast is noticeably decreased\nby precipitation heat treatment.\nDifferences in type, volume fraction,\nsize, and distribution of the precipitated\nparticles govern properties as well as the\nchanges observed with time and tempera-\nture, and these are all affected by the initial\nstate of the structure. The initial structure\nmay vary in wrought products from unre-\ncrystallized to recrystallized and may ex-\nhibit only modest strain from quenching or\nadditional strain from cold working after\nsolution heat treatment. These conditions,\nas well as the time and temperature of\nprecipitation heat treatment, affect the final\nstructure and the resulting mechanical prop-\nerties.\nBecause mechanical properties and other\ncharacteristics change continuously with\ntime and with temperature, as shown in Fig\n22(a), (b), and (c) by typical curves for three\nwrought alloys, treatment to produce a\ncombination of properties corresponding to\na specific alloy-temper combination re-\nquires one or more rather specific and co-\nordinated combinations of time and temper-\nature, with both parameters being subject to\npractical limitations. Recommended com-\nmercial treatments often are compromises\nbetween time and cost factors and the prob-\nability of obtaining the intended properties,\nwith consideration of allowances for varia-\nbles such as composition within specified\nrange and temperature variations within the\nfurnace and load. Use of higher tempera-\ntures may reduce treatment time; but if the\ntemperature is too high, characteristic fea-\ntures of the precipitation-hardening process\nreduce the probability of obtaining the re-\nquired properties.\nT6 and T7 Tempers. Precipitation heat\ntreatment following solution heat treatment\nand quenching produces T6- and T7-type\ntempers. Alloys in T6-type tempers gener-\nally have the highest strengths practical\nwithout sacrifice of the minimum levels of\nother properties and characteristics found\nby experience to be satisfactory and useful\nfor engineering applications. Alloys in T7\ntempers are overaged, which means that\nsome degree of strength has been sacrificed\nor \"traded off\" to improve one or more\nother characteristics. Strength may be sac-\nrificed to improve dimensional stability,\nparticularly in products intended for service\nat elevated temperatures, or to lower resid-\nual stresses in order to reduce warpage or\ndistortion in machining. T7-type tempers\nfrequently are specified for cast or forged\nengine parts. Precipitation heat-treating\ntemperatures used to produce these tem-\npers generally are higher than those used to\nproduce T6-type tempers in the same al-\nloys.\nTwo important groups of T7-type\ntempers-the T73 and T76 types-have\nbeen developed for the wrought alloys of\nthe 7xxx series, which contain more than\nabout 1.25% copper. These tempers are\nintended to improve resistance to exfolia-\ntion corrosion and stress-corrosion crack-\ning, but as a result of overaging, they also\nincrease fracture toughness and, under\nsome conditions, reduce rates of fatigue-\ncrack propagation. The T73-type temper\n862 / Heat Treating of Nonferrous Alloys\nTensile strength, MPa\nYield strength, MPa\nElongation, % in 50 mm or 2 in.\n600\n2014\n500\nRT\n400\n0 \u00b0C (32 \u00b0F).\n60\n300\n200\n80\n2024\n600\n80\nT\n500\n70\n70\nRT\nTensile strength, ksi\nTensile strength, MPa\n0 \u00b0C (32\u00b0F)\n60\n400\n-18 \u00b0C (0 \u00b0F)\n50\n--18\u00b0C (0\u00b0F)\n40\n300\nT\nT\n200\n1 year\n20\n30 min\n1 day\n1 week 2 months.\n100\n0.1\n1\n102\n103\n104\n30\n1 year\n30 min\n1 day\n1 week 2 months\n100\n0.1\n1\n102\n103\n104\n500\n400\n10\nElapsed time after quenching, h\n1\n70\nT\n20\nI\n1\n60\n50\n50\n300\nRT\n40\n0\u00b0C (32\u00b0F)\n30\n200\n100\nYield strength, ksi\nYield strength, MPa\n500\n400\n300\n200\n10\nElapsed time after quenching, h\nRT\nT\n50\n40\n30\n20\n10\n70\n60\n50\nTensile strength, ksi\n30\nYield strength, ksi\n0\u00b0C (32 F)\n40\n30\n-18 C (OF)\n20\n20\n20\n-18 \u00b0C (0 \u00b0F)\n100\n1 year\n10\n1 year\n10\n30 min\n1 day\n1 week 2 months\n30 min\n1 day\n1 week 2 months\n11\n0\n0\n0\n0.1\n1\n102\n103\n104\n0\n0.1\n1\n102\n103\n104\n40\n30\n20\n20\n10\n10\nElapsed time after quenching, h\n-18 \u00b0C (0 \u02daF)\n0 \u00b0C (32\u00b0F)\nElongation, % in 50 mm or 2 in.\n40\n30\n10\nElapsed time after quenching, h\n-18 \u00b0C (0\u00b0F)\n0\u00b0C (32\u00b0F)\nRT\n20\n10\n10\n1 year\n1 week 2 months\n1 year\n30 min\n1 day\n1 week 2 months\n30 min\n1 day\nI\nI\n0\n0\n0.1\n1\n10\n102\n103\n104\n0.1\n1\n10\n1\n102\nElapsed time after quenching, h\n11\n103\nElapsed time after quenching, h\n104\nFig 21 Aging characteristics of aluminum sheet alloys at room temperature, at 0 \u00b0C (32 \u00b0F), and at -18 \u00b0C (0 \u00b0F)\nhas greatly minimized stress-corrosion\ncracking of large and complex machined\nparts made of these alloys, which occa-\nsionally occurred with T6-type tempers.\nThe precipitation heat treatment used to\nproduce the T73- and T76-type tempers\nconsist either of a two-stage isothermal\nprecipitation heat treatment or of heating\nat a controlled rate to a single treatment\ntemperature. The microstructural/\nelectrochemical relationships that are re-\nquired in order to achieve the desired\ncorrosion-resisting characteristics can be\ndeveloped by using only a single-stage\nprecipitation heat treatment above about\n150 \u00b0C (300 \u00b0F), but higher strength is\nobtained by preceding this with a lower-\ntemperature stage or with a slow-con-\ntrolled heatup. Extended natural aging can\nprovide the same results, but the times\nrequired at room temperature are imprac-\ntical. Either during the preliminary stage\nor during slow heatup, a fine, high-density\ndispersion of GP zones is nucleated. Either\nthe time and temperature of the first step\nor the rate of heating must be controlled to\nproduce GP zones that will not dissolve\nbut will transform to the n' precipitate\nwhen heated to the aging temperature\nabove 150 \u00b0C (300 \u00b0F). The aging practice\nthat produces the results in the shortest\ntime depends on the GP-zone solvus tem-\nperature. This temperature, in turn, de-\npends on vacancy concentration, a factor\ninfluenced by solution heat-treating tem-\nperature and quench rate, and on compo-\nsition. If first-step aging time is too short,\nif first-step aging temperature is too far\nbelow the GP-zone solvus, or if heating\nrates are too high, the GP zones will\ndissolve above 150 \u00b0C (300 \u00b0F), and the\nresultant coarse and widely distributed\nprecipitate will provide lower strength.\nThe T76-type treatments have the same\noperational sequence but employ second-\nstage heating only long enough to develop\na resistance to exfoliation corrosion higher\nthan that provided by the T6-type tempers.\nMaterials in the T73-type temper also have\nhigh resistance to exfoliation corrosion.\nRecommended treatments to produce T5-\nand T6-type tempers, and those of the T7-\ntype employed for dimensional and proper-\nty stabilization, provide adequate tolerance\nfor normal variations encountered with\ngood operating practices. On the other\nhand, the T73, T74 (formerly T736), and\nT76 tempers for alloys 7049, 7050, 7075,\n7175, and 7475 involve changes in strength\nthat occur significantly more rapidly at the\ntemperatures employed in the second stage\nof the T7x precipitation heat-treatment cy-\ncle compared to the changes occurring at\nthe temperatures employed to produce the\nT6 temper.\nAs illustrated in Fig 23, variations in soak\ntime of several hours, and variations in soak\ntemperature of up to 11 \u00b0C (20 \u00b0F) from the\nnominal aging practice of 24 h at 120 \u00b0C (250\n\u00b0F) affect the strength of 7075-T6 by as\nmuch as 28 MPa (4 ksi). In contrast, similar\nvariations in second-step soak time and\ntemperature for 7075-T73-that is, varia-\ntions for 24 h at 165 \u00b0C (325 \u00b0F)-affect\nstrength by up to 150 MPa (22 ksi).\nConsequently, control of both tempera-\nture and time to achieve the mechanical\nproperties and corrosion resistance speci-\nfied for these tempers is more critical than\nthe control required in producing the T6\ntemper. Moreover, rate of heating from the\nfirst to the second aging step must be con-\nsidered, because precipitation occurs dur-\ning this period.\nHeat treaters attempt to adjust these new\nproblems by empirically modifying soak\ntimes to compensate for precipitation dur-\ning heating and for effects of soaking at\ntemperatures above or below the nominal.\nA method has been developed (Ref 8) that\npermits quantitative compensation for the\neffects of precipitation during heating and of\nsoaking either above or below the recom-\nmended temperature. For overaging, these\neffects can be described by the following\nequation:\nYS = Y exp\n-\n(\n+ 0\nFYS\n(Eq 5)\nwhere YS is yield strength; Y is a term\nhaving units of strength that is dependent on\nalloy, fabrication, and test direction; to is\ntime at soak temperature; Fys is a temper-\nature-dependent term; and\nYS\nTensite strength, MPa\nYield strength, MPa\nElongation, \" in 50 mm or 2 in.\n6061\n600\n500\n400\n300\n200\n30 min\n100\n0.1\n1\n500\n400\n300\n200\n10\nRT\n0 \u00b0C (32\u00b0F)\nT\nT\nT\n--\n80\n70\n60\n50\n40\n10\nTensile strength, ksi\nTensile strength, MPa\n600\n7050\n00\n80\n500\n70\nRT\n60\n400\n300\nT\n50\n40\nHeat Treating of Aluminum Alloys / 863\nTensile strength, ksi\nTensile strength, MPa\n600\n500\n7075\n80\n70\n0\u00b0C (32\u00b0F)\nRT\n60\n400\n300\n-18 \u00b0C (0\u00b0F)\n50\n40\n40\nT\nT\n30\n200\n30\n200\n-18 \u00b0C (0 \u00b0F)\n30\n30\n1 year\n1 year\n-\n1 day 1 week 2 months\n20\n20\n1 year\n30 min\n1 day\n1 week\n2 months\n20\n| L\n102\n11\n100\n1\n\u2610 1\n30 min\n\u2610\n100\n1 day\n1\n1 week 2 months\n11\n103\n104\n0.1\n1\n10\n102\n103\n104\n0.1\n1\n10\n102\n103\n104\n24\n20\nElapsed time after quenching, h\nT\nT\nT\n500\n70\n60\n50\n40\n30\nYield strength, ksi\nYield strength, MPa\n400\n300\n200\nElapsed time quenching, h\nRT\n500\n70\n0\n60\nT\nT\nT\n60\n400\n50\n50\n40\n40\n30\nYield strength, ksi\nYield strength, MPa\nElapsed time after quenching, h\n300\nRT\n0 \u00b0C (32\u00b0F)\n200\nT\n70\n660\n650\n40\n30\nRT\n20\n20\nT\n20\n20\n100\n0 \u00b0C (32 F:\n100\n-18 \u00b0C (0 \u00b0F)\n100\n-18 \u00b0C (0 \u00b0F) 1 year 10.\n1 year\n10\n10\n30 min\n1 day 1 week\n2 months\n30 min\n1 day\n1 week 2 months\n1 year\n0\n0.1\n1\n\u2610\u2610\n10\n11\n102\n0\n0\n11\n\u2610\n\u2610\n0\n30 min\n1 day\n1 week 2 months\n103\n104\n0.1\n1\n10\n102\n103\n104\n0\n1\n11\n0\n0.1\n1\n10\n102\n103\n104\nElapsed time after quenching, h\nElapsed time after quenching, h\n40\n30\n20\n10\n-18 \u00b0C (0\u00b0F)\n0 \u00b0C (32\u00b0F)\nRT\n1 year\n30 min\n1 day\nT\n0\n0.1\n1\n10\n1 week 2 months\n\u2610\n103\n102\nElapsed time after quenching, h\nFig 21 (continued)\nElongation, % in 50 mm or 2 in.\n40\n30\nRT\n1 year\n30 min\n1 day\n1 week 2 months.\n104\n0\n0.1\n\u2610\n1\n\u2610 1\n1\n10\n102\n103\n104\nElapsed time after quenching, h\nElongation, % in 50 mm or 2 in.\n40\n30\nElapsed time after quenching, h\n20\n0 \u00b0C (32\u00b0F)\n10\n-18 \u00b0C (0 \u00b0F)\nRT\n1 year\n30 min\n0\n1 day\n1\n0.1\n1\n10\n1 week 2 months\n\u2610\n102\n103\n104\nElapsed time after quenching, h\nTensile strength, ksi\nYield strength, ksi\n0=\nS\ndt\nFys\n(Eq 6)\nFys 1.45\u00d710-16\nexp (\n18 090\n)\nTK\n(Eq 7b)\nwhere t is time during heating.\nEquation 5 provides the basis for selec-\ntion of a nominal aging time that will result\nin the desired yield strength and gives the\nfurnace operator a method of compensating\nfor heating rate and for differences between\ndesired and attained soak temperatures.\nSpecifics will be illustrated using data for\nalloy 7050. The value of Fys (in units of\nhours) for 7050 can be calculated by the\nfollowing equation:\nFys 1.45\u00d710-16 exp (\n32 562\nTF+460\n)\n(Eq 7a)\nwhere T is temperature in \u00b0F, or\nwhere Tk is temperature in K.\nIn one experiment, lengths of 7050-W (4\ndays) extrusions were aged at 24 h at 120\n\u00b0C (250 \u00b0F) plus the equivalent of 3 to 42 h\nat 165 \u00b0C (325 \u00b0F). For the second step, a\nlogarithmic heatup was used in which 10 h\nwere required for the load to reach 155 \u00b0C\n(315 \u00b0F), and nominal soak temperature\nwas 165 \u00b0C (325 \u00b0F). Figure 24 indicates\nthat yield strength generally agreed with\nvalues predicted using Eq 5. The deviation\nof the curve for short-transverse strength\nat the short aging times indicates that the\nmethod is inadequate for predicting\nstrength on the underaging side of the\naging curve.\nThe effects of neglecting to compensate\nfor soaking at temperatures other than the\nnominal can be large (Fig 25). For exam-\nple, the calculated difference in strength\nbetween alloy 7050 extrusions soaked 29 h\nat 160 \u00b0C (320 \u00b0F) and at 165 \u00b0C (325 \u00b0F) is\nabout 50 MPa (7 ksi), and the calculated\ndifference in strength between 7050 extru-\nsions soaked 29 h at 155 \u00b0C (315 \u00b0F) and at\n170 \u00b0C (335 \u00b0F) is about 100 MPa (14 ksi).\nNeglecting to compensate for time spent\nheating the work to the soak temperature\nwill increase the variability. Strength loss\nattributed to heatup was 14 MPa (2 ksi).\nThese kinetic relationships also can as-\nsist in selection of equivalent aging times\nfor alternate second-step aging tempera-\ntures. Equations 5 and 7 can be rearranged\nElongation, % in 50 mm or 2 in.\nYield strength, MPa\nTensile strength, MPa\n864 / Heat Treating of Nonferrous Alloys\n50\n500\n450\n400\n350\n300\n260 \"C\n(500 \u00b0F)\n250\n200\n0\n0.01\n0.1\n500\n400\n550\n500\n105 C\nRT\n(225 F) -70\n450\n60\n135 \"C\n(275 F)\n150 \u00b0C\n(300 \"F)\n175 \"C\n(350 (F)\n190 \"C\n(375 \u00b0F)\n205 \"C\n230 \"C\n(450 \"F)\n(400 \"F)\nT\nTensile strength, ksi\nTensile strength, MPa\n400\n350\n300\n40\n250\n\u2610\n100 C\n(212 F)\n150 C\nNW(320F) (300)\n190 C\n(375 \u00b0F)\nIT\n70\n50\n60\nTensile strength, ksi\n205 C\n260 \u00b0C\n(500 \u02daF)\n40\n(400 \u00b0F)\n30 min\n30 min\n1 day\n1\n10\n1 day 1 week 2 months 1 year\n1 week 2 months 1 year.\n30\n00\n200\n1\n100\n103\n104\n105\n0\n0.01\n0.1\n1\n10\n100\n103\n1|\n104\n30\n105\nDuration of precipitation heat treatment, h\nAging time, h\n500\n70\n105 \u00b0C\n(225\u00b0F)\n60\n+\n135\u00b0C\n400\nRT\n(275 \u00b0F)\n50\n175 C\n(350 \u00b0F)\n130 \u00b0C (265\u00b0F) 150 \u00b0C\n(300 \u02daF)\n70\n10\n60\n60\n300\n150 \u00b0C\n40\n(300 \u00b0F)\n200\n175 \u00b0C (350 \u00b0F) - 30\n190 \u00b0C (375 \u00b0F)\nYield strength, ksi\nYield strength, MPa\n100\u00b0C\n(212 \"F)\n50\n300\n190 \u00b0C\n(375 \"F)\n160 \u00b0C\n(320 \u00b0F)\n200\n40\n40\n-205 \u00b0C (400 \u00b0F)\n30\n205 C\n(400 F)\n20\n260 \u00b0C (500 \u00b0F).\n100\n20\n230 \u00b0C (450 \u00b0F).\n100\n00\n260 \u02daC\n10\n10\n(500 \"F)\n30 min\n1 day 1 week 2 months 1 year\n30 min\n1 day 1 week 2 months 1 year\n0\n0\n0.01\n0.1\n1\n10\n100\n11\n103\n1\n0\n1\n\u2610\n\u2610\n0\n0\n104\n105\n0\n0.01\n0.1\n1\n10\n100\n103\n104\n105\nDuration of precipitation heat treatment, h\nAging time, h\n40\n30 205 \u00b0C\n190 \u00b0C\n(375 \u00b0F)\n(400 \u00b0F)\n175 \u00b0C\n(350 \u00b0F)\n150 \"C 135 C\n(300 \u02daF) (275 \u00b0F)\n20\nElongation, % in 50 mm or 2 in.\n20\nRT\n15\n10\n260 \u00b0C\n(500 \u00b0F)\n100 C\n(212 \u00b0F)\n130 C\n(265\u00b0F)\n205 C\n(400 F)\n150 C\n(300 F)\n10\n105 C\n(225 FI\n5\n190 \u00b0C 1\n(375 \"F)\n160 \u00b0C\n260 \u00b0C\n(500 \u00b0F)\n230 C\n(450 \u00b0F)\n30 min\n1 day\n0\n1 week 2 months 1 year\n\u2610 1\n30 min\n\u2020175 C\n(350 F) (320 \u02daF)\n1 day 1 week 2 months 1 year\n0\n0.01\n0.1\n1\n10\n100\n103\n\u2610\n104\n0\n1\n1\n105\n0\n0.01\n0.1\n1\n10\n100\n11\n103\n||\n104\n105\nAging time, h\nDuration of precipitation heat treatment, h\nFig 22(a) Aging characteristics of alloy 2014 sheet\nto yield the following equation:\nt2=t\u2081 exp\n32 562 32 562\nT\u2081+460 T\u2082+460\n(Eq 8)\nwhere 11 is aging time at temperature T\u2081, t\u2082 is\naging time at temperature T\u2082 that will provide\nequivalent yield strength, and T, and T\u2082 are in\n\u00b0F. For example, the time at 175 \u00b0C (350 \u00b0F)\nequivalent to aging alloy 7050 for 29 h at 165\n\u00b0C (325 \u00b0F) is calculated as follows:\nFig 22(b) Aging characteristics of alloy 2024 sheet (see also Fig 22d)\n1350 29/exp (1.28)=29/3.6=8 h\nThermomechanical effects on aging occur\nfrom deformation after solution heat treat-\nment. The deformation step may be warm\nor cold and before, after, or during aging.\nThe simplest thermomechanical practices\nare those of the conventional T3, T8, or T9\ntempers. The rate and extent of precipita-\ntion strengthening are distinctly increased\nin some alloys by cold working after\nquenching, whereas other alloys show little\nor no added strengthening when treated by\nthis sequence of operations.\nAlloys of the 2xxx series such as 2014,\n2124, and 2219 are particularly responsive\nto cold work between quenching and aging,\nand this characteristic is the basis for the\nhigher-strength T8 tempers. The strength\nimprovement accruing from the combina-\ntion of cold working and precipitation heat\ntreating is a result of nucleation of addition-\nYield strength, ksi\nTensile strength, MPa\nYield strength, MPa\nElongation, % in 50 mm or 2 in.\n350\n325\n(250 \u00b0F)\n300\n275\n250\n225\n200\n150 \u00b0C\n(300 \u00b0F)\n170 \u00b0C\n(340 \u00b0F)\n205 \u00b0C\n(400 \u00b0F)\n230 \u00b0C\n(450 \u00b0F)\n260 \u00b0C\n(500 \u00b0F)\n30 min\n1 day\n175\n1 week 2 months 1 year\n\u2610 1\nJ\n0\n0.01\n0.1\n1\n10\n100\n103\n104\n105\n300\n250\n200\n150\n100\nDuration of precipitation heat treatment, h\n50\n0\n0.01\n0.1\n40\n445\n40\n35\n35\n30\n40\n120 \u00b0C\n(250 \u00b0F)\n150 \u00b0C\n(300 \u00b0F)\n35\n170 \u00b0C\n(340 \u00b0C)\n30\n205 \u00b0C\n(400 \u00b0F)\n1\n230 \u00b0C\n(450 \u00b0F)\n25\n20\n20\n15\n260 \u00b0C\n(500 \u00b0F)\n30 min\n1 day\n1 week 2 months 1 year\n10\n103\n104\n105\n1\n10\n100\nDuration of precipitation heat treatment, h\n30\n230 \u00b0C 205 \u00b0C\n(450 \u00b0F) (400 \u00b0F)\n170 \u00b0C\n(340 \u00b0F)\n_ 150 \u00b0C\n(300 \u00b0F)\n260 \u00b0C\n(500 \u00b0F)\n120\u00b0C\n(250 \u00b0F)\n201\n10-260 \u00b0C\n(500 \u00b0F)\n0\n0\n0.01\n0.1\n30 min\n10\n1 day 1 week 2 months 1 year\n100\n1\n103\nDuration of precipitation heat treatment, h\nFig 22(c) Aging characteristics of alloy 6061 sheet\nal precipitate particles by the increased\nstrain. In some alloys of the 2xxx series,\nstrain introduced by cold working after so-\nlution heat treatment and quenching also\ninduces nucleation of a finer precipitate\ndispersion that increases strength. Depend-\ning on the aging temper, however, tough-\nness may be adversely affected, as illustrat-\ned in Fig 26 for 2024 sheet.\nStrengthening from thermomechanical\nprocessing is the basis for the higher-\n11\n104\n105\nTensile strength, ksi\nYield strength, ksi\nstrength T8-type tempers of alloys 2011,\n2024, 2124, 2219, and 2419, which are\nproduced by applying controlled amounts\nof cold rolling, stretching, or combinations\nof these operations. Normally, cold work\nis introduced by stretching; however, oth-\ner methods such as cold rolling can be\nused. Recently, 2324-T39 was developed.\nThe T39 temper is obtained by cold rolling\napproximately 10% after quenching fol-\nlowed by stretching to stress relieve. This\nHeat Treating of Aluminum Alloys / 865\ntype of approach results in strengths sim-\nilar to those obtained with T8 processing\nbut with the better toughness and fatigue\ncharacteristics of T3 products. Alloys\n2024, 2124, and 2219 in T8-type tempers\nare particularly well suited for supersonic\nand military aircraft; alloy 2219 in such\ntempers, and alloy 2014-T65, were the\nprincipal materials for the fuel and oxidiz-\ner tanks (which also served as the primary\nstructure) of the Saturn V space vehicles.\nRe-solution heat treatment of mill prod-\nucts supplied in these tempers can result in\ngrain growth and in substantially lower\nstrength than is normal for the original\ntemper. Such reheat treatment is not rec-\nommended.\nAlloys of the 7xxx series do not respond\nfavorably to the sequence of operations\nused to produce T8-type tempers, and no\nsuch tempers are standard for these alloys.\nThe strains associated with stretching or\ncompressing of 7xxx alloys have relatively\nlittle effect on the mechanical properties of\nmaterial precipitation heat treated to T6-\ntype tempers. On the other hand, these\noperations have measurable detrimental\neffects on final strength when T73-, T736,\nor T76-type tempers are produced, partic-\nularly in the direction opposite the direc-\ntion of cold work. Accordingly, specifica-\ntion properties are somewhat lower for\nthe stress-relieved versions of these tem-\npers. Decreasing the overaging time to\ncompensate for the loss in strength is not\nadvisable, because this would impair de-\nvelopment of the desired corrosion char-\nacteristics.\nTemperature control and uniformity pre-\nsent essentially the same problems in pre-\ncipitation heat treating as they do in solu-\ntion heat treating.\nGood temperature control and uniformity\nthroughout the furnace and load are re-\nquired for all precipitation heat treating.\nRecommended temperatures are generally\nthose that are least critical and that can be\nused with practical time cycles. Except for\n7xxx alloys in T7x tempers, these tempera-\ntures generally allow some latitude and\nshould have a high probability of meeting\nproperty specification requirements. Fur-\nnace radiation effects seldom are trouble-\nsome except in those few furnaces that are\nused for both solution and precipitation heat\ntreating. Generally, such situations should\nbe avoided, because the high heat capacity\nneeded for the higher temperatures may be\ndifficult to control at normal aging temper-\natures.\nSoak time in precipitation heat treating is\nnot difficult to control; the specified times\ncarry rather broad tolerances. Heavier loads\nwith parts racked closer together, and even\nnested, are not abnormal. The principal haz-\nard is undersoaking due to gross excesses in\nloading practices. Some regions of the load\nmay reach soak temperature long after soak\n866 / Heat Treating of Nonferrous Alloys\nElongation, % in\n50 mm or 2 in.\nYield strength, MPa\nTensile strength, MPa\n600\n2024-T4\n(not cold worked)\n80\n500\n190 \u00b0C (375 \u00b0F).\n70\n205 \u00b0C (400 \u00b0F)\n60\n60\n400\n300\n600\n500\n220 \u00b0C (425 \u00b0F) - 50\nTensile strength, ksi\nTensile strength, MPa\n600\n2024-T3\n(cold worked 1 to 2%)\n500\n400\n220 \u00b0C (425 \u00b0F)\n80\n190 \u00b0C (375 \u00b0F)\n205 \u00b0C (400 \u00b0F)\n70\n60\n50\n50\nTensile strength, ksi\nTensile strength, MPa\n600\n2024-T36\n(cold worked 5 to 6%)\n190 \u00b0C (375 \u00b0F) 175 \u00b0C (350 \u00b0F)-80\n000\n500\n400\nF205 \u00b0C (400 \u00b0F)\n70\n220 \u00b0C (425 \u00b0F)\n300\n300\n0\n4\n8\n12\n16\n0\n4\n8\n12\n16\n0\n4\nAging time, h\nAging time, h\n80\n70\n60\n400\n190 \u00b0C (375 \u00b0F)-\n300\n205 \u00b0C (400 \u00b0F)\n40\n220 \u00b0C (425 \u00b0F)\n200\n0\n4\n30\n8\nAging time, hi\nYield strength, ksi\nYield strength, MPa\n600\n500\n600\n80\n80\n400\n190 \u00b0C (375 \u00b0F)\n205 \u00b0C (400 \u00b0F)\n220 \u00b0C (425 \u00b0F)\nT\n70\n60\n50\n300\n30\n200\n12\n16\n0\n4\n190 \u00b0C (375 \u00b0F)\n205 \u00b0C (400 \u00b0F)\n220 \u00b0C (425 \u00b0F)\n0\n4\n8\nElongation, % in\n50 mm or 2 in.\n30\n8\nAging time, h\n205 \u00b0C (400 \u00b0F)\n50\nYield strength, ksi\nYield strength, MPa\n- 40\n8\nAging time, h\nT\n12\n16\n190 \u00b0C (375 \u00b0F)\n175 \u00b0C (350 \u00b0F) - 80\n500\n70\n205 \u00b0C (400 \u00b0F)\n60\n400\n220 \u00b0C (425 \u00b0F)\n300\nT\n50\n40\n30\n30\n200\n12\n16\n0\n4\n8\n12\n16\nAging time, h\n190 \u00b0C (375 \u00b0F)\n220 \u00b0C (425 \u00b0F)\n12\n16\n0\n4\n8\nElongation, % in\n50 mm or 2 in.\n30\n190 \u00b0C (375 \u00b0F)\n175 \u00b0C (350 \u00b0F)\n205 \u00b0C (400 \u00b0F)\n220 \u00b0C (425 \u00b0F)\n12\n16\n0\n4\n8\n12\n16\nAging time, h\n60\n60\n50\n50\nYield strength, ksi\nAging time, h\nAging time, h\nFig 22(d) Effects of cold work after quenching and before aging on tensile properties of alloy 2024 sheet\ntime has been called. Placement of load ther-\nmocouples is critical, and limiting the size and\nspacing of a load may be necessary for aging\nAging temperature, \u00b0C\n225\n200\n175\n60\n150\nto the T73 and T76 tempers. As discussed\nabove, soak time is not as critical for peak-\naged (T6 and T8) tempers.\n360\n385\n400\n410\nMPa\n440\n470\n350\n500\nT73 aging\n525\n550\n300\n64\n68\n125\nksi\n72-\n76\n80\n100\n0.01\n0.1\n1\n10\nAging time, h\n250\nT6 aging\n100\n1000\nHardening of Cast Alloys\nIn general, the principles and procedures\nfor heat treating wrought and cast alloys are\nsimilar. The major differences between so-\nlution-treating conditions for castings and\nthose for wrought products are found in\nsoak times and quenching media. Solution\nof the relatively large microconstituents\npresent in castings requires longer soaking\nAging temperature, \u00b0F\nYield strength, MPa\n700\n600\nLongitudinal\n500\nShort transverse\nPredicted from Eq 3\n\u043e\n\u2022 Actual\n400\n2\n4 6 8 10\nEquivalent aging time\n100\n90\n70\n88\n80\n]\n60\n20 30 40 60\nat 165 \u00b0C (325 \u00b0F), h\nYield strength, ksi\nFig 23 Iso-yield-strength curves for alloy 7075\nFig 24\nActual versus predicted yield strengths for\nalloy 7050 extrusions\nTensile strength, ksi\n100\nAging temperature, \u00b0F\n310\n320\n330\n50\n0\nEffect on yield strength, MPa\n50\n-50\nCompared for standard\n-100\nconditions of 29 h\nat 165 \u00b0C\n340\n10\n10\n5\nC\nG\nEffect on yield strength, ksi\nTear strength/Yield strength\n-15\n-150\n-20\n150\n155\n160 165 170 175\nAging temperature, \u00b0C\nEffect of aging temperature on yield strength\nFig 25 of alloy 7050-1736\nperiods than those used for wrought prod-\nucts (Table 3). When heat treatment of\ncastings must be repeated, solution times\nbecome similar to those for wrought prod-\nucts, because the gross solution and homog-\nenization has been accomplished and is\nirreversible under normal conditions. Re-\nduction of stresses and distortion from\nquenching are also important, because cast-\nings generally are complex shapes with vari-\nations in section thickness.\nDifferent casting processes and foundry\npractices also result in microstructural dif-\nferences with relevance to heat-treatment\npractice, because the coarser microstruc-\ntures associated with slow solidification\nrates require a longer solution heat treat-\nment exposure. Therefore, the time re-\nquired at temperature to achieve solution is\nprogressively shorter for investment, sand,\nand permanent mold castings. Foundry\npractice (chills, gating, type of mold) also\nplays an important role in the response of a\ncasting, or a portion of a casting, to heat\ntreatment. For example, thin-wall sand\ncastings produced with extensive use of\nchills can often display finer microstruc-\ntures than heavy-section permanent mold\nparts produced in such a way that process\nadvantages are not exploited.\nFor these reasons, solution heat-treat-\nment practices can be optimized for any\nspecific part to achieve solution with the\nshortest reasonable cycle once production\npractice is finalized, even though most\nfoundries and heat treaters will standardize\na practice with a large margin of safety.\nThere also exists a fundamental difference\nbetween unmodified and modified alloys in\nwhich the size and shape of silicon crystals\nare modified with additions of elements\nsuch as calcium, sodium, strontium, or an-\ntimony. Modified alloys undergo rapid\nspheroidization while complete sphe-\nroidization is not achieved in unmodified\nalloys even after very long times. The prac-\ntical implication is that shorter solution heat\ntreatment could be employed in fully mod-\nTransverse yield strength, ksi\n40 45 50 55 60 65 70\n1.8\n5000\nT4\nFe Si\nCu Mg Mn\n4.2 1.4 0.6 0.34 0.14\n4500\n1.6\nT3\nStretched 5%\nT7\n4000\n1.4\n1.2\nT6\nNot stretched\n1.0\n0.8\n250\nFig 26\n3500\n3000\n2500\nT7\nT8)\n2000\n300 350 400 450 500\nTransverse yield strength, MPa\nEstimated K, MPav mm\nEffect of stretching and aging on the tough-\nness and yield strength of 2024 sheet\nified castings. The microsegregation of sili-\ncon and magnesium is not severe in the\naluminum-silicon-magnesium casting al-\nloys, and hence it takes only a short time to\nhomogenize the alloy and to place the\nMg2Si into solution.\nQuenchants. Quenching of aluminum\ncastings is often done in boiling water or a\nmilder medium to reduce quenching stress-\nes in complex shapes. A commercially im-\nportant variety is a mixture of polyalkylene\nglycol and water, which has no detrimental\neffect on properties for thicknesses under\napproximately 3.2 mm (0.125 in.). Quen-\nchant additions can be made for the follow-\ning purposes:\n\u2022 To promote stable vapor film boiling by\nthe deposition of compounds on the sur-\nface of parts as they are submerged in the\nquench solution\n\u2022 To suppress variations in heat flux by\nincreasing vapor film boiling stability\nthrough chemically decreased quench so-\nlution surface tension\n\u2022 To moderate quench rate for a given water\ntemperature\nThe key to the compromise between\ngoals involving property development and\nthe physical consequences of quenching is\nuniformity of heat extraction, which is in\nturn a complex function of the operable heat\nextraction mechanism. Nucleate, vapor\nfilm, and convective boiling occur with dra-\nmatically different heat extraction rates at\ndifferent intervals. Differences in section\nthickness, load density, positioning, racking\nmethods, surface condition, and casting ge-\nometry also influence the results.\nProperty Development, Yield strength is\nlargely controlled by the limiting hardening-\nelement level, and tensile strength (in a\ngeneral sense) is related to the ductility at a\ngiven yield strength. Ductility, however, is\ncontrolled for a given yield strength by\nsoundness and microstructural fineness,\nand is thus determined in the foundry and\nHeat Treating of Aluminum Alloys / 867\nnot by the heat treater. This effect of casting\nmethods on property development is shown\nin Fig 27. Because of the finer cast structure\nand higher supersaturation of the more rap-\nidly solidified permanent mold castings,\ntheir tensile properties are superior to those\nof sand castings of the same composition\nsimilarly heat treated.\nTempers. Cast products of heat-treatable\naluminum alloys have the highest combina-\ntions of strength, ductility, and toughness\nwhen produced in T6-type tempers. Devel-\noping T6-type tempers in cast products re-\nquires the same sequence of operations\nemployed in developing tempers of the\nsame type in wrought products-solution\nheat treating, quenching, and precipitation\nheat treating. Premium-quality casting spec-\nifications such as MIL-A-21180 can require\ndifferent strengths and ductility levels in the\nsame casting.\nAmong precipitation treatments unique to\ncastings are those resulting in the T5 and T7\ntempers. The T5 temper is produced merely\nby applying a precipitation treatment to the\nas-cast casting, without previous solution\ntreatment. A moderate increase in strength\nis achieved without warpage and subse-\nquent straightening. High hardness and di-\nmensional and strength stability at elevated\ntemperatures account for the almost univer-\nsal use of materials in T5 tempers for pis-\ntons and other engine parts. Some applica-\ntions demand combinations of strength,\ntoughness, and dimensional stability that\ncannot be met by heat treating to T5-, T6-,\nor T8-type tempers. For these applications,\nT7-type tempers are developed by solution\nheat treating, quenching in a medium that\nprovides a moderate cooling rate, and then\nprecipitation heat treating at a temperature\nhigher than those used to develop T5-, T6-,\nand T8-type tempers. Heat treating to T7-\ntype tempers results in lower strength than\nthat of material in T6- or T8-type tempers,\ndevelops high ductility and toughness, and\ncarries precipitation far enough to minimize\nfurther precipitation during service.\nStress Relief\nImmediately after being quenched, most\naluminum alloys are nearly as ductile as\nthey are in the annealed condition. Conse-\nquently, it is often advantageous to stress\nrelieve parts by working the metal immedi-\nately after quenching. Numerous attempts\nalso have been made to develop a thermal\ntreatment that will remove, or appreciably\nreduce, quenching stresses. Normal precip-\nitation heat-treating temperatures are gener-\nally too low to provide appreciable stress\nrelief. Exposure to higher temperatures (at\nwhich stresses are relieved more effective-\nly) results in lower properties. However,\nsuch treatments are sometimes utilized\nwhen even moderate reduction of residual\nstress levels is important enough so that\nTensile strength, MPa\n868 / Heat Treating of Nonferrous Alloys\n350\n300\nPermanent mold castings\nSand castings\n350\n50\n150 \u00b0C (300\u00b0F)\n40\n250\n205 \u00b0C (400)\n\u00b0F)-\n200\n150\n260 \u00b0C (500 \u00b0F)\n20\nTensile strength, ksi\nTensile strength, MPa\n300\n250\n200\n150 \u00b0C (300 \u00b0F)\n40\n205 \u00b0C (400 \u00b0F)\n30\n150\n260 \u00b0C (500 \u00b0F)\n20\n7079\n50\nSpar forging\nTensile strength, ksi\nB\nA\nPlanes of\nsaw cuts\n-160 mm\nB\nA\nDirection of\nYield strength, MPa\nElongation, %\nin 50 mm or 2 in.\n100\n100\n0\n10\n20\n30\n0\n10\nAging time, h\n250\n200\n30\n-150 \u00b0C (300 \u00b0F)-\n150\n205 \u00b0C (400 \u00b0F)\n260 \u00b0C (500 \u00b0F)\n100\n20\n10\n50\n50\n0\n10\n20\n30\nAging time, h\n2\n1\n260 \u00b0C (500 \u00b0F)\n205 \u00b0C (400 \u00b0F)\n10\nAging time, h\n150 \u00b0C (300 \u00b0F).\n20\n30\nYield strength, ksi\nYield strength, MPa\nElongation, %\nin 50 mm or 2 in.\nAging time, h\n250\n200\n150 \u00b0C (300 \u00b0F)\n150\n150\n100\n50\n1\n20\n30\n-205 \u00b0C (400 \u00b0F):\n260 \u00b0C (500 \u00b0F)\nT\n20\n10\n0\n10\n20\n30\nAging time, h\nYield strength, ksi\nDeflection, \u03bcm\nDeflection, um\ncompressive deformation\n7500\nSaw cut A-A\nT6 parallel\n250\n200\n4500\nT6 normal\n150\n100\n1500\n50\n0\n-1500\nT652 parallel\nT652 normal\n-50\nLength, 1.8 m\n2500\nSaw cut B-B\nT6 parallel\n80\nT6 normal\n1500\n60\n40\nT652 normal\n500\nT652 parallel\n20\n0\n-20\n205 \u00b0C (400 \u00b0F)\n260 \u00b0C (500 \u00b0F)\n-500\n-150 \u00b0C (300 \u00b0FH\n10\n20\n30\nAging time, h\nFig 27 Comparison of the precipitation-hardening characteristics of 356.0-T4 sand and permanent mold\ncastings\nsome sacrifice in mechanical properties can\nbe accepted. The T7 temper for castings is a\ntypical example of this kind of treatment.\nMechanical Stress Relief. Deformation\nconsists of stretching (bar, extrusions, and\nplate) or compressing (forgings) the product\nsufficiently to achieve a small but controlled\namount (1 to 3%) of plastic deformation. If\nthe benefits of mechanical stress relieving\nare needed, the user should refrain from\nreheat treating.\nFigure 28 illustrates the beneficial effect\nof 3% permanent deformation in compres-\nsion on a large forging.\nThese methods are most readily adapt-\nable to mill and forge shop products and\nrequire equipment of greater capacity than\nthat found in most manufacturing plants.\nApplication of these methods to die forgings\nand extrusions usually requires construc-\ntion of special dies and jaws. Stretching\ngenerally is limited to material of uniform\ncross section; however, it has been applied\nsuccessfully to stepped extrusions and to a\n3 by 14 m (10 by 47 ft) aircraft wing skin\nroll-tapered to a thickness range of 7.1 to\n3.2 mm (0.280 to 0.125 in.).\nSpecific combinations of the supplemen-\ntal digits are used to denote the tempers\nproduced when mechanical deformation is\nused primarily to relieve residual stresses\ninduced during the quenching operation.\nFor products stress relieved by stretching,\nthe digits 51 follow the basic Tx designation\n(T451, for example). For products stress\nrelieved by compressive deformation, the\nsupplementary digits are 52.\nAn additional digit is added to designa-\ntions for extrusions: an added zero specifies\nthat the product has not been straightened\nafter final stretching; an added one indicates\nthat straightening may have been performed\nafter final stretching.\nEffect of Precipitation Heat Treating on\nResidual Stress. The stresses developed dur-\ning quenching from solution heat treatment\nare reduced during subsequent precipitation\nheat treatment. The degree of relaxation of\nstresses is highly dependent upon the time\nand temperature of the precipitation treat-\nment and the alloy composition. In general,\nthe precipitation treatments used to obtain\nthe T6 tempers provide only modest reduc-\ntion in stresses, ranging from about 10 to\n35%. To achieve a substantial lowering of\nquenching stresses by thermal stress relax-\nation, higher-temperature treatments of the\nT7 type are required. These treatments are\nused when the lower strengths resulting\nfrom overaging are acceptable.\nOther thermal stress-relief treatments,\nknown as subzero treatment and cold stabi-\nLength, -1.8 m\nDeflection, 0.001 in.\nDeflection, 0.001 in.\nEffect of 3% permanent deformation in com-\nFig 28\npression (T652 treatment) on distribution of\nstress in a large forging. Parallel and normal refer to\nwarpage directions with respect to the plane of the\nsaw cut.\nlization, involve cycling of parts above and\nbelow room temperature. The temperatures\nchosen are those that can be readily ob-\ntained with boiling water and mixtures of\ndry ice and alcohol-namely, 100 and -73\n\u00b0C (212 and -100 \u00b0F)-and the number of\ncycles ranges from one to five. The maxi-\nmum reduction in residual stress that can be\neffected by these techniques is about 25%.\nThe maximum effect can be obtained only if\nthe subzero step is performed first, and\nimmediately after quenching from the solu-\ntion-treating temperature while yield\nstrength is low. No benefit is gained from\nmore than one cycle.\nA 25% reduction in residual stress is\nsometimes sufficient to permit fabrication\nof a part that could not be made without this\nreduction. However, if a general reduction\nis needed, as much as 83% relief of residual\nstress is possible by increasing the severity\nof the uphill quench\u2014that is, more closely\napproximating the reverse of the cooling-\nrate differential during the original quench.\nThis may be accomplished by a patented\nprocess that involves extending the subzero\nstep to 195 \u00b0C (-320 \u00b0F) and then very\nrapidly uphill quenching in a blast of live\nsteam (Fig 29). The rate of reheating is\nextremely critical, and therefore, to ensure\nproper application of the steam blast, a\nPlate\nthickness\n50 mm\nHeat Treating of Aluminum Alloys / 869\nTable 9 Reheating schedules for wrought aluminum alloys\nThe schedules given in this table normally will not decrease strength more than 5%.\nReheating time at a temperature of:\n150 \u00b0C\n165 \u00b0C\nAlloy and temper\n(300 \u00b0F)\n(325 \u00b0F)\n175 \u00b0C\n(350 \u00b0F)\n190 \u00b0C\n(375 \u00b0F)\n205 \u00b0C\n(400 \u00b0F)\n220 \u00b0C\n(425 \u00b0F)\n230 \u00b0C\n(450 \u00b0F)\nA B C D\nRange of\nresidual\n2014-T4\n(a)\n(a)\n(a)\n(a)\n(a)\n(a)\n(a)\nstress\nMPa ksi\n2014-T6\n2-50 h\n8-10 h\n2-4 h\n1/2-1 h\n5-15 min\n(b)\n(b)\n2024-T3, 2024-T4\n(a)\n(a)\nA 28\nB 86\n4.0\n12.4\n2024-T81, 2024-T86\n20-40 h\n(a)\n2-4 h\n(a)\n(a)\n(a)\n(a)\n1 h\n12 h\n15 min\n5 min\n6061-T6, 6062-T6,\nC 130\n6063-T6\n19.0\nD 165\n24.0\n7075-T6, 7178-T6\n100-200 h\n10-12 h\n50-100 h\n1-2 h\n8-10 h\n1-2 h\n1-2 h\n1/2-1 h\n1/2 h\n5-10 min\n15 min\n(b)\n5 min\n(a)\n(a) Reheating not recommended. (b) Bring to temperature\nFig 29\nTreatment\nA: Cooled to -195 \u00b0C, then uphill\nquenched in a steam blast\nB: Cooled to -75 \u00b0 C, then uphill\nquenched in a steam blast\nC: Cooled to -75 or -195 \u00b0 C, then\nuphill quenched in boiling water\nD: Standard specimen, quenched and\naged to T6 temper in conventional\nmanner with no further treatment\nEffectiveness of various uphill quenching\ntreatments in reducing residual quenching\nstresses in 2014 plate. Note: uphill quenching treat-\nments (single-cycle only) were applied from 1/2 to 11/2 h\nafter quenching from the recommended solution-\ntreating temperature. All specimens were aged to the\nT6 temper after uphill quenching.\nspecial fixture usually is required for each\npart.\nThis process will not solve all problems of\nwarpage in machining. It may reduce\nwarpage internally but increase warpage of\nthe extreme outer layers, although in the\nopposite direction (Fig 30). Also, the effect\nof the altered residual-stress pattern on per-\nformance must be evaluated carefully for\neach part. This is particularly important for\nparts subjected to cyclic loading or exposed\nto corrosive environments such as marine\natmospheres, especially if the process is\nintroduced after the start of production and\noriginal performance tests are not repeated.\nFurther disadvantages are the cost and haz-\nard involved in handling liquid nitrogen and\nlive steam.\nEffects of Reheating\nThe precipitation characteristics of alumi-\nnum alloys must be considered frequently\nduring evaluation of the effects of reheating\non mechanical properties and corrosion re-\nsistance. Such evaluations are necessary for\ndetermining standard practices for manu-\nfacturing operations, such as hot forming\nand straightening, adhesive bonding, and\npaint and dry-film lubricant curing, and for\nevaluating the effects of both short-term\nand long-term exposure in elevated temper-\natures in service.\nThe stage of precipitation that exists in an\nalloy at the time of reheating plays a signif-\nicant role in the effects of reheating. Con-\nsequently, it is extremely dangerous to re-\nheat material in a solution heat-treated\ntemper without first carefully testing the\neffects of such reheating. In one such test,\n2024-T4 sheet was found to be very suscep-\ntible to intergranular corrosion when sub-\njected to a 15-min drying operation at 150 \u00b0C\n(300 \u00b0F) during the first 8 h after quenching;\nno susceptibility was evident when the\nsame drying operation was performed more\nthan 16 h after quenching. In another test,\n7075-W (0.2 to 600 h) bar and plate were\nreheated for hot forming at 175 \u00b0C (350 \u00b0F)\nfor 20 min. Strengths after aging to the T6\ntemper were 10 to 15% lower than those for\nstandard 7075-T6. In contrast, similar re-\nheating of T6 material for up to 1 h at 175 \u00b0C\n(350 \u00b0F) produced no detrimental effect.\nIf reheating is performed on material in\nthe W or T4 condition, its effect can be\nestimated from families of precipitation\nheat-treating curves such as those present-\ned in Fig 22. Such curves can also be used\nfor reheating of precipitation heat-treated\nmaterial at the precipitation heat-treating\ntemperature. For reheating at other temper-\natures, other data may be needed (Fig 31).\nThe heat-treating and reheating curves may\nbe used as the bases for limitations on\nreheating (Table 9).\nAnnealing\nAnnealing treatments employed for alu-\nminum alloys are of several types that differ\nin objective. Annealing times and tempera-\ntures depend on alloy type as well as on\ninitial structure and temper.\nFull Annealing. The softest, most ductile,\nand most workable condition of both non-\nheat-treatable and heat-treatable wrought\nalloys is produced by full annealing to the\ntemper designated \u201cO.\u201d Strain-hardened\nproducts in this temper normally become\nrecrystallized, but hot-worked products\nmay remain unrecrystallized. In the case of\nheat-treatable alloys, the solutes are suffi-\nciently thoroughly precipitated to prevent\nnatural age hardening. A higher maximum\ntemperature than that used for stress-relief\nannealing, controlled cooling to a lower\ntemperature, and additional holding time at\nthe lower temperature generally are em-\nployed.\nFor both heat-treatable and non-heat-treat-\nable aluminum alloys, reduction or elimina-\ntion of the strengthening effects of cold work-\ning is accomplished by heating at a\ntemperature from about 260 to about 440 \u00b0C\n(500 to 825 \u00b0F). The rate of softening is\nstrongly temperature-dependent; the time re-\nquired to soften a given material by a given\namount can vary from hours at low tempera-\ntures to seconds at high temperatures.\nIf the purpose of annealing is merely to\nremove the effects of strain hardening, heat-\ning to about 345 \u00b0C (650 \u00b0F) will usually\nsuffice. If it is necessary to remove the\nhardening effects of a heat treatment or of\ncooling from hot-working temperatures, a\ntreatment designed to produce a coarse,\nwidely spaced precipitate is employed. This\nusually consists of soaking at 415 to 440 \u00b0C\n(775 to 825 \u00b0F) followed by slow cooling (28\n\u00b0C/h, or 50 \u00b0F/h, max) to about 260 \u00b0C (500\n\u00b0F). The high diffusion rates that exist dur-\ning soaking and slow cooling permit maxi-\nmum coalescence of precipitate particles\nand result in minimum hardness.\nAs a result of this treatment, only partial\nprecipitation occurs in 7xxx alloys, and a\nsecond treatment (soaking at 230 \u00b1 6 \u00b0C, or\n450 \u00b1 10 \u00b0F, for 2 h) is required. When the\nneed arises for small additional improve-\nments in formability, cooling at 28 \u00b0C/h (50\n\u00b0F/h) should be extended to 230 \u00b0C (450 \u00b0F),\nand the material should be soaked at 230 \u00b0C\nfor 6 h. The effects of eliminating or pro-\nlonging the 230 \u00b0C second step on the duc-\ntility of 7075-0 sheet are compared with the\nstandard treatment in Table 10.\nIn annealing, it is important to ensure that\nthe proper temperature is reached in all\nportions of the load; therefore, it is common\nto specify a soaking period of at least 1 h.\nThe maximum annealing temperature is\nmoderately critical; it is advisable not to\nexceed 415 \u00b0C (775 \u00b0F), because of oxidation\nand grain growth. The heating rate can be\ncritical, especially for alloy 3003, which\nusually requires rapid heating for preven-\ntion of grain growth. Relatively slow cool-\ning, in still air or in the furnace, is recom-\nmended for all alloys to minimize distortion.\nTypical annealing conditions used for some\nalloys in common use are listed in Table 11.\n870 / Heat Treating of Nonferrous Alloys\nDeflection, m\n-1300\n-2600\n2600\n+1300\n+100\n25-by-25-mm (1-by-1-in.) bar\n0\n0\n-100\n-3900\n-5200\n7075\no Control specimen\nQuenched from 75 \u00b0C -200\n(-100 \u00b0F) to steam\nA Quenched from liquid\nnitrogen to steam\n-6500\n1\n2\n3\n4\nDeflection, 0.001 in.\nTensile strength, MPa\n475\n450\n65\n425\nO 120 \u00b0C\n60\n400\n\u2022 135 \u00b0C\nReheating\nA 150 \u00b0C\n375\ntemperature\n55\n99\nTensile strength, ksi\n425\n60\n400\n375\n55\n350\n50\nO 120 \u00b0C\n325\n135 \u00b0C\nYield strength, ksi\nA 150 \u00b0C\n1\n300\nElongation in\n50 mm (2 in), %\nYield strength, MPa\n10\n0\n10\n100\n103\nDeflection, um\nDeflection, um\n+1300\n50-by-50-mm\nTine number\n50\n+50\n0\n0\n(2-by-2-in.) bar\n-50\n-1300\n-2600\n50\n-100\n7075\n-150\n-3900\n\u3002 Control specimen\n\u2022 Quenched from -75 \u00b0C (-100 \u00b0F) to steam\n-5200\n2600\n+1300\n0\n-1300\n-2600\nA Quenched from liquid nitrogen to steam\n2\n3\n4\nTine number\n-3900\n1\n2\n3\n-200\n5\n6\nDeflection, 0.001 in.\n1 2 3 4 5 6\n100\n75-by-75-mm (3-by-3-in.) bar\n7075\no Control specimen\nQuenched from -75 \u00b0C (-100 \u00b0F) to steam\nA Quenched from liquid nitrogen to steam\n4\n5\nTine number\n6\n7\n500\n+50\n-50\n-100\nA-150\n8\nDeflection, 0.001 in.\nEffect of uphill quenching on deflection of tines. Six-tine specimen was machined from 50 by 50 mm\nFig 30\n(2 by 2 in.) bar. Similar specimens machined from 25 by 25 mm (1 by 1 in.) and 75 by 75 mm (3 by 3 in.)\nbars had four and eight tines, respectively.\nProducts that can be heated and cooled\nvery rapidly, such as wire, are annealed by\ncontinuous processes that require a total\nheating and cooling time of only a few\nseconds. Continuous annealing of coiled\nsheet is accomplished in a total time of a\nfew minutes. For these extremely rapid\noperations, maximum temperature may ex-\nceed 440 \u00b0C (825 \u00b0F).\nAlthough material annealed from the pre-\ncipitation-hardened condition usually has\nsufficient ductility for most forming opera-\nFig 31\nDuration of reheating, days\nEffects of reheating on tensile properties of\nalclad 2024-T81 sheet\ntions, this ductility often is slightly lower\nthan that of material that has not been\nsubjected to prior heat treatment\u2014that is,\nmaterial annealed at the producing source.\nTherefore, when maximum ductility is re-\nquired, annealing of a previously heat-treat-\ned product is sometimes unsuccessful.\nPartial Annealing. Annealing of cold-\nworked non-heat-treatable wrought alloys to\nobtain intermediate mechanical properties\n(H2-type tempers) is referred to as partial\nannealing or recovery annealing. Tempera-\ntures used are below those that produce ex-\ntensive recrystallization, and incomplete soft-\nening is accomplished by substructural\nchanges in dislocation density and rearrange-\nment into cellular patterns (polygonization).\nBendability and formability of an alloy an-\nnealed to an H2-type temper generally are\nsignificantly higher than those of the same\nalloy in which an equal strength level is de-\nveloped by a final cold-working operation\n(H1-type temper). Treatments to produce H2-\ntype tempers require close control of temper-\nature to achieve uniform and consistent me-\nchanical properties.\nFigure 32 shows changes in yield strength\nas functions of temperature and time for\nsheet of two non-heat-treatable alloys (1100\nand 5052) initially in the highly cold-worked\ncondition (H18 temper). From these curves,\nit is apparent that, by selection of appropri-\nate combinations of time and temperature,\nmechanical properties intermediate to those\nof cold-worked and fully annealed material\ncan be obtained. It is also evident that yield\nstrength depends much more strongly on\ntemperature than on time of heating.\nStress-Relief Annealing. For cold-worked\nwrought alloys, annealing merely to remove\nthe effects of strain hardening is referred to\nas stress-relief annealing. Such treatments\n\n\nTable 10 Effects of annealing treatments on ductility of 7075-0 sheet\nAnnealing\ntreatment\nfor thickness of:\nElongation in bending(c), %\nin 50 mm (2 in.) for\nthickness of:\nElongation in tension(a), % in 50 mm (2 in.)\nBend angle(b), degrees, for\nthickness of:\n0.5 mm\n(0.020 in.)\n1.6 mm\n(0.064 in.)\n2.6 mm\n(0.102 in.)\n1.6 mm\n(0.064 in.)\n2.6 mm\n(0.102 in.)\n1.6 mm\n(0.064 in.)\nTreatment 1(d)\n12\n12\nTreatment 2(e)\n14\n12\n14\n82\n73\n48\nTreatment 3(f)\n16\n91\n92.5\n58\n56\n14\n16\n76\n84\n2.6 mm\n(0.102 in.)\n50\n57\n60\n(a) Uniform elongation of gridded tension specimens. (b) Bend angle at first fracture. (c) Elongation in bend test for 1.3 mm (0.05 in.)\ngage spanning fracture. (d) Soak 2 h at 415 14 \u00b0C (775 \u00b1 25 \u00b0F); furnace cool to 260 \u00b0C (500 \u00b0F) at 30 \u00b0C/h (50 \u00b0F/h); air cool. (c) Soak\n2 h at 425 \u00b0C (800 \u00b0F), air cool; soak 2 h at 230 \u00b0C (450 \u00b0F), air cool. (f) Soak 1 h at 425 \u00b0C (800 \u00b0F); furnace cool to 230 \u00b0C (450 \u00b0F) at\n30 \u00b0C/h (50 \u00b0F/h); soak 6 h at 230 \u00b0C (450 \u00b0F), air cool\nemploy temperatures up to about 345 \u00b0C\n(650 \u00b0F), or up to 400 \u00b1 8 \u00b0C (750 \u00b1 15 \u00b0F)\nfor 3003 alloy, and cooling to room temper-\nature. No appreciable holding time is re-\nquired. Such treatment may result in simple\nrecovery, partial recrystallization, or full\nrecrystallization. Age hardening may follow\nstress-relief annealing of heat-treatable al-\nloys, however, because a concentration of\nsoluble alloying elements sufficient to cause\nnatural aging remains in solid solution after\nsuch treatments.\nA special form of stress-relief temper is\nused for heat-treatable alloy products that\nsubsequently will be inspected ultrasonical-\nly. The product is heated to its normal\nsolution heat-treating temperature, then\ncooled in still air to room temperature. This\ntemper is referred to as the O1 temper.\nControlled-Atmosphere Annealing and\nStabilizing. Aluminum alloys that contain\nTable 11\neven very small amounts of magnesium will\nform a surface magnesium oxide unless the\natmosphere in the annealing furnace is free\nof moisture and oxygen. Examples include\nalloy 3004, which is used for cooking uten-\nsils, and alloys of the 5xxx series.\nAnother problem that control of the an-\nnealing atmosphere helps to overcome or\navoid is oil staining by oil-base roll lubri-\ncants that do not burn off at lower annealing\ntemperatures. If the oxygen content of the\nfurnace atmosphere is kept very low during\nsuch annealing, the oil will not oxidize and\nstain the work.\nTemperature control for full and partial\nannealing is somewhat more critical than for\nstress-relief annealing; the temperatures\nand times specified are selected to produce\nrecrystallization and, in the case of heat-\ntreatable alloys, a precipitate of maximum\nsize; for this the cooling rate must be close-\nTypical full annealing treatments for some common wrought aluminum alloys\nThese treatments, which anneal the material to the O temper, are typical for various sizes and methods of\nmanufacture and may not exactly describe optimum treatments for specific items.\nMetal temperature\nAlloy\n\u00b0C\n\u00b0F\nApproximate\ntime at\ntemperature, h\nMetal temperature\nAlloy\n\u00b0C\n\u00b0F\nApproximate\ntime at\ntemperature, h\n1060\n345\n650\n(a)\n5457\n345\n650\n(a)\n1100\n345\n650\n(a)\n5652\n345\n650\n(a)\n1350\n345\n650\n(a)\n6005\n415(b)\n775(b)\n2-3\n2014\n415(b)\n775(b)\n2-3\n6009\n415(b) 775(b)\n2-3\n2017\n415(b)\n775(b)\n2-3\n6010\n415(b) 775(b)\n2-3\n2024\n415(b)\n775(b)\n2-3\n6053\n415(b) 775(b)\n2-3\n2036\n385(b)\n725(b)\n2-3\n6061\n415(b) 775(b)\n2-3\n2117\n415(b)\n775(b)\n2-3\n6063\n415(b) 775(b)\n2-3\n2124\n415(b)\n775(b)\n2-3\n6066\n415(b) 775(b)\n2-3\n2219\n415(b)\n775(b)\n2-3\n7001\n415(c)\n775(c)\n2-3\n3003\n415\n775\n(a)\n7005\n345(d)\n650(d)\n2-3\n3004\n345\n650\n(a)\n7049\n415(c)\n775(c)\n2-3\n3105\n345\n650\n(a)\n7050\n415(c)\n775(c)\n2-3\n5005\n345\n650\n(a)\n7075\n415(c)\n775(c)\n2-3\n5050\n345\n650\n(a)\n7079\n415(c)\n775(c)\n5052\n345\n650\n(a)\n7178\n415(c) 775(c)\n2-3\n2-3\n5056\n345\n650\n(a)\n7475\n415(c)\n775(c)\n2-3\n5083\n345\n650\n(a)\n5086\n345\n650\nBrazing sheet\n5154\n345\n650\n5182\n345\n650\n(a)\n5254\n345\n650\n(a)\nNo. 11 and 12\n345\n650\n5454\n345\n650\n(a)\nNo. 21 and 22\n345\n650\n5456\n345\n650\n(a)\nNo. 23 and 24\n345\n650\n(a)\n(a)\n(a)\n(a) Time in the furnace need not be longer than necessary to bring all parts of the load to appealing temperature. Cooling rate is\nunimportant. (b) These treatments are intended to remove the effects of solution treatment and include cooling at a rate of about 30 \u00b0C/h\n(50 \u00b0F/h) from the annealing temperature to 260 \u00b0C (500 \u00b0F). Rate of subsequent cooling is unimportant. Treatment at 345 \u00b0C (650 \u00b0F),\nfollowed by uncontrolled cooling, may be used to remove the effects of cold work or to partly remove the effects of heat treatment. (c)\nThese treatments are intended to remove the effects of solution treatment and include cooling at an uncontrolled rate to 205 \u00b0C (400 \u00b0F)\nor less, followed by reheating to 230 \u00b0C (450 \u00b0F) for 4 h. Treatment at 345 \u00b0C (650 \u00b0F), followed by uncontrolled cooling, may be used\nto remove the effects of cold work or to partly remove the effects of heat treatment. (d) Cooling rate to 205 \u00b0C (400 \u00b0F) or below is less\nthan or equal to 30 \u00b0C/h (50 \u00b0F/h).\nHeat Treating of Aluminum Alloys / 871\nly controlled. Even allowing the load to cool\nin the furnace may result in an excessively\nhigh rate. Similarly, lowering the furnace-\ncontrol instrument by 28 \u00b0C (50 \u00b0F) each\nhour may produce stepped cooling, which is\nnot satisfactory for severe forming opera-\ntions. For maximum softening, a continu-\nous cooling rate of not more than 28 \u00b0C/h (50\n\u00b0F/h) is recommended.\nAnnealing of castings for 2 to 4 h at\ntemperatures from 315 to 345 \u00b0C (600 to 650\n\u00b0F) provides the most complete relief of\nresidual stresses and precipitation of the\nphases formed by the excess solute retained\nin solid solution in the as-cast condition.\nSuch annealing treatments provide maxi-\nmum dimensional stability for service at\nelevated temperatures. The annealed tem-\nper is designated \u201cO.\" (This temper was\ndesignated \"T2\" prior to 1975.)\nGrain Growth\nMany of the aluminum alloys in common\nuse are subject to grain growth during solu-\ntion treatment or annealing. This phenome-\nnon can occur during or after recrystallization\nof material that has been subjected to a small\ncritical amount of prior cold work. It is usu-\nally manifested by surface roughening during\nsubsequent fabrication operations and fre-\nquently results in rejections for appearance or\nfunctional reasons. Less frequently, some de-\nterioration of mechanical properties is en-\ncountered, and this is undesirable regardless\nof surface-roughening effects.\nDegree of susceptibility to grain growth\nvaries with alloy, structure, and chemical-\ncomposition variation, and from one prod-\nuct form to another. The critical range of\ncold work is ordinarily about 5 to 15%.\nUsually, temperatures of 400 \u00b0C (750 \u00b0F)\nand above must be reached before grain\ngrowth occurs, but some growth has been\nencountered at temperatures as low as 345\n\u00b0C (650 \u00b0F). Grain growth that occurs during\ninitial recrystallization is more a function of\ncomposition, structure, and degree of cold\nwork than of temperature per se; tempera-\ntures in excess of 455 \u00b0C (850 \u00b0F) in common\nalloys can lead to secondary-recrystalliza-\ntion grain-growth problems. The common\nsymptom indicating moderately large-grain\nmaterial is roughening or \"orange peel\" on\nthe external surfaces of bends. Severe\ngrowth of grains to fingernail size and larger\nsometimes is evident in parts made from\nannealed (O temper) material by stretch\nforming and then thermal treating or similar\noperations. This type of grain growth often\nis detected during subsequent anodizing,\netching, and chemical milling operations.\nCracking during welding or brazing is\nanother characteristic which may indicate\nthat severe grain growth has occurred. In\nsuch instances, cracks propagate along\ngrain boundaries that provide little obstruc-\ntion to their progress.\n872 / Heat Treating of Nonferrous Alloys\n200\n\u201c100 B\n150\n175 \u00b0C (350 \u00b0F)-\nYield strength, MPa\nYield strength, MPa\n100\n50\n350\n300\n250\n205 \u00b0C (400 \u00b0F)\n230 \u00b0C (450 \u00b0F)\n260 \u00b0C (500 \u00b0F)\n288 and 316 \u00b0C (550 and 600 \u00b0F)\n1\n1\n0\n0.5\n1\n1.5\n2\n2.5\nTime, h\n3\n3.5\n4\n4.5\n5\n20\n10\n0\n\u25ac\u25ac\u25ac\u25ac\u25ac 5052-+18 7 50\n175 \u00b0C (350 \u00b0F)\n230 \u00b0C (450 \u00b0F)\n200\n260 \u00b0C (500 \u00b0F)\n150\n100\n50\n290 \u00b0C (550 \u00b0F)\n\u2610 315 \u00b0C (600 \u00b0F)\n205 \u00b0C (400 \u00b0F)\n0\n0\n0.5\n1\n1.5\n2\n2.5\n3\n3.5\n4\n4.5\nTime, h\nFig 32 Representative isothermal annealing curves for alloys 1100-H18 and 5052-H18\nIf the surface roughening is objectionable\nfrom either an appearance or a functional\naspect, the desirability of surface-smooth-\ning operations, such as sanding or buffing,\nmust be evaluated. If reductions in mechan-\nical properties are suspected, these must be\nestablished by test and evaluated in relation\nto the anticipated service.\nIn one application, a part that had been\nmade by stretch forming O-temper 2 mm\n(0.080 in.) sheet and heat treating exhibited\nsignificantly lower tensile and yield\nstrengths in portions where severe grain\ngrowth had occurred than in portions hav-\ning normal grain size:\nYield strength\nksi\nTensile strength\nTest\nGrain\nstructure\nMPa\nksi\nMPa\nTransverse\n1\nCoarse\n265\n38.5\n247\n35.8\n2\nCoarse\n263\n38.2\n241\n35.0\n3\nFine\n311\n45.1\n261\n37.8\nLongitudinal\n1\nCoarse\n259\n37.6\n243\n35.3\n2\nCoarse\n269\n39.0\n245\n35.6\n3\nFine\n305\n44.2\n270\n39.1\n5\n40\n40\n30\n20\n10\nYield strength, ksi\nYield strength, ksi\nIn other similar investigations, no detri-\nmental effects have been discovered, and in\nmany cases such parts have served satisfac-\ntorily in critical applications.\nWhen a grain-growth problem is discov-\nered, it is too late to change the condition of\nthe parts in question, but several possible\nmethods are available for preventing recur-\nrence of the difficulty. The simplest of these\nis relieving the causative stress by interject\ning a stress-relief anneal into the manufac-\nturing sequence immediately prior to the\nsolution-treating or full-annealing cycle in\nwhich the grain growth occurred. This ap-\nproach is usually successful and practical.\nAnother possibility is to adjust the amount\nof stress present in the part immediately\nprior to the critical heat treatment so that\nthe stress level is outside the critical range.\nThis may be done by adding a cold-working\noperation before forming, such as pre-\nstretching of blanks, or by forming in mul-\ntiple stages with a stress-relief anneal before\neach stage.\nA third method that is sometimes suc-\ncessful consists of increasing the heating\nrate during the critical heat treatment by\nreducing the size of furnace loads or by\nchanging from an air furnace to a salt bath.\npaired\nIn one application, severe grain growth was\nfound during bending of alloy 1100 rectan-\ngular tubing. The roughening of the inside\nsurfaces of the parts, which occurred during\nforming of the large-grain material, im-\ntheir functioning as radar\nwaveguides. Investigation disclosed that, to\nminimize handling marks, the material was\nprocured in the strain-hardened (H14) tem-\nper and was stress-relief annealed at 345 \u00b0C\n(650 \u00b0F) immediately prior to forming. Grain\ngrowth occurred during annealing as a re-\nsult of the moderate amount of cold work\nintroduced at the mill. The problem was\neliminated by changing the stress-relieving\noperation to a 5-min heating period in an air\nfurnace operating at 540 \u00b0C (1000 \u00b0F). The\nexplanation advanced for the success of this\ntreatment was that, due to the rapid heating\nrate, the temperature of the material was\nraised through the recrystallization range\nfor the less severely cold-worked grains\nbefore the critically cold-worked grains had\ntime to grow appreciably.\nHeating Equipment and\nAccessories\nThe general methods for heat treating\naluminum alloys include the use of molten\nsalt baths, air-chamber furnaces, and induc-\ntion heaters. The choice of heating equip-\nment depends largely on the alloy and the\nconfiguration of the parts to be processed.\nThe type of heat treatment can also influ-\nence the choice of heating equipment. For\nexample, both molten salt baths and air-\nchamber furnaces are suitable for solution\ntreating of aluminum alloys, while induction\nheating requires additional analysis to de-\nfine the proper temperature range for solu-\ntion treatment. Induction methods can pro-\nvide high heating rates, which affect\ntransformation behavior (see, for example,\nthe section \"Nonequilibrium Melting\" in\nthis article).\nMolten salt baths and air-chamber furnac-\nes both have advantages and disadvantages\nin solution heat treatments, as discussed\nbelow. Oil and gas-fired furnaces, in de-\nsigns that allow the products of combustion\nto come in contact with the work, are\nusually unsatisfactory because they pro-\nmote high-temperature oxidation.\nSalt baths heat the work faster (see Table\n2) than air furnaces, provided that the\namount of work introduced at any one time\nis controlled to prevent the temperature\nfrom falling below the desired range. If the\ntemperature is permitted to fall below the\nminimum limit, much of the advantage of\nthe salt bath is lost, because of the necessity\nfor reheating the large mass of salt.\nSalt baths are also more readily adapted\nto the introduction, at any time, of small\namounts of work requiring different soaking\nTable 12 Frequency selection for induction heating with a longitudinal-flux coil and a\ntransverse-flux coil\nMinimum part thickness, mm (in.), for a frequency of:\n200 Hz\n1 kHz\n3 kHz\n10 kHz\nHeat Treating of Aluminum Alloys / 873\nProduction rate, tons/h\n0\n1\n2\n1800\n3\n4\n5\n6\n7\nSolution heat treat at 590 \u00b0C\n1500\nFull anneal at 425 \u00b0C\nPartial anneal at 315 \u00b0C\n1200\nMaterial\n60 Hz\nSolenoid (longitudinal-flux) coil\nSteel below Curie temperature\n>38 (1.5)\n13 (0.5)\nSteel above Curie temperature\nBrass\n>175 (7.0)\n100 (4.0)\n5 (0.2)\n43 (1.7)\n2.3 (0.09)\n25 (1.0)\n>50 (2.0)\n28 (1.1)\n13 (0.5)\nAluminum\n>38 (1.5)\n22 (0.85)\n9 (0.375)\n7 (0.28)\n5 (0.2)\n1 (0.04)\n13 (0.5)\n4 (0.16)\n3 (0.12)\nTransverse-flux coil\nAluminum\n>5 (0.2)\n1.3 (0.05)\nBrass\nSteel\n>10 (0.4)\n>50 (2.0)\n2.5 (0.1)\n13 (0.5)\n0.25 (0.01)\n0.5 (0.02)\n2.5 (0.1)\n0.08 (0.003)\n0.15 (0.006)\n0.9 (0.035)\n0.04 (0.0016)\n0.08 (0.0032)\n0.5 (0.020)\nPower source, kW\n900\n600\n300\n0\n0\n5\n10\n15 20\nLine speed, m/min\n25\n30\nperiods. (Economical utilization of air fur-\nnaces usually dictates accumulation of a\nlarge load of parts of similar thickness be-\nfore charging.) Also, the buoyant effect of\nthe salt reduces distortion during heating,\nand the large reservoir of heat facilitates\ntemperature control and uniformity.\nSalt bath operation entails special house-\nkeeping requirements. Dragout is costly and\nunsightly. Because residual salt on parts\nmay result in corrosion, all salt must be\ncompletely removed, including that from\ncrevices and blind holes. In addition, salt\nresidue from the quench water must be kept\nto a minimum by a constant water overflow\nor by providing a fresh-water rinse for all\nparts after quenching. When these provi-\nsions are impractical, corrosion can be in-\nhibited by adding 14 g (\u00bd oz) of sodium or\npotassium dichromate to each 45 kg (100 lb)\nof the molten salt.\nPrecautions. Molten salt baths are poten-\ntially hazardous and require special precau-\ntions. Operators must be protected from\nsplashing and dripping of the hot salt. Be-\ncause heated nitrates are powerful oxidizing\nagents, they must never by allowed to come\nin contact with combustibles and reducing\nagents, such as magnesium and cyanides.\nMost authorities advise against inserting\naluminum alloys containing more than a few\npercent of magnesium into molten nitrate.\nTo avoid exposure of personnel to nitrous\nfumes produced during decomposition of\nnitrates, good ventilation is essential.\nWhen molten nitrates are being used, the\npossibilities of explosions resulting from\nboth physical and chemical reactions must\nbe avoided. The former result from rapid\nexpansion of gases entrapped beneath the\nsurface of the bath. Hence, parts entering\nthe bath must be clean and dry; they must\nalso be free of pockets or cavities that\ncontain air or other gases. Chemical-reac-\ntion explosions result from rapid break-\ndown of the nitrates due to overheating or\nreaction with the pot material. Stainless\nsteel pots (preferably of type 321 or 347) are\nmore resistant to scaling than those made of\ncarbon steel or cast iron and therefore pre-\nsent a lower probability of local overheat-\ning. Sludge or sediment accumulations in\nbottom-heated pots can also lead to local\noverheating. Overheat controls are essen-\ntial to ensure against temperatures exceed-\ning 595 \u00b0C (1100 \u00b0F).\nIt is vitally important that water be kept\naway from a nitrate tank. In controlling a\nnitrate fire, do not use water or any fire\nextinguisher containing water. The best ex-\ntinguisher is dry sand, a supply of which\nshould be kept near the tank.\nExtra sacks of salt should be stored in a\ndry place, distant from the tank. If the fresh\nsalt being added to the bath is even slightly\ndamp, it should be added very slowly or\nwhen the bath is frozen.\nAir furnaces are used more widely than salt\nbaths because they permit greater flexibility\nin operating temperature. When production\nschedules and the variety of alloys requiring\nheat treatment necessitate frequent changes\nin temperature, the time and cost of adjusting\nthe temperature of a large mass of salt makes\nthe use of an air furnace almost mandatory.\nHowever, waiting periods are often required\nto allow the walls of air furnaces to stabilize at\nthe new temperature before parts are intro-\nduced. Otherwise, parts may radiate heat to\ncolder walls or absorb radiant heat from hot-\nter walls, and the temperature indicated by\nthe control instrument will not reflect actual\nmetal temperature in the usual manner. Air\nfurnaces are also more economical when the\nproduct mix includes a few rather large parts;\nholding the temperature of a large volume of\nsalt in readiness for an occasional large part is\nfar more expensive than heating an equal\nvolume of air.\nInduction heating with either solenoid\n(longitudinal-flux) coils or transverse-flux\ncoils provides an efficient method for in-line\nheating of flat-rolled products (particularly\nstrip). Solenoid coils create a longitudinal\nflux, which can produce efficient heating for\nthicker and/or lower resistivity materials.\nSolenoid coils can also be used efficiently in\nthe heating of thinner magnetic material\n(see, for example, steel below the Curie\ntemperature in Table 12).\nFor several nonferrous materials (alumi-\nnum, copper, brass), however, efficiency\nand power factors with solenoid coils are\nsignificantly lower than for ferrous materi-\nPower requirement for transverse-flux induc-\nFig 33\ntion heating of aluminum strip 1 mm (0.04 in.)\nthick and 1270 mm (50 in.) wide. Source: Ref 9\nals. Therefore, transverse-flux coils are ide-\nally suited for heating nonferrous materials,\nbecause transverse-flux lines do not exhibit\nthe degree of current cancellation associat-\ned with longitudinal flux lines. This aspect\nof transverse-flux coils improves efficiency\nand also permits the use of lower frequen-\ncies (Table 12). This reduces the capital\nequipment costs, and where it shifts from\nrequiring RF frequencies, the power source\nconversion efficiency is also significantly\nimproved. Aluminum, brass, copper, and\naustenitic stainless steel strip lines are ide-\nally suited for transverse-flux heating. Each\nof these materials often requires in-line pro-\ncesses like partial or full annealing or solu-\ntion heat treating to provide necessary me-\nchanical properties for subsequent finishing\noperations.\nTransverse-flux induction heating offers\nseveral benefits for in-line strip heating and\nhas been known for many years. However,\nit requires specially designed iron-cored\nlaminated inductor coils and tighter control\nof the power, strip handling, and process\nparameters. Frequency selection is dictated\nby the resistivity and thickness of the ma-\nterial, while power requirements depend on\nthe production rates, the specific heat, and\nthe processing temperatures for a given\nmaterial. Figure 33 shows typical power\nsource requirements for transverse-flux\nheating of aluminum strip.\nFurnace Temperature Control\nThe importance of close temperature con-\ntrol in solution treating has been noted in\nthe previous section on solution treating.\nEach control zone of each furnace should\ncontain at least two thermocouples. One\nthermocouple, with its instrument, should\nact as a controller, regulating the heat input;\nthe other should act independently as a\nsafety cutoff, requiring manual reset if its\nset temperature (usually the maximum of\n874 / Heat Treating of Nonferrous Alloys\nthe specified range) is exceeded during the\nsolution-treating cycle.\nSafety cutoffs are mandatory for salt\nbaths to guard against explosions and often\nhave paid for themselves in air furnaces by\nsaving a load of parts or even the furnace\nitself. It is important, however, that they be\ntested periodically (by deliberately over-\nshooting the empty furnace) to guard\nagainst \"frozen\" corroded contacts result-\ning from prolonged periods of idleness.\nAt least one of the instruments for each\nzone should be of the recording type, and\nboth instruments should have restricted\nscales for instance, 400 to 600 \u00b0C (750 to\n1110 \u00b0F), rather than 0 to 600 \u00b0C (32 to 1110\n\u00b0F). This is required for maximum accuracy\nbecause manufacturers' guarantees are\nspecified in terms of percent of scale.\nIn the placement of instruments, expo-\nsure to extremes in ambient temperature,\nhumidity, vibration, dust, and corrosive\nfumes should be avoided. Ambient temper-\natures between 5 and 50 \u00b0C (40 and 120 \u00b0F)\nare satisfactory, but temperature changes of\n6 \u00b0C/h (10 \u00b0F/h) or more should be avoided.\nIt is also essential that instruments and\nthermocouple circuits be shielded from\nelectromagnetic fields commonly associat-\ned with the leads of high-amperage furnace\nheating elements.\nTemperature-sensing elements must be\ncapable of responding more rapidly to tem-\nperature changes than the materials being\nprocessed. Therefore, thermocouple wire\ndiameter should not exceed 1\u00bd times the\nthickness of the minimum-gage material to\nbe heat treated, and should in no case\nexceed 14 gage. Thermocouples for salt\nbaths should be enclosed in suitable protec-\ntion tubes. Air-furnace thermocouples\nshould be installed in open-end protection\ntubes, with the thermocouple junction ex-\ntending sufficiently beyond the tube to pre-\nvent any loss in sensitivity.\nTemperature-sensing elements should be\nlocated in the furnace work chamber, not in\nducts and plenums, and should be as close\nas possible to the working zone. Specifica-\ntion MIL-H-6088C restricts distance be-\ntween the sensing element and the working\nzone to a maximum of 100 mm (4 in.). The\nsafety-cutoff thermocouple should be locat-\ned to reflect the highest temperature in the\nworking zone. The control thermocouple\nshould be located in a position where it will\nread a temperature approximately halfway\nbetween the hottest and coldest tempera-\ntures.\nProbe Checks. After the temperature-\nmeasurement equipment is properly in-\nstalled, it must be checked frequently for\naccuracy. This is accomplished by inserting\na calibrated probe thermocouple into the\nfurnace adjacent to each furnace thermo-\ncouple and comparing its reading on a cali-\nbrated test potentiometer with that indicat-\ned by the furnace instrument. Correction\nfactors should be applied after each probe\ncheck, but if the correction required ex-\nceeds \u00b13 \u00b0C (\u00b15 \u00b0F), the source of the\ndeviation should be corrected. MIL-H-6088\nrecommends that this check be made week-\nly, but many operators make the check as\nfrequently as once each shift.\nTemperature-Uniformity Surveys. In con-\ntrolling the temperature of parts that are\nbeing heat treated it must first be deter-\nmined that the temperature indicated by the\nfurnace instruments truly represents the\ntemperature of the nearby air or salt. Sec-\nond, the uniformity of temperature within\nthe working zone must be shown to be\nwithin a range of 11 \u00b0C, or 20 \u00b0F (6 \u00b0C, or 10\n\u00b0F, for precipitation heat treatment of alloy\n2024). This is accomplished by measuring\nthe temperature at several test locations,\nusing calibrated test thermocouples and a\ncalibrated test potentiometer, and reading\nfurnace instruments nearly simultaneously.\nMIL-H-6088 recommends monthly surveys\nwith one test location per 1.1 m\u00b3, or 40 ft\u00b3\n(0.7 m\u00b3, or 25 ft\u00b3, for air furnaces on initial\nsurvey), but with a minimum of nine test\nlocations distributed as shown in Fig 34.\nDespite the large size of some furnaces,\nrather surprising temperature uniformities\nhave been reported. In one instance the\ninitial survey of an air furnace measuring\n12.5 by 1.2 by 3.0 m (41 by 4 by 10 ft)\nshowed maximum temperature variations of\n+1.7, \u22121.1 \u00b0C (+3, \u22122 \u00b0F). When a parti-\ntion 0.3 m (1 ft) thick was lowered, convert-\ning the furnace to two chambers 6.1 by 1.2\nby 3.0 m (20 by 4 by 10 ft) each, the spread\nwas 1.1, 0.6 \u00b0C (+2, -1 \u00b0F) in one\nsection and +0.6, \u22121.1 \u00b0C (+1, \u22122 \u00b0F) in\nthe other.\nFor each furnace load, one thermocouple\n(the \"cold\" couple) should be placed in the\ncoldest area of the furnace and another (the\n\"hot\" couple) in the hottest area. In addi-\ntion to these two thermocouples, a load\nthermocouple should be installed. The load\ncouple should be of approximately the same\ngage as the sheet or other product being\nheat treated. If heavy plate, forgings, or\ncastings are being heat treated, a similar\ndiscarded item should be used at the con-\ntrolling load couple. The thermocouple\nshould be placed in a drilled hole and\npacked to hold it firmly in place during the\nheat-treating cycle. In some instances, the\nitems being heat treated can be used as the\nload couples. The thermocouples can be\nplaced in holes drilled in areas that will be\nremoved in making the finished article.\nIt is important that items of different\nthicknesses-1 mm (0.040 in.) sheet and 25\nmm (1 in.) plate, for example-not be heat\ntreated in the same furnace load.\nIn salt baths, uniformity surveys usually\nare made by holding a probe thermocouple\nin each location until thermal equilibrium is\nreached; in air furnaces, a mock heat-treat-\ning cycle is required. First, the air furnace is\n6\n5\n5-194\n8\n=\nRectangular furnace\n\u798f\nCylindrical salt bath\nWINE\n6 5 4\n8\nCylindrical air furnace\nLocation of thermocouples for surveying\nFig 34\ntemperature uniformity in the working\nzones of air furnaces and salt baths\nstabilized at the test temperature. Then a\nrack containing the test thermocouples is\ninserted into the furnace. By using multiple\nswitches or a multipoint recording instru-\nment, all test thermocouples and furnace\ninstruments are read every 5 min. As the\ntemperature approaches the test range, it is\nadvisable to increase the frequency of read-\nings to detect possible overshooting. After\nthermal equilibrium is reached, readings\nshould be continued until the recurrent tem-\nperature pattern is established.\nSurveys of salt baths generally are con-\nsidered acceptable whether they are made\nwhile the bath is empty or filled with work.\nIt is controversial whether surveys of air\nfurnaces should be made with or without a\nload. Undoubtedly, recovery overshoots\nare most likely to occur with a very light\nload and would not be detected if a heavier\nload were used. Certainly, if all loads are\nessentially alike, surveys should be made\nwith typical loads. With widely varying\nloads, the optimum approach is to make\nseveral surveys initially, including one with\nan empty furnace, and then to make suc-\nceeding surveys with an empty furnace to\nensure against changes in furnace charac-\nteristics. If any changes are made in the\nfurnace that might affect temperature distri-\nbution, such as repair of vanes or louvers,\nseveral surveys should be repeated.\nAnother aspect of the problem of temper-\nature control in air furnaces is the necessity\nof ensuring that the temperature of the parts\nis the same as that of the surrounding air.\nFurnace components whose temperature\ndiffers from the air temperature must be\nsuitably shielded to prevent radiation to or\nfrom the parts being heat treated. In a\nfurnace used for solution heat treating of\nrivets, unshielded heating elements have\nbeen known to produce part temperatures\nas much as 20 \u00b0C (35 \u00b0F) higher than the\ncontrol temperature, resulting in eutectic\nmelting and cracking. In two other instanc-\nes, reradiation through inadequate shielding\nproduced a radiation effect of as much as 11\n\u00b0C (20 \u00b0F). One of these problems was\nsolved by painting the shield with reflective\naluminum paint and the other by adding a 13\nmm (1/2 in.) thick layer of asbestos to the 1.6\nmm (1/16 in.) stainless steel shield.\nFurnace-wall temperatures that differ ap-\npreciably from the temperature of the parts\nalso must be avoided. Consequently, when\nthe operating temperature of an air furnace\nis changed, waiting periods are required\nafter the furnace instrument indicates sta-\nbility, to allow the furnace walls to stabilize\nat the new temperature. The magnitude of\nthis limitation is directly proportional to the\nefficiency of the furnace as an insulated\nchamber, but possibilities of such radiation\nshould be recognized even in thin-wall fur-\nnaces.\nRadiation effects are potentially danger-\nous because they often cannot be detected\nby ordinary thermocouples. Specially pre-\npared radiation panels with thermocouples\nattached are used, and their readings are\ncompared with adjacent free thermocou-\nples. These panels normally are made of\nmaterial of the same gage as the thinnest\nparts to be heat treated and should have a\nsingle surface area of about 650 cm\u00b2 (100\nin.2). A thermocouple is attached to the\ncenter of the panel by welding or peening.\nIn order to detect the maximum effect,\npanel surfaces should be darkened so that\ntheir emissivity is at least as high as that of\nany material to be processed. During the\ntest, the panel surfaces should be parallel to\nthe suspected source or recipient of radia-\ntion. As an example of the number of panels\nrequired, several aerospace companies\nspecify one panel for every 1.5 linear meters\n(5 linear feet) of furnace wall.\nInstrument Calibration. All instruments\nand thermocouples must be accurately cal-\nibrated, and it is essential that the calibra-\ntions be traceable directly to the National\nBureau of Standards. The chain of trace-\nability should consist of not more than four\nlinks for sensing elements and three links\nfor measuring elements. To illustrate, if the\narticle calibrated by the National Bureau of\nStandards is called a primary standard, then\nthe chain of traceability of measuring ele-\nments should consist of primary standard,\ntest potentiometer, and furnace instrument.\nSimilarly, the chain for sensing elements\nshould consist of primary standard, second-\nary standard, test thermocouple, and fur-\nnace thermocouple. Every effort should be\nmade to ensure that the temperature indi-\ncated by the furnace instruments is as close\nas possible to the actual temperature. To\nachieve this, it is necessary to apply correc-\ntion factors obtained during calibration to\nthe next lower echelon of accuracy. Even\nthen, if all errors inherent in the chain are in\nthe same direction, a considerable differ-\nence will exist between the measured and\nactual temperatures. Therefore, it is advis-\nable to operate as close to the mean of the\ndesired range as possible.\nDimensional Changes during\nHeat Treatment\nIn addition to the completely reversible\nchanges in dimensions that are simple func-\ntions of temperature change and are caused\nby thermal expansion and contraction, di-\nmensional changes of more permanent char-\nacter are encountered during heat treat-\nment. These changes are of several types,\nsome of mechanical origin and others\ncaused by changes in metallurgical struc-\nture. Changes of mechanical origin include\nthose arising from stresses developed by\ngravitational or other applied forces, from\nthermally induced stresses or from relax-\nation of residual stresses. Dimensional\nchanges also accompany recrystallization,\nsolution, and precipitation of alloying ele-\nments.\nSolution Heat Treatment. Distortion as a\nresult of creep during solution heat treat-\nment should be avoided by proper loading\nof parts in baskets, racks, or fixtures, or by\nprovision of adequate support for long piec-\nes of plate, rod, bar, and extrusions heat\ntreated in horizontal roller hearth furnaces.\nSheet is provided with air-pressure support\nin continuous heat-treating furnaces to\navoid scratching, gouging, and distortion. If\nparts are to be solution heat treated in\nfixtures or racks made of materials (such as\nsteel) with coefficients of thermal expansion\nlower than that of the aluminum being treat-\ned, allowance should be made for this dif-\nferential expansion to ensure that expansion\nof the aluminum is not restricted. Straight-\nening immediately after solution heat treat-\ning may be preferable to fixturing.\nSolution of phases formed by major alloy-\ning elements causes volumetric expansion\nor contraction, depending on the alloy sys-\ntem, and this may have to be taken into\naccount in heat treatment of long pieces.\nFor example, solution heat treatment and\nquenching of alloy 2219 causes lengthwise\ncontraction of about 2 mm/m (0.002 in./in.).\nSolution heat treatment and quenching of\nalloys of the 7xxx series is accompanied by\nlengthwise expansion-about 0.6 mm/m\n(0.0006 in./in.) for alloy 7075 rod or plate.\nQuenching. The most troublesome\nchanges in dimensions and shape are those\nthat occur during quenching or that result\nfrom stresses induced by quenching. Due to\nits nonuniform cooling, quenching may pro-\nduce warpage or distortion, particularly in\nHeat Treating of Aluminum Alloys / 875\nLongitudinal stress range, MPa\n500\n400\n300\nAverage cooling rate at center\nof cylinders, \u00b0F/s\n200\n\u00b0\n100\n0\nPoo\n10\n10\n100\n0\n100\nAverage cooling rate at center\nof cylinders, \u00b0C/s\n1660\n40\n20\n20\nLongitudinal stress range, ksi\nEffect of quenching rate on longitudinal\nFig 35\nstress ranges in alloy 2014-T4 cylinders\nquenched in various media. Cylinders were 75 mm (3\nin.) in diameter by 230 mm (9 in.) long. Cooling rate\nwas measured from 400 to 290 \u00b0C (750 to 555 \u00b0F). Stress\nrange is maximum tensile stress plus maximum com-\npressive stress.\nthin material and in thin sections of parts\nthat contain variations in thickness. For\nthick-section products or parts, changes in\nexternal shape may be small because of\nrigidity, but the interior-to-surface temper-\nature gradients that form with rapid cooling\ncreate residual stresses; these stresses nor-\nmally are compressive at the surfaces and\ntensile in the interior.\nAs previously discussed, warpage or dis-\ntortion of thin-section material can be re-\nduced by using a quenching medium that\nprovides slower cooling; however, cooling\nmust be sufficient to produce the required\nproperties. Slower quenching can also re-\nduce the magnitude of residual stresses in\nthicker parts or pieces, as shown in Fig 9 for\ncylindrical specimens of alloy 6151\nquenched in cold or boiling water. Stress\nrange (maximum tensile stress plus maxi-\nmum compressive stress) for a cylinder with\na radius of 89 mm (3.5 in.) is about 205 MPa\n(30 ksi) when the cylinder is quenched in\ncold water but less than 70 MPa (10 ksi)\nwhen it is quenched in boiling water. The\neffects of average cooling rate through the\ntemperature range from 400 to 290 \u00b0C (750\nto 550 \u00b0F) on longitudinal stress ranges\ndeveloped in alloy 2014 cylinders 75 mm (3\nin.) in diameter are shown in Fig 35.\nHigh stresses induced by rapid quenching\ngenerally are reduced only modestly by the\nprecipitation heat treatments used to pro-\nduce T6- or T8-type tempers. Consequent-\nly, for the alloys that require rapid cooling\nto develop the properties of these tempers,\nthose incorporating mechanical stress relief\n(Tx51, Tx52) usually are specified when\nsubstantial metal must be removed to pro-\nduce final shapes. Other T8-type tempers,\nsuch as T86 and T87, also have low residual\nstress as a result of the stretching required\nto produce them.\nHeat Treatments for Precipitation and Sta-\nbilization. The most significant dimensional\nchanges associated with precipitation heat\n876 / Heat Treating of Nonferrous Alloys\ntreatments and stabilizing heat treatments\narise from concurrent dilution of the solid\nsolution (which changes lattice parameter)\nand formation of precipitate. Changes in\ndensity and specific volume resulting from\nthese changes in metallurgical structure are\nthe reverse of those caused by solution of\nthe alloy phases. However, because the\nstrongest tempers are those in which the\nprecipitate is present in nonequilibrium\ntransition forms, the amount of change dur-\ning precipitation heat treatment does not\ntotally compensate for the previous (and\nopposite) change that occurred during solu-\ntion heat treatment. Most of the heat-treat-\nable alloys expand (grow) during precipita-\ntion heat treatment. Exceptions are alloys\nof the 7xxx wrought series and the 7xx.0\ncasting series, which exhibit contraction.\nIn alloys of the 2xxx series, the amount of\ngrowth decreases with increasing magne-\nsium content. Thus, growth of about 1.5\nmm/m (0.0015 in./in.) can be expected dur-\ning precipitation heat treatment of alloy\n2219-T87, about 0.5 mm/m (0.005 in./in.) for\ntreatment of alloy 2014-T6 and less than 0.1\nmm/m (0.0001 in./in.) for treatment of alloy\n2024-T851. Alloys 7050 and 7075, on the\nother hand, contract about 0.3 mm/m\n(0.0003 in./in.) on precipitation heat treating\nfrom the W temper to the T6 temper and\nabout 0.7 mm/m (0.0007 in./in.) on treating\nfrom the W temper to the T73 temper.\nStabilizing T7-type treatments cause great-\ner amounts of growth than the T5-, T6-, or\nT8-type treatments for the same alloys. This\nincreased growth is associated either with\nformation of increased amounts of transition\nprecipitates or with transformation of transi-\ntion precipitates to equilibrium phases.\nDimensional Stability in Service\nDimensional stability of heat-treated\nparts in service depends on alloy, temper,\nand service conditions. Of the latter, ex-\ncluding mechanical conditions such as ap-\nplied loads, the most important is service\ntemperature range relative to the range in\nwhich precipitation occurs. Residual stress-\nes constitute another source of dimensional\nchanges. Stress relief minimizes changes\ndue to residual stresses, and most mill prod-\nucts usually are supplied in tempers that\ninclude stress relief. Potential dimensional\nchange as a result of further precipitation in\nparts that operate at elevated temperatures\nis minimized for wrought products by use of\nT7-type stabilizing treatments and for cast-\nings by use of T5-type treatments. Howev-\ner, components of high-precision equip-\nment, such as instruments for aerospace\nguidance systems and optical and telescopic\ndevices, may require special supplementary\ntreatments during manufacture to further\nreduce stresses or subsequent precipitation.\n(These treatments are discussed below, un-\nder \"Stability of Precision Equipment.\")\nThe T3- and T4-type tempers are the least\nstable dimensionally because of possible\nprecipitation in service. Alloys 2024 and its\nvariants have the smallest dimensional\nchange in aging; the total change from the\nquenched to the average state is of the order\nof 0.06 mm/m (0.00006 in./in.), less than the\nchange due to a temperature variation of 3\n\u00b0C (5 \u00b0F). These alloys therefore can be used\nin the T3- and T4-type tempers, except for\nprecision equipment. For all other alloys,\nT6- or T8-type tempers should be used,\nbecause in these tempers all the alloys have\ngood dimensional stability.\n% of tests\nYield strength, ksi\n60\n65\n50\nT\n1\n40\n180 specimens\nfrom a single sheet\n30\nde 20\n10\n70\nT\n7075-T6\nAlclad sheet\n4290 routine\nmill tests\n75\n400\n425\n450\nYield strength, MPa\n475\n500\n525\nin heat-treated 7075-T6 clad sheet product\nwith distribution in a single sheet. A is 95% probability\nthat not more than 1% of all material will fall below this\nvalue; B is 95% probability that not more than 10% of\nall material will fall below this value. (A and B refer\nonly to curve representing 4290 routine mill tests.)\nStability of Precision Equipment. Proper\nmaintenance of high-precision devices,\nsuch as gyros, accelerometers, and optical\nsystems, requires use of materials in which\ndimensional changes from metallurgical in-\nstability are limited from 10 \u00b5m/m (10 \u03bcin./ Fig 36 Comparison of distribution of yield strength\nin.). Several laboratory investigations and\nconsiderable practical experience have\nshown that wrought alloys 2024 and 6061\nand casting alloy 356.0 are well suited and\ngenerally preferred for such applications.\nDimensional changes were no greater than\n10 \u03bcm/m when alloys 2024-T851 and -T62,\n6061-T651 and -T62, and 356.0-T51, -T6,\nand -T7 were tested for more than a year at\nroom temperature and for several months at\n70 \u00b0C (160 \u00b0F), and then the same alloys\nwere tested with repeated thermal cycling\nbetween 20 and -70 \u00b0C (68 and -94 \u00b0F).\nBecause stresses applied or induced by\nacceleration in such devices generally are\nnot high, strength levels lower than those of\nthe highest-strength tempers frequently are\nsatisfactory. To increase precision of ma-\nchining to intended dimensions, as well as\nto promote maximum stability, it is com-\nmon practice to apply additional thermal\ntreatments for stress relief and precipitation\nof 1 to 2 h at temperatures of 175 to 205 \u00b0C\n(350 to 400 \u00b0F) after rough machining. These\nadditional treatments sometimes are repeat-\ned at successive stages of processing, and\neven after final machining. In addition, it\nhas been claimed that one or two cyclic\ntreatments consisting of cooling to -100 \u00b0C\n(-150 \u00b0F), holding for 2 h, heating to 232 to\n240 \u00b0C (450 to 465 \u00b0F) and again holding for\n2 h can improve dimensional stability of\n356-T6 castings.\nQuality Assurance\nQuality-assurance criteria that heat-treat-\ned materials must meet always include min-\nimum tensile properties and, for certain\nalloys and tempers, adequate fracture\ntoughness and resistance to detrimental\nforms of corrosion (such as intergranular or\nexfoliation attack) or to stress-corrosion\ncracking. All processing steps through heat\ntreatment must be carefully controlled to\nensure high and reliable performance.\nTensile Tests. In general, the relatively\nconstant relationships among various prop-\nerties allow the use of tensile properties\nalone as acceptance criteria. The minimum\nguaranteed strength is ordinarily that value\nabove which it has been statistically pre-\ndicted with 95% probability that 99% or\nmore of the material will pass. The inherent\nvariability within lots and among specimens\nfrom a given piece is shown in Fig 36.\nTesting provides a check for evidence of\nconformance; process capability and pro-\ncess control are the foundations for guaran-\nteed values.\nPublished minimum guaranteed values\nare applicable only to specimens cut from a\nspecific location in the product, with their\naxes oriented at a specific angle to the\ndirection of working as defined in the appli-\ncable procurement specification. In thick\nplate, for example, the guaranteed values\napply to specimens taken from a plane\nmidway between the center and the surface,\nand their axes parallel to the width dimen-\nsion (long transverse). Different properties\nshould be expected in specimens taken from\nother locations, or in specimens whose axes\nwere parallel to thickness dimension (short\ntransverse). However, the specified \"refer-\nee\" locations and orientations do provide a\nuseful basis for lot-to-lot comparisons, and\nconstitute a valuable adjunct to other pro-\ncess-control measures.\nTensile tests can be used to evaluate the\neffects of changes in the process, provided\nspecimens are carefully selected. A varia-\ntion in process that produces above-mini-\nmum properties on test specimens, howev-\ner, is not necessarily satisfactory. Its\nacceptability can be judged only by compar-\ning the resulting properties with those de-\nveloped by the standard process on similar-\nly located specimens. Finally, variations in\nheat-treating procedure are likely to affect\nHeat Treating of Aluminum Alloys / 877\n600\nTable 13 Typical acceptable hardness values for wrought aluminum alloys\nAcceptable hardness does not guarantee acceptable properties; acceptance should be based on acceptable\nhardness plus written evidence of compliance with specified heat-treating procedures. Hardness values\nhigher than the listed maximums are acceptable provided that the material is positively identified as the\ncorrect alloy.\n550\n85\n80\n80\n75\nT\nT\n70\n65\n60\n55\n50\n45\n80\n90\n100\n70\nHardness, HRB\nTensile strength versus hardness for various\naluminum alloys and tempers\n\u2022 Use a specimen that has at least 19 cm\u00b2 (3\nin.2) of surface area\nHardness\nAlloy and temper\nProduct form(a)\nHRB\nHRE\nHRH\nHRIST\n2014-T3, T4, -T42\n2014-T6, -T62, -T65\nAll\n65-70\n87-95\nSheet(b)\n80-90\n103-110\nAll others\n81-90\n104-110\n2014-T61\n2024-T3\nAll\n100-109\nNot clad(c)\n69-83\n97-106\n111-118\n82.5-87.5\nTensile strength, MPa\n500\n450\n400\nClad, \u22641.60 mm (0.063\nin.)\n52-71\n91-100\n109-116\n80-84.5\nClad, 1.60 mm (0.063\nin.)\n52-71\n93-102\n109-116\n350\n2024-T36\nAll\n76-90\n2024-T4, -T42(d)\nNot clad\n69-83\n100-110\n97-106\n111-118\n85-90\n82.5-87.5\nClad, 1.60 mm (0.063\nin.)\n300\n52-71\n91-100\n109-116\n80-84.5\n50\n60\nClad, >1.60 mm (0.063\nin.)\n52-71\n93-102\n109-116\n2024-T6, -T62\nAll\n74.5-83.5\n99-106\n2024-T81\nNot clad\n74.5-83.5\n99-106\n84-88\n84-88\nFig 37\nClad\n99-106\n2024-T86\nAll\n83-90\n105-110\n87.5-90\n6053-T6\nAll\n79-87\n74.5-78.5\n6061-T4(d)\nSheet\n60-75\n88-100\n64-75\nExtrusions; bar\n70-81\n82-103\n67-78\n\u2022\nNot clad, 0.41 mm\n6061-T6\n(0.016 in.)\n75-84\nNot clad, \u22650.51 mm\n(0.020 in.)\n47-72\n85-97\nClad\n84-96\n6063-T5\nAll\n55-70\n89-97\n62.5-70\n6063-T6\nAll\n70-85\n6151-T6\nAll\n91-102\n7075-T6, -T65\nNot clad(e)\n85-94\n106-114\nClad:\n\u22640.91 mm (0.036 in.)\n102-110\n>0.91 1.27 mm\n(>0.036 0.050 in.)\n78-90\n104-110\n>1.27 1.57 mm\n(>0.050 \u2264 0.062 in.)\n76-90\n104-110\n>1.57 1.78 mm\n(>0.062\u2264 0.070 in.)\n76-90\n102-110\n>1.78 mm (0.070 in.)\n73-90\n102-110\n7079-T6, -T65\n7178-T6\nAll(e)\n81-93\n104-114\nNot clad(f)\n85 min\n105 min\nClad:\n\u22640.91 mm (0.036 in.)\n102 min\n>0.91 1.57 mm\n(>0.036\u2264 0.062 in.)\n>1.57 mm (0.062 in.)\n85 min\n88 min\n78-84\n87.5-92\n86-90\n87.5-92\n88 min\n86 min\n(a) Minimum hardness values shown for clad products are valid for thicknesses up to and including 2.31 mm (0.091 in.); for heavier-gage\nmaterial, cladding should be locally removed for hardness testing or test should be performed on edge of sheet. (b) 126 to 158 HB (10\nmm ball, 500 kg load). (c) 100 to 130 HB (10 mm ball, 500 kg load). (d) Alloys 2024-T4, 2024-T42 and 6061-T4 should not be rejected for\nlow hardness until they have remained at room temperature for at least three days following solution treatment. (e) 136 to 164 HB (10\nmm ball, 500 kg load). (f) 136 HB min (10 mm ball, 500 kg load)\nthe relationships among tensile properties\nand other mechanical properties. In appli-\ncations where other properties are more\nimportant than tensile properties, the other\nproperties should be checked also.\nHardness tests are less valuable for ac-\nceptance and rejection of heat-treated alu-\nminum alloys than they are for steel. Nev-\nertheless, hardness tests have some utility\nfor process control. Typical hardness val-\nues for various alloys and tempers are given\nin Table 13. Figure 37 shows the general\nrelationship between longitudinal tensile\nstrength and hardness for aluminum alloys.\nIntergranular-Corrosion Test. The extent\nof precipitation during elevated-temperature\naging of alloys 2014, 2219, and 2024 markedly\ninfluences the type of corrosion attack and\nthe corrosion resistance. With thin-section\nproducts quenched at rates sufficiently rapid\nto prevent precipitation in the grain bound-\naries during the quench, short periods of\nprecipitation heat treating produce localized\ngrain boundary precipitates adjacent to the\ndepleted areas, producing susceptibility to\nintergranular corrosion. Additional heating,\nhowever, induces extensive general precipi-\ntation within the grains, lowering the corro-\nsion potential differences between the grains\nand the boundary areas, thus removing the\ncause of the selective corrosion.\nThe most common test for susceptibility\nto intergranular corrosion is carried out as\nfollows:\n\u2022\nRemove any cladding by filing or etching\nClean the specimen by immersing it for 1\nmin in a solution containing 5% concen-\ntrated nitric acid and 0.5% hydrofluoric\nacid at a temperature of 95 \u00b0C (200 \u00b0F);\nrinse in distilled water. Immerse for 1 min\nin concentrated nitric acid at room tem-\nperature; rinse in distilled water\n\u2022 Immerse the specimen for 6 h in a freshly\nprepared solution containing 57 g of sodi-\num chloride and 10 mL of 30% hydrogen\nperoxide per liter of water at a tempera-\nture of 30 \u00b1 5 \u00b0C (86 \u00b1 9 \u00b0F). More than\none specimen may be corroded in the\nsame container provided that at least 4.6\nmL of solution is used for each square\ncentimeter (30 mL/in.\u00b2) of specimen sur-\nface and that the specimens are electrical-\nly insulated from each other\n\u2022 After the immersion period, wash the\nspecimen with a soft-bristle brush to re-\nmove any loose corrosion product. Cut a\ncross-sectional specimen at least 19 mm\n(3/4 in.) long through the most severely\ncorroded area; mount and metallographi-\ncally polish this specimen\nExamine the cross-sectional specimen mi-\ncroscopically at magnifications of 100\u00d7\nand 500\u00d7 both before and after etching\nwith Keller's reagent\n\u2022\n\u2022 Describe the results of the microscopic\nexamination in terms of the five degrees\nof severity of intergranular attack illus-\ntrated in Fig 38\nElectrical Conductivity. For control of the\ncorrosion and stress-corrosion characteris-\ntics of certain tempers, notably the T73 and\nT76 types, the materials must meet combi-\nnation criteria of yield strength plus electri-\ncal conductivity. Although these criteria are\nbased on indirect measurements of proper-\nties, their validity for ensuring the intended\ncorrosion and stress-corrosion resistance\nTensile strength, ksi\n878 / Heat Treating of Nonferrous Alloys\nSheet and plate\nSheet, plate, extrusions, and\nforgings\nAlloy\nCondition\nProduct form\n2048\n2124\n2419\n(a)\n(b)\n7049\n7050\n** EE\nT8\nSheet and plate\nT3, T8\nT8\nT7\nT7\n7150\nT6\n7175\nT6, T7\n(c)\n(d)\n7475\nT6, T7\n(e)\nFive degrees of severity of intergranular at-\nFig 38\ntack. Severity of intergranular attack (sche-\nmatic), as observed microscopically in transverse sec-\ntions after test for susceptibility to intergranular\ncorrosion. Top of each area shown in surface exposed\nto corrosive solution\nhas been firmly established by extensive\ncorrelation and testing.\nLow tensile strengths may be accompa-\nnied by high levels of electrical conductiv-\nity, so electrical conductivity is sometimes\nused as a quality-assurance diagnostic\ntool. However, because the correlation\nbetween strength and electrical conductiv-\nity is strongly a function of chemical com-\nposition and fabricating practice, use of\nelectrical conductivity is not recommend-\ned except for rough screening. This\nscreening must be followed by hardness\ntesting, and then by tensile testing if the\nhardness tests indicate that the heat treat-\nment was suspect.\nFracture Toughness Indices. Fracture\ntoughness is rarely, if ever, a design con-\nsideration in the 1000, 3000, 4000, 5000,\nand 6000 series alloys. The fracture tough-\nness of these alloys is sufficiently high that\nthicknesses beyond those commonly pro-\nduced would be required to obtain a valid\ntest.\nFracture toughness is a meaningful de-\nsign-related parameter for some conven-\ntional high-strength alloys and all the con-\ntrolled-toughness, high-strength alloys.\nConventional aerospace alloys for which\nfracture toughness minimums may be use-\nful in design include 2014, 2024, 2219,\n7075, and 7079. These alloys have tough-\nness levels that are inferior to those of\ntheir controlled-toughness counterparts.\nConsequently, these products are not used\nin fracture-critical applications, although\nfracture toughness can be a meaningful\ndesign parameter. Fracture toughness is\nnot guaranteed in conventional high\nstrength alloys.\nFracture toughness quality control and\nmaterial procurement minimums are appro-\npriate for controlled-toughness, high-\nstrength alloys. The alloys and tempers\ncurrently identified as controlled-tough-\nness, high-strength products include:\nPlate, forgings, and extrusions\nSheet, plate, forgings, and\nextrusions\nSheet and plate\nSheet, plate, forgings, and\nextrusions\nSheet and plate\nThe fracture toughness of these alloys and\ntempers range in measured K\u2081c values from\nabout 20 MPa\u221a\u221a\u221am (18 ksi\u221ain.) upward. Con-\ntrolled-toughness alloys are often derivatives\nof conventional alloys. For example, 7475\nalloy is a derivative of 7075 with maximum\ncompositional limits on some elements that\nwere found to decrease toughness.\nIn products of the newer controlled-\ntoughness high-strength alloys 2090, 2091,\n2124, 2224, 2324, 7050, 7149, 7150, 7175,\n7475, and 8090, which provide guaranteed\nlevels of fracture toughness, minimum val-\nues of the applicable indices, K\u2081c or Kc,\nare established by accumulation of statis-\ntical data from production lots as a basis\nfor guaranteed minimum values. If the\nminimum specified fracture toughness val-\nue is not attained, the material is not\nacceptable. Some specifications allow use\nof less-expensive screening tests (such as\nthe notch tensile or chevron-notched short\nbar) as a basis for release of high-tough-\nness alloy products. In these instances,\ncorrelations between K\u2081 and the screening\ntest result is used to establish the appro-\npriate notch-yield ratio as a lot-release\ncriterion.\nTemper Designations for\nHeat-Treatable Aluminum Alloys\nThe temper designations used in the Unit-\ned States for heat-treatable aluminum alloys\nare part of the system that has been adopted\nas an American National Standard (ANSI\nH35.1). Used for all wrought and cast prod-\nuct forms except ingot, the system is based\non the sequences of mechanical or thermal\ntreatments, or both, used to produce the\nvarious tempers. The temper designation\nfollows the alloy designation and is separat-\ned from it by a hyphen. Basic temper des-\nignations consist of individual capital let-\nters. Major subdivisions of basic tempers,\nwhere required, are indicated by one or\nmore digits following the letter. These digits\ndesignate specific sequences of treatments\nthat produce specific combinations of char-\nacteristics in the product. Variations in\ntreatment conditions within major subdivi-\nsions are identified by additional digits. The\nconditions during heat treatment (such as\ntime, temperature, and quenching rate)\nused to produce a given temper in one alloy\nmay differ from those employed to produce\nthe same temper in another alloy.\nDesignations for the common heat-treat-\ned tempers, and descriptions of the se-\nquences of operations used to produce\nthose tempers, are given in the following\nparagraphs. (For the entire aluminum alloy\ntemper designation system, including desig-\nnations for non-heat-treatable alloys, see\nVolume 2 of this Metals Handbook series.)\nBasic temper designations for heat-treated\nconditions include the codes O, W, and T.\nOther basic temper designations are F (as\nfabricated) and H (strain hardened).\nO, annealed. Applies to wrought prod-\nucts that are annealed to obtain lowest\nstrength temper and to cast products that\nare annealed to improve ductility and di-\nmensional stability. The O may be followed\nby a digit other than zero.\nW, solution heat treated. An unstable\ntemper applicable to any alloy that naturally\nages (spontaneously ages at room tempera-\nture) after solution heat treatment. This\ndesignation is specific only when the period\nof natural aging is indicated-for example,\nWh. (See also the discussion of the Tx51,\nTx52, and Tx54 tempers, in the section\nbelow on subdivision of the T temper.)\nT, heat treated to produce stable tempers\nother than O. Applies to products that are\nthermally treated, with or without supple-\nmentary strain hardening, to produce stable\ntempers. The T is always followed by one or\nmore digits, as discussed below.\nMajor Subdivisions of T Temper. In T-type\ndesignations, the T is followed by a number\nfrom 1 to 10; each number denotes a specif-\nic sequence of basic treatments, as de-\nscribed below.\nTI, cooled from an elevated-temperature\nshaping process and naturally aged to a\nsubstantially stable condition. Applies to\nproducts that are not cold worked after an\nelevated-temperature shaping process such\nas casting or extrusion, and for which me-\nchanical properties have been stabilized by\nroom-temperature aging. If the products are\nflattened or straightened after cooling from\nthe shaping process, the effects of the cold\nwork imparted by flattening or straightening\nare not recognized in specified property\nlimits.\nT2, cooled from an elevated-temperature\nshaping process, cold worked, and natural-\nly aged to a substantially stable condition.\nApplies to products that are cold worked\nspecifically to improve strength after cool-\ning from a hot-working process such as\nrolling or extrusion, and for which mechan-\nical properties have been stabilized by\nroom-temperature aging. The effects of cold\nwork, including any cold work imparted by\nflattening or straightening, are recognized in\nspecified property limits.\nT3, solution heat treated, cold worked,\nand naturally aged to a substantially stable\ncondition. Applies to products that are cold\nworked specifically to improve strength af-\nter solution heat treatment, and for which\nmechanical properties have been stabilized\nby room-temperature aging. The effects of\ncold work, including any cold work impart-\ned by flattening or straightening, are recog-\nnized in specified property limits.\nT4, solution heat treated and naturally\naged to a substantially stable condition.\nApplies to products that are not cold\nworked after solution heat treatment, and\nfor which mechanical properties have been\nstabilized by room-temperature aging. If the\nproducts are flattened or straightened, the\neffects of the cold work imparted by flatten-\ning or straightening are not recognized in\nspecified property limits.\nT5, cooled from an elevated-temperature\nshaping process and artificially aged. Ap-\nplies to products that are not cold worked\nafter an elevated-temperature shaping pro-\ncess such as casting or extrusion, and for\nwhich mechanical properties or dimensional\nstability, or both, have been substantially\nimproved by precipitation heat treatment. If\nthe products are flattened or straightened\nafter cooling from the shaping process, the\neffects of the cold work imparted by flatten-\ning or straightening are not recognized in\nspecified property limits.\nT6, solution heat treated and artificially\naged. Applies to products that are not cold\nworked after solution heat treatment, and\nfor which mechanical properties or dimen-\nsional stability, or both, have been substan-\ntially improved by precipitation heat treat-\nment. If the products are flattened or\nstraightened, the effects of the cold work\nimparted by flattening or straightening are\nnot recognized in specified property limits.\nT7, solution heat treated and stabilized.\nApplies to products that have been precip-\nitation heat treated to the extent that they\nare overaged. Stabilization heat treatment\ncarries the mechanical properties beyond\nthe point of maximum strength to provide\nsome special characteristic, such as en-\nhanced resistance to stress-corrosion crack-\ning or to exfoliation corrosion.\nT8, solution heat treated, cold worked,\nand artificially aged. Applies to products\nthat are cold worked specifically to improve\nstrength after solution heat treatment, and\nfor which mechanical properties or dimen-\nsional stability, or both, have been substan-\ntially improved by precipitation heat treat-\nment. The effects of cold work, including\nany cold work imparted by flattening or\nstraightening, are recognized in specified\nproperty limits.\nT9, solution heat treated, artificially\naged, and cold worked. Applies to products\nthat are cold worked specifically to improve\nstrength after they have been precipitation\nheat treated.\nT10, cooled from an elevated-tempera-\nture shaping process, cold worked, and\nartificially aged. Applies to products that\nare cold worked specifically to improve\nstrength after cooling from a hot-working\nprocess such as rolling or extrusion, and for\nwhich mechanical properties or dimensional\nstability, or both, have been substantially\nimproved by precipitation heat treatment.\nThe effects of cold work, including any cold\nwork imparted by flattening or straighten-\ning, are recognized in specified property\nlimits.\nOther Subdivisions of T Temper Codes for\nStress-Relieved Products. When it is desir-\nable to identify a variation of one of the ten\nmajor T tempers described above, addition-\nal digits, the first (x) of which cannot be\nzero, may be added to the designation.\nThe following specific sets of additional\ndigits have been assigned to stress-relieved\nwrought products.\nTx51, stress relieved by stretching. Ap-\nplies to the following products when\nstretched to the indicated amounts after\nsolution heat treatment or after cooling\nfrom an elevated-temperature shaping pro-\ncess:\nProduct form\nPlate\nRod, bar, shapes, extruded\ntube\nDrawn tube\nPermanent set, %\n112-3\n1-3\n1/2-3\nTx51 applies directly to plate and to rolled\nor cold finished rod and bar. These products\nreceive no further straightening after\nstretching. Tx51 also applies to extruded\nrod, bar, shapes, and tubing, and to drawn\ntubing, when designated as follows:\n\u2022\nTx510. Products that receive no further\nstraightening after stretching\n\u26ab Tx511. Products that may receive minor\nstraightening after stretching to comply\nwith standard tolerances\n\u26ab Tx52. Stress relieved by compressing. Ap-\nplies to products that are stress relieved\nby compressing after solution heat treat-\nment, or after cooling from a hot-working\nprocess to produce a permanent set of 1\nto 5%\n\u2022 Tx54. Stress relieved by combining\nstretching and compressing. Applies to\ndie forgings that are stress relieved by\nrestriking cold in the finish die. (These\nsame digits and 51, 52, and 54-may be\nadded to the designation W to indicate\nunstable solution heat-treated and stress-\nrelieved tempers)\nHeat Treating of Aluminum Alloys / 879\nTemper designations T42 and T62 have\nbeen assigned to wrought products heat\ntreated from the O or the F temper to\ndemonstrate response from the heat treat-\nment described below. Temper designations\nT42 and T62 also may be applied to wrought\nproducts heat treated from any temper by\nthe user when such heat treatment results in\nthe mechanical properties applicable to\nthese tempers.\n\u2022 T42. Solution heat treated from the O or\nthe F temper to demonstrate response to\nheat treatment and naturally aged to a\nsubstantially stable condition\n\u2022 T62. Solution heat treated from the O or\nthe F temper to demonstrate response to\nheat treatment and artificially aged\nSubdivision of the O Temper. In temper\ndesignations for annealed products, a digit\nfollowing the O indicates special character-\nistics. For example, 01 denotes that a prod-\nuct has been heat treated according to a\ntime/temperature schedule approximately\nthe same as that used for solution heat\ntreatment, and then air cooled to room\ntemperature, to accentuate ultrasonic re-\nsponse and provide dimensional stability;\nthis designation applies to products that are\nto be machined prior to solution heat treat-\nment by the user.\nREFERENCES\n1. S. Hirano et al., Quench Sensitivity in\nAl-Li Based Alloys, Proceedings of Con-\nference on Aluminum-Lithium Alloys\n(Vol 1), Materials and Component Engi-\nneering Publications, 1989, p 335-344\n2. T. Sheppard, Mater. Sci. Technol., Vol\n4, July 1988, p 636\n3. J.E. Hatch, in Aluminum Properties and\nPhysical Metallurgy, American Society\nfor Metals, 1984, p 165-166\n4. C.E. Bates, Selecting Quenchants to\nMaximize Tensile Properties and Mini-\nmize Distortion in Aluminum Parts, J.\nHeat Treat., Vol 5 (No. 1), 1987, p 27-40\n5. T. Croucher, Critical Parameters for\nEvaluating Polymer Quenching of Alu-\nminum, Heat Treat., Vol 19 (No. 12),\nDec 1987, p 21-25\n6. W.L. Fink and L.A. Willey, Quenching\nof 75S Aluminum Alloy, Trans. AIME,\nVol 175, 1948, p 414-427\n7. J.W. Evancho and J.T. Staley, Kinetics\nof Precipitation in Aluminum Alloys dur-\ning Continuous Cooling, Metall. Trans.\nA, Vol 5A, Jan 1974, p 43-47\n8. J.T. Staley, Industrial Heating XLIV,\nOct 1977, p 6-9\n9. G.F. Bobart, J. Heat Treat., Vol 6 (No.\n1), 1988, p 47-52\n"}, "expected_output": {"claims": [{"unit": "h", "value": 24, "evidence": ["7075(i)", "Sheet", "T6", "24", "nominal aging practice of 24 h at 120 \u00b0C (250 \u00b0F) affect the strength of 7075-T6", "7150, aged 24 h at 120 \u00b0C", "7050-W (4 days) extrusions were aged at 24 h at 120 \u00b0C (250 \u00b0F)"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_89c5ce1e87c296bc"}}
{"id": "0c691392d23445742288db62", "input": {"query": "What is the total water consumption including hydrolysis water additions and steam injection in kg water per kg IPA produced?", "source_url": "https://mohdazizan.files.wordpress.com/2010/06/presentation-new-edit.pdf", "document_text": "ISOPROPYL ALCOHOL\n\n\nTHE PROCESS FLOW DIAGRAM IPA\n\n\nIUPAC\n\n\nECONOMIC\n\n\nPRODUCTION\n\n\nISOPROPYL\nALCOHOL\n\n\nHEALTH\nAND SAFETY\n\n\nCHEMICAL\nPROPERTIES\n\n\nPHYSICAL\n\n\nUSES\n\n\nPROPERTIES\n\n\nPHYSICAL PROPERTIES\n\n\nODOUR\n\n\n-STRONG\n\n\nSOLUBILITY\n\n\nALCOHOL\n\n\nCOLOUR\n\n\n-SOLUBLE IN\n\n\nALL ORGANIC\n\n\n-COLOURLESS\n\n\n-INSOLUBLE IN\n\n\nSALT\n\n\nMOLAR MASS\n\n\nPhysical\nproperties\n\n\nBOILING\n\n\n-60.09\n\n\nPOINT\n\n\nmol 1\n\n\ng\n\n\n82.5\u00b0C\n\n\nFairly\n\n\nMICSIBLE\n\n\nvolatile\n\n\nevaporates\neasily\n\n\nIN WATER\n\n\nMELTING\n\n\nPOINT\n\n\n-89\u00b0C\n\n\nDehydrogenation\n\n\nOxidation\n\n\nEtherification\n\n\n\u2022Halogenations\n\n\nCHEMICAL PROPERTIES\n\n\nPRODUCTION METHOD\n\n\n* Direct\n\n\nIndirect\n\n\nSOUTHWAY\n\n\nSOUTH\n\n\nAM\n\n\nUSES\n\n\n* pharmaceutical industry\n\n\n\u00d7 Construction field -wood\n\n\n* Plantation industry -pesticides\n\n\n* Cosmetics -baby oils, lotions, etc.\n* Aerosol -hair spray, detergents, etc.\n* Removing sticky residue\n\n\nHEALTH AND SAFETY\n\n\n*Industry (not harm)\n\n\n*Highly\n\n\nflammable/explode-away\n\n\nfrom heat, open flame, ignition\n\n\nsources and incompatibles\n\n\n*Defect to human skin -use glove\n\n\nECONOMICS\n\n\n*rose by 5% to $1240/tonne from\n$1200/tonne in mid-April\n\n\n*Geo Chem Trading located in\nPenang Port, Butterworth Penang.\n\n\nBLOCK DIAGRAM\nFOR INDIRECT\n\n\nHYDRATION OF\n\n\nI SOPROPYL\n\n\nALCOHOL\n\n\nINPUT (RAW MATERIAL)\n\n\n-SULPHURIC ACIDS\n- PROPYLENE GAS\n\n\nUNABSORBED\n\n\nREACTION OF\n\n\nSULPHURIC ACID WITH\n\n\nPROPELYNE GAS-\nRECYCLED\n\n\nPROPYLENE\n\n\nMIXER CHAMBER-\nWATER FEED\n\n\nFLASH ZONE\n\n\nSEPARATION\n\n\nMIXING CHAMBER-\nWATER FEED\n\n\nACID\n\n\nIPA\nGENERATION\n\n\nRECONCENTRATION\n\n\n-RECYCLED\n\n\nDISTILLATION COLUMN\n-ETHER COLUMN\n-DRIYING COLUMN\n\n\nOUTPUT\n\n\n-DIISOPROPYL ETHER\n-ISOPROPYL ALCOHOL IN CDM\n\n\nAND ANHYDROUS\n\n\nPROCESS FLOW\n\n\nDIAGRAM\n\n\nINDIRECT HYDRATION\n\n\nMETHOD\n\n\nPROPYLENE\nCOLUMN\nC104\n\n\nWATER\nDIISOPROPYL ETHER\n\n\n18\n\n\n28\n\n\nSODIUM HYDROXIDE\nFEED\n\n\nENTRAINER\n\n\nSCRUBBER\nC103\n\n\nC100\n\n\n21\n\n\nDECANTER\nF101\n\n\nC102\n\n\nWASH WATER\n\n\nDIISOPROPYL\nCOLUMN\nC109\n\n\nPROPYELENE\nFEED\n\n\nPROPANE\n\n\nABSORBER\nC100\n\n\nC101\n\n\nCONCENTARTED\nSULFURIC ACID FEED\n\n\n22\n\n\nIPA\nGENERATIONS\nC102\n\n\nWATER\nLAYER\nSTRIPPERS\nC108\n\n\n20\n\n\nETHER\nCOLUMN\nC105\n\n\nALCOHOL\nCOLUMN\nC106\n\n\nDRYING\nCOLUMN\nC107\n\n\nSTEAM\n\n\nMIXER\nM100\n\n\nMIXER\nM101\n\n\nCRUDE ISOPROPYL\nALCOHOL TANK\nF100\n\n\nFLASH VESSEL\nC101\n\n\nWATER\n\n\nSULPHURIC ACID\nTO ACID CONTETRATOR\n\n\nWATER\nFEED\n\n\nWATER\n\n\nCBM OR\nAZEOTROPIC\nISOPROPYL\nALCOHOL\n\n\nWATER\nFEED\n\n\n12\n\n\n11\n\n\nANHYDROUS\nISOPROPYL\nALCOHOL\n\n\nV100\n\n\nP100\n\n\nTitle: PFD study of the Production of Isopropyl Alcohol-indirect hydration method\n\n\nINITIAL INPUT:-\n\n\nPRODUCT:-\n\n\nGROUP MEMBERS:\n\n\nRAW MATERIALS CONSIST OF\n\n\nUNIVERSITI\nKEBANGSAAN\nMALAYSIA\n\n\n1. ANHYDROUS ISOPROPYL ALCOHOL\n\n\n1.MOHD AZIZAN BIN RAZAL @ ABD RAZI\n\n\nI. SULPHURIC ACIDS (60%-70% WT)\n\n\n2.AZETROPE ISOPROPYL ALCOHOL\n\n\n2.CHONG KIN HOW\n\n\n2. PROPYLENE GAS (45% WT)\n\n\n3. DIISOPROPYL ETHER\n\n\n3.NORASMANIRA BINTI MUSA\n\n\n4.NORFATYHAH BINTI ABD AZIZ\n\n\nNational University of Malaysia\n\n\nNUR IZAYU BINTI BADARUDDIN\n\n\nPFD DESCRIPTION\n\n\nSTEP 1\n\n\nStream 1 (concentrated Sulphuric Acid(60-70\nwt%)) + stream 2(Propylene Gases (45wt%))\n\n\n. Occur in the absorber C100\n\n\n. Condition for the absorber C100 are:\n\n\n1. At the pressure 150-300 psig.\n\n\n2. At temperature 85-115\u00b0C.\n\n\nPROPYELENE\nFEED\n\n\nABSORBER\n\n\nC100\n\n\nCONCENTARTED\nSULFURIC ACID FEED\n\n\nPFD DESCRIPTION\n\n\nSTEP 2\n\n\nextract liquid in the absorber\n\n\nStream 1+ Stream 2\n\n\nC100.\n\n\nExtract liquid consist of:-\n\n\n40 to 65 wt% of Sulphuric Acid,\n\n\n1.\n\n\nabout 10 to 40 wt% of absorbed Propylene values,\n3. balance water,\n\n\n2.\n\n\n4. Absorbed propylene-all molecule in the liquid (propyl\n(CH), and propylene (CH))\n\n\n5. IPA at 10-45 wt%,\n\n\n6. Diisopropyl Ether at 3-6 wt%,\n\n\nThe extract liquid is withdrawn from the bottom portion of\nabsorber via Stream 3 to the mixer M100.\n\n\nENE\n\n\nABSORBER\nC100\n\n\nMIXER\nM100\n\n\nPFD DESCRIPTION\n\n\nSTEP 3\n\n\nUnabsorbed gas- Propylene gas are withdrawn\nand compressed to propylene column by c104 via\nstream 4.\nIt will refined- produced recovered Propylene Gas\nand some Propane.\n\n\n\u2022 The recovered gas-being recycled by stream 5.\n\n\nPROPYLENE\nCOLUMN\n\n\nC104\n\n\nSODIUM HYDROXIDE\n\n\nFEED\n\n\nSCRUBBER\n\u0421109\n\n\nC100\n\n\nC102\n\n\nPROPYELENE\n\n\nPROPANE\n\n\nFEED\n\n\nPFD DESCRIPTION\n\n\nSTEP 4\n\n\n. In the mixer M100.\n\n\nStream 3 mixes with stream 7 which feed with\nwater.\n\n\n2. Water is added in an amount from about 1.5 to 6\nwt% water, based on the weight of absorbed\nPropylene values in the extract liquid.\n3. It will make sure the moles of water\n\n\nstoichiometrically required to form the amount\nof Diisopropyl Ether and Isopropyl Alcohol\nwithdrawn as vapor in the flash vessel C101.\n\n\nMIXER\nM100\n\n\nFLASH VESSEL\n\n\n[\n\n\nC101\n\n\nW\n\n\nF\n\n\nWATER\nFEED\n\n\nPFD DESCRIPTION\n\n\nSTEP 5\n\n\nFlash vessels C101 is operate in the condition of:-\n1. At temperature between 80 to 160\u00baC.\n\n\n2. At pressure from about 5 to 10 psig.\n\n\nThe overhead products is withdrawn and being compressed by\nC101 via stream 10 comprising :-\n\n\n1. Diisipropyl Ether from about 30 to 50 wt%,\n\n\n2. Isopropyl Alcohol from about 30 to 40 wt%,\n\n\n3. Propylene from about 10 to 20 wt%.\n\n\nLiquid which is withdrawn from the flash vessel via stream 9\ncomprises the depleted extract :-\n\n\n1. Absorbed Propylene - 5 to 35 wt%,\n2. Sulphuric acids-45 to 65 wt%,\n\n\n3. Balance water.\n\n\nC101\n\n\n8\n\n\nFLASH VESSEL\n\n\nC101\n\n\nWAT\n\n\nFEE\n\n\nPFD DESCRIPTION\n\n\nSTEP 6\n\n\n\u26ab Depleted extract have 2 portions:-\n\n\nFirst portion (40 to 60 wt% of total) will be recycled, pump\nby P100 through stream 11 and 1 to absorber C100.\n\n\nA control valve V100 -increase or decrease liquid flow between\nstream 11 and 12 and control Diisopropyl Ether and\nIsopropyl Alcohol production.\n\n\n\u25ba Second portion is sent to second water mixing zone M101.\n\n\nWATER\nFEED\n\n\n\u2611\n\n\nV100\n\n\nP100\n\n\nPFD DESCRIPTION\n\n\nSTEP 7\n\n\n. The depleted liquids in stream 12 flow to mixer M101.\nIn the mixer M101:-\n\n\nDepleted extract mixes stream 13 which feed with water (35\nto 100 parts by weight of water per part by weight of\n\n\nabsorbed Propylene Values).\n\n\n\u25cfThe liquid residence time in water mixing zone ranges from\nabout 1 to 30 minutes.\n\n\n13\n\n\nMIXER\n\n\nM101\n\n\nWATER\nFEED\n\n\n12\n\n\nPFD DESCRIPTION\n\n\nSTEP 8\n\n\nThe mixture sent via stream 14 to alcohol generator C102\nto be heated.\n\n\nDirect injection steam at stream 15 to generator.\n\n\nUnit operation for Alcohol generator are:-\n\n\n1. At temperature about 70 to 130\u00baC\n\n\n2. Pressure about 1 to 5 psig\n\n\nAlcohol vapor are formed as overhead product and sent to\nscrubber C103 via stream 17-comprised IPA and Diisopropyl\nEther.\n\n\nSulphuric acid about 45 to 55 wt% is withdrawn from lower\nportion of alcohol generator to acid concentrator via stream\n16.\n\n\nC102\n\n\nIPA\n\n\nGENERATIONS\n\n\nC102\n\n\n15\n\n\n14\n\n\nSTEAM\n\n\n16\n\n\nMIXER\nM101\n\n\nSCRUI\n\n\nALI\n\n\nSULPHURIC ACID\n\n\nTO ACID CONTETRATOR\n\n\nPFD DESCRIPTION\n\n\nSTEP 9\n\n\nProduct stream 10 mixed with product in stream 17-\n\n\nflow to the scrubber C103.\n\n\nProduct in scrubber have been neutralized using NaOH at\nconcentration 10 wt% of the alcohol and Sulphuric mixture.\nUnit operation of the Scrubber C103:-\n\n\n1. At temperature 373K.\n\n\n2. Pressure 14.7 psi.\n\n\nThe product at 55 wt% IPA is withdrawn from a lower part to\ncrude isopropyl alcohol tank F100.\n\n\nFrom the tank, mixture is sent to ether column C105\n\n\n18\n\n\nSODIUM HYDROXIDE\n\n\nFEED\n\n\nSCRUBBER\n\n\nC103\n\n\nC102\n\n\n19\n\n\nC101\n\n\nIPA\n\n\nPFD DESCRIPTION\nSTEP 9-continue\n\n\nMixture in F100 will send to ether column C105 via\n\n\nstream 20.\n\n\n\u26ab Unit operations for Ether column C105 are:-\n\n\n1. At the temperature 362 K.\n\n\n2. Pressure 14.7 psi.\n\n\nDistillation take place to separate Diisopropyl Ether and\nHexene.\n\n\nDiisopropyl Ether is withdrawn to Diisopropyl Column C109.\nWashed to produce washed Ether at 95 wt.%.\n\n\nPROPYLENE\n\n\nCOLUMN\n\n\nWATER\n\n\nC104\n\n\nDIISOPROPYL ETHER\n\n\nDE\n\n\nSCRUBBER\n\n\nC103\n\n\n21\n\n\n8\n\n\nWASH WATER\n\n\nDIISOPROPYL\nCOLUMN\nC109\n\n\nPROPANE\n\n\n19\n\n\n22\n\n\nETHER\nCOLUMN\n\n\nALCOHOL\nCOLUMN\n\u0421106\n\n\n20\n\n\nC105\n\n\nSTEAM\n\n\n25\n\n\nCRUDE ISOPROPYL\nALCOHOL TANK\nF100\n\n\nWATER\n\n\nPFD DESCRIPTION\n\n\nSTEP 10\n\n\nWet IPA from Ether column C105 will be send to Alcohol\ncolumn C106 via stream 22.\n\n\n\u26ab Unit operation for Alcohol Column C106:-\n\n\n1. At pressure about 3 to 7 atm.\n\n\nAzeotroping agent is feed near the top of column C106 via\nstream 23 which is produce in drying column C107.\n\n\n\u26ab CBM or azeotropic isopropyl alcohol at 87 wt% is\nwithdrawn from upper part of alcohol column via stream 24\n(product) while water is withdrawn from bottom of column\nvia stream 25.\n\n\nWASH WATER\n\n\nDIISOPROPYL\n\n\nCOLUMN\nC109\n\n\n23\n\n\n22\n\n\nALCOHOL\nCOLUMNN\nC106\n\n\n24\n\n\n25\n\n\nWATER\n\n\nCBM OR\nAZEOTROPIC\n\n\nISOPROPYL\n\n\nALCOHOL\n\n\nPFD DESCRIPTION\n\n\nSTEP 11\n\n\nUnit operation for Drying column C107:-\n\n\n1. At pressure 14.7 psi.\n\n\n2. Temperature 347 K.\n\n\nProduct form: Ternery Azeotrope.\n\n\nLatter concentrate at head column in concentration about 60%\nis piped into condenser.\n\n\nLiquefaction of vapor enter decanter F101 to separate the\nlayers into virtually water-free upper layer and lower layer\ncontaining the res\u00eddual water.\n\n\nThe upper layers in F101 (mainly azeotroping agent and\nalcohol) is return to the top of column C107.\n\n\nWhile anhydrous Isopropyl Alcohol at 99.8 wt% is removed\nfrom the lower part of column C107 via stream 29.\n\n\nEzz\n\n\nENTRAINER\n\n\nDECANTER\nF101\n\n\nWATER\nLAYER\nSTRIPPERS\nC108\n\n\nDRYING\nCOLUMN\nC107\n\n\nWATER\n\n\nANHYDROUS\nISOPROPYL\nALCOHOL\n\n\nPFD DESCRIPTION\n\n\nSTEP 12\n\n\nThe lower layer in the decanter F101 is mostly water.\nIt is fed into stripping column C108.\n\n\nUnit operation for Stripping column C108:-\n\n\n1. a temperature 466K.\n\n\n2. pressure about 155 psi.\n\n\nThe waste water in F101 is send to C108 via stream 30\nfor recovery of Isopropyl Alcohol and azeotroping agent.\nThis product is send back into column C107 via stream\n31 and waste the water at stream 32.\n\n\nCONCLUSION\n\n\n\u2610 The indirect process = Propene with sulfuric\nacid to form a mixture of sulfate esters and\nsubsequent hydrolysis will produce isopropyl\nalcohol.\n\n\n\u25a0 Direct hydration = Propene and water, either\nin gas or liquid phases, at high pressures in the\npresence of solid or supported acidic catalysts\n\n\nUse of IPA in direct solvent applications accounted for 62% of\ntotal IPA demand in 2008.\n\n\nGlobal IPA-based acetone production is expected to decrease\nwith the increase of phenol capacity (acetone is a coproduct\nof phenol by the cumene peroxidation process). Acetone will\nbe producing after 2010 in Western Europe.\n\n\nIn 2008, global supply/demand was relatively balanced as\ndemand decreased because of the weakened global economy.\n\n\nThe future demand of IPA is expected to remain flat or to\ngrow slightly.\n\n\nIts main usage as a chemical intermediate is growing, and this\nshould offset the pressure on use as a solvent from tighter\nvolatile organic chemicals (VOC) regulations\n\n\nTHE END\n"}, "expected_output": {"claims": [{"unit": "wt%", "value": 1.5, "evidence": ["Water is added in an amount from about 1.5 to 6 wt% water, based on the weight of absorbed Propylene values in the extract liquid.", "2. Water is added in an amount from about 1.5 to 6\nwt% water, based on the weight of absorbed\nPropylene values in the extract liquid.\n3. It will make sure the moles of water", "stoichiometrically required to form the amount\nof Diisopropyl Ether and Isopropyl Alcohol\nwithdrawn as vapor in the flash vessel C101."]}, {"unit": "wt%", "value": 6, "evidence": ["Water is added in an amount from about 1.5 to 6 wt% water, based on the weight of absorbed Propylene values in the extract liquid.", "2. Water is added in an amount from about 1.5 to 6\nwt% water, based on the weight of absorbed\nPropylene values in the extract liquid.\n3. It will make sure the moles of water", "stoichiometrically required to form the amount\nof Diisopropyl Ether and Isopropyl Alcohol\nwithdrawn as vapor in the flash vessel C101."]}]}, "metadata": {"product_category": "Chemical products", "request_id": "req_bfeeb4c1fd409542"}}
{"id": "ff43bd9e0d987a6ac5cd2ec4", "input": {"query": "What is the total mass loss rate (in percentage or g/g) from virgin glass batch during melting based on TGA data?", "source_url": "https://www.hanford.gov/files.cfm/Effect%20of%20Alumina%20....pdf", "document_text": "Effect of alumina source on ease of melting of glass batch\n\n\nDavid A. Pierce, Pavel Hrma, Jos\u00e9 Marcial, Brian J. Riley, Michael J. Schweiger\n\n\nPacific Northwest National Laboratory, Richland, WA 99354\n\n\nAbstract\n\n\nThe selection of raw materials affects the rate of batch-to-glass conversion. In all-electric melters,\nfoam under the batch blanket limits the heat flux from the molten glass, thus slowing the rate of melting.\nOur study, in which we compare the melting behaviors of three batches formulated to vitrify high-alumina\nhigh-level waste, shows that a slowly dissolving refractory component can cause excessive foaming.\nFaster melting batches with gibbsite [Al(OH)3] or boehmite [AlO(OH)] as an alumina source produced\nsubstantially less foaming than a batch with corundum (Al2O3). While gibbsite and boehmite dissolved\nbelow 500\u00b0C, corundum was still present in the batch up to 900\u00b0C; hence, the glass-forming melt lacked\nalumina in the batch with corundum. The low viscosity of that batch caused the open pores to close\nprematurely at 660\u00b0C, trapping gases and expanding to foam. This would explain the literature-reported\nslow melting rate of a batch with corundum, as compared to batches with gibbsite and boehmite.\n\n\nIntroduction\u00b9\n\n\nOur study was motivated by the effect of alumina source on the rate of melting observed in [1], a\n\n\npublicly available report that provides a summary of studies of the effects of composition on the rate of\n\n\n\u00b9 This work was supported by the U.S. Department of Energy Federal Project Office Engineering Division for the\n\n\n1\n\n\nHanford Tank Waste Treatment and Immobilization Plant.\n\n\nmelting in continuous electric melters of various scales conducted to assess the performance of the\n\n\nHanford Waste Treatment and Immobilization Plant currently under construction in Washington State.\nFor information regarding nuclear waste vitrification, we refer the reader to the review by Vienna [2] and\n\n\nthe literature cited there.\n\n\nGenerally, the selection of batch materials affects the rate of glass melting. In particular, additions\nof refractory components, such as crystalline alumina (corundum) [1], zinc oxide [3], or zirconia\n(baddeleyite) decrease the rate of melting of glass batches in continuous melters. This effect can be\nconsiderable. Table 1 lists the results of melter experiments with three batches reported in [1]. These\nbatches were formulated to vitrify a simulated high-alumina high-level waste and each was prepared with\na different alumina precursor: one with gibbsite [Al(OH)3], another with boehmite [AlO(OH)], and the\nother with corundum (Al2O3). Table 1 shows that batches with gibbsite and boehmite melted substantially\nfaster than the batch prepared with corundum.\n\n\nThe list of alumina sources used for commercial glasses is extensive [4]. In nuclear waste glasses,\nalumina can be a waste component, typically in the form of gibbsite and boehmite, or a glass-forming\nadditive. Whereas alumina rarely is a major component in commercial glasses, nuclear waste glass can\ncontain up to 30 mass % of alumina [5].\n\n\nThe considerable effect of batch materials on the rate of melting has important economic\nconsequences. The life cycle of high-level waste vitrification at Hanford depends on the rate of melting,\nand thus will be affected by the selection of glass-forming and modifying additives used for waste\nvitrification. Additionally, the chemical and mineralogical form of waste simulants influences prediction\nof the performance of waste-glass melters. Thus, corundum appears a poor substitute for gibbsite and\nboehmite, which are common in Hanford high-level wastes.\n\n\nThe very fact that the corundum, or any refractory oxide, slows down the melting of glass is\n\n\nintriguing because it can hardly be attributed to the refractoriness alone if the residence time of glass in\nthe melter is long enough to allow a refractory oxide to dissolve and homogenize.\n\n\nIn an all-electric melter, the rate of melting of the batch blanket, or the cold cap, is determined by\nthe heat flux delivered to the cold cap from the pool of molten glass on which the cold cap floats.\nAccordingly, the presence of a refractory oxide, such as corundum, must affect the heat transfer to or\nwithin the cold cap.\n\n\nIndeed, laboratory examination revealed that the batches with corundum produced copious foam\nwhereas the batches with gibbsite did not [6, 7]. Foam accumulated under the cold cap insulates the cold\ncap from the melt, thus reducing the heat flow to the cold cap [8-11]. The intriguing question is about the\nmechanism by which corundum, a stable oxide that does not participate in any gas-evolving reaction,\nintensifies foam accumulation.\n\n\nTwo types of foam are associated with glass melting. Primary foam, term coined by Gerrard and\nSmith [12], is batch expansion caused by the evolution of batch-reaction gases, such as CO2, after enough\nglass-forming melt was produced in the batch to close the open pores. Secondary foam is generally a term\nreserved for foam created in molten glass from fining-reaction gases [12, 13]. In high-level-waste glass\nmelts, secondary foam results from oxidation-reduction reactions, mainly involving ferric oxide. Neither\nfoam is directly associated with the dissolution of corundum in molten glass.\n\n\nThis study employs several experimental techniques to compare the melting behavior of three\nhigh-alumina batches with identical compositions except for sources of alumina. It is focused on the\ndifference between batches containing corundum and gibbsite, although experiments were also performed\nwith a batch containing boehmite. As argued in the Discussion section, excessive foaming of the\n\n\ncorundum-containing batch can be attributed to the early closure of open pores caused by the low\nviscosity of the glass-forming melt.\n\n\nExperimental Approach\n\n\nTable 2 shows the composition of a batch used in previous studies [10, 14] with gibbsite, Al(OH)3\n(Almatis, Lot No. 0724131923), as the alumina source. The other two batches examined in this study\ncontained equivalent fractions (g per g glass) of corundum (Sigma-Aldrich\u00ae, Lot No. 342688) and\nboehmite (Nabaltec, Lot No. 047235), i.e., 0.240 of Al2O3 and 0.280 of AlO(OH). Fractions of all other\ncomponents remained identical. All batches were made with quartz particles of 75 \u03bcm as the source of\nsilica. The particle size of corundum, by image analysis, was 45.6 \u00b1 8.6 \u03bcm. The glass composition, in\nterms of targeted mass fractions, was SiO2 (0.305), Al2O3 (0.240), B2O3 (0.152), Na\u2082O (0.096), CaO\n(0.061), Fe2O3 (0.059), Li\u2082O (0.036), Bi\u2082O3 (0.011), P2O5 (0.011), F (0.007), Cr2O3 (0.005), PbO (0.004),\nNiO (0.004), ZrO2 (0.004), SO3 (0.002), K\u2082O (0.001), MgO (0.001), and ZnO (0.001). As described\nelsewhere [14], batches were prepared as slurry and were dried at 105\u00b0C overnight in an oven.\n\n\nTo measure the expansion of a batch during heating, ~1.5 g of batch was pressed at ~7 MPa into a\ncylindrical pellet ~13 mm in diameter and ~6.5 mm high. Each pellet was placed into a Deltech\u00ae (Denver,\nCO, U.S.A) furnace with a silica-glass viewing window and heated at 5\u00b0C/min on an alumina plate to\n1000\u00b0C. Pellets were photographed at regular intervals, and their profile areas were measured with\n\n\n\u24c7\n\n\nPhotoshop (Adobe\u00ae, San Jose, CA, U.S.A.). Next to the pellet was a Pt wire, 10-mm in length, which\n\n\nwas used as a scale gauge. The profile area, A, was normalized to the profile area of a hemispherical body\nof the glass, i.e., to A\u2081 = (9\u03c0/32)\u00b9/\u00b3 (mpl.fpg)\u00b2/\u00b3 = 0.9596(m\u00f8lfp\u2082) 2/\u00b3, where mp is the initial pellet mass,\n\n\n2/3\n\n\ns, fis the\n\n\nbatch mass fraction per glass, and pg is the glass density.\n\n\nFor X-ray diffraction (XRD) and microscopy examinations, approximately 10-g samples were\nheated at 5\u00b0C/min to temperatures of 400 to 1200\u00b0C in either a porcelain crucible (samples heated to\n<700\u00b0C) or Pt/10%Rh crucibles (samples heated to 800 to 1200\u00b0C) and air quenched. Each sample was\nweighed before and after the heat treatment. Half of the specimens were powdered and mixed with\n\n\n~5 mass% CaF2 as an internal standard for XRD. The other half of specimens, heat treated to 1100 and\n1200\u00b0C, were used to make thin sections for optical microscopy.\n\n\nXRD was performed with a Bruker\u24c7 D8 Advance diffractometer (Bruker AXS Inc., Madison,\nWI) equipped with a Cu Ka target at a power level of 40 kV and 40 mA. The instrument utilized a\nLynxEyeposition-sensitive detector with an angular range of 3\u00b0 20. The mass fractions of crystalline\nphases were obtained from XRD patterns with Jade\u00ae and RIQAS\u00ae software (MDI, Livermore, CA,\nU.S.A.).\n\n\nOptical microscopy on samples heat-treated to 800 to 1200\u00b0C was performed with Olympus\u24c7\n(Center Valley, PA, U.S.A.) SZH10 and PMG-3 microscopes. Samples of batches containing corundum\nand gibbsite, heated to 600 and 700\u00b0C, were viewed with a JEOL\u00ae (Akishima, Tokyo, Japan) 5900\nscanning electron microscope (SEM) and analyzed with an EDAX\u00ae (Mahwah, NJ, U.S.A.) lithium-drifted\nsilicon energy-dispersive X-ray spectrometer (EDS). Since the samples were not fully sintered, they were\ncast in an epoxy resin and polished to an optical quality finish before analysis.\n\n\nFinally, batch samples, ~20 to 50 mg, were analyzed with TA Instruments' (New Castle, DE,\nU.S.A.) SDT Q-600 for thermo-gravimetric analysis (TGA).\n\n\nResults\n\n\nFrom TGA, the total mass losses on melting were -0.27 g/g glass for the gibbsite batch, ~0.17 g/g\nglass for the boehmite batch, and ~0.15 g/g glass for the corundum batch, roughly corresponding to the\ndifferent content of water in the alumina sources, but were slightly smaller than those obtained from the\nstoichiometry of the batched chemicals\u2014see Table 3. As seen in Figure 1, the peak heights and positions\nin the TGA plots reveal large differences in gas-releasing reactions between batches that differ solely in\nthe alumina source. Further investigation of these reactions and associated off-gas is presently underway.\n\n\nFigure 2 displays photographs of the cylindrical pellets taken at specific temperatures and\n\n\nFigure 3 shows the plots of the normalized profile area, A/Ag, against temperature. After mild gradual\nexpansion, the batch with corundum shrank starting at ~600\u00b0C, reached a minimum profile area at\n~650\u00b0C, and then rapidly expanded to more than twice the minimum profile area at -850\u00b0C (see also\nshrinking and expansion of pellets in Figure 2). Unlike the pellet with corundum, the pellet from the batch\nwith gibbsite maintained a nearly constant profile area until it shrank at ~700\u00b0C, reaching a minimum\nprofile area at ~800\u00b0C. The pellet profile area then began to increase, becoming ~40% larger at ~950\u00b0C\nthan the minimum profile area. The pellet from the batch with boehmite responded to heating similarly to\nthe batch with gibbsite except that both the minimum and maximum profile areas occurred at -50\u00b0C\nlower temperatures. All pellets finally collapsed to a bubbly glass that spread over the alumina support\nplate.\n\n\nFigure 4 shows the undissolved quartz and corundum fractions versus temperature, in which\narctangent trend lines were fitted to the data. A notable difference in the extent of quartz dissolution can\nfirst be seen at ~400\u00b0C where the batch containing corundum has lost -20% of the initial quartz while\nbatches with gibbsite and boehmite lost <10%. The first derivatives of the arctangent trend lines, Figure 5,\nshow that the temperature of the maximum rate of quartz dissolution increased in the order corundum-\ngibbsite-boehmite; the maximum rate of dissolution also slightly increased in that order, just enough for\nthe quartz to be fully dissolved at around the same temperature in all batches.\n\n\nCorundum dissolved rapidly between 700 and 1000\u00b0C. Hence, corundum had remained inert up\nto ~700\u00b0C. In contrast with corundum, gibbsite and boehmite could not be detected by XRD (Figure 6)\neven at temperatures as low as 500\u00b0C. Boehmite was only detected at temperatures \u2264 400\u00b0C in both\ngibbsite and boehmite batches. Immediately after the boehmite disappeared at 500\u00b0C, nepheline was\nproduced and existed up to 900\u00b0C in the feed with gibbsite, and up to 800\u00b0C in the feed with boehmite.\n\n\nSodalite was produced in batches within the range of 500\u00b0C (600\u00b0C in the feed with corundum) to\n1000\u00b0C. In all three batches, hematite was present up to ~1000\u00b0C whereas spinel formed at 900\u00b0C.\n\n\nQuartz and corundum particles can be seen in SEM micrographs, Figure 7. At 600\u00b0C, the\ncorundum particles are seen in region b of Figures 7A and 7B. Grayish regions, (a in Figures 7A and B,\nand c and d in Figure 7C) are rich in Fe2O3. The round dark object, region a, Figure 7C (the gibbsite\nbatch, 700\u00b0C), is a quartz particle surrounded by a thin diffusion layer.\n\n\nOptical micrographs, Figure 8, show sections of melts from batches with corundum and gibbsite\nheated to 1100 and 1200\u00b0C. At 1100\u00b0C, the melts somewhat differ in the size and distribution of bubbles\n\n\nand appear fairly similar in homogeneity. Dark regions contain tiny crystals of spinel (magnetite). At\n1200\u00b0C, all bubbles are gone and the crystals are nearly uniformly distributed within each melt.\n\n\nDiscussion\n\n\nThe alumina source substantially influences batch behavior throughout the whole temperature\ninterval of the batch-to-glass conversion, including the early stages. Reactions between waste components\nand additives occur as soon as they are mixed in the slurry and then during drying, as the differences\nbetween as-batched and dried masses, listed in Table 3, indicate. These reactions include those between\nacids and bases (NaOH and H3BO3), formation of solid solutions of salts, and interactions of aluminum\nhydroxide and oxyhydrate with dissolved components. However, the important differences between the\nbehaviors of batches with gibbsite or boehmite and that of the batch with corundum occur during the heat\ntreatment. These differences influence the initial gas-evolving reactions (Figure 1), the kinetics of quartz\ndissolution (Figures 4 and 5), the volume expansion or foaming (Figures 2 and 3), and the formation and\nmotions of bubbles (Figure 8).\n\n\nThe initial gas-evolving reactions and the kinetics of quartz dissolution are related through the\n\n\ngeneration of glass-forming melt in a process that influences volume expansion (foaming) [10, 14, 15,\n16]. The initial reactions evolve the chemically bonded water from salts, hydroxides, oxyhydrates, and\nacids. Nearly simultaneously with these reactions, the oxyionic salts produce eutectic melts, the borates\ngenerate the first glass-forming melt, and the hydroxides and oxyhydrates become amorphous oxides,\nmainly Al2O3, but also a portion of Fe2O3 (the other portion becomes hematite). Amorphous oxides, in\nturn, dissolve in both molten salts and molten borates. The addition of Al2O3 to the borate melt\nconsiderably increases its viscosity [17].\n\n\nThe batch melting reactions are numerous and complex. Therefore, it is difficult to identify the\nchemistry of various overlapping TGA peaks seen in Fig. 1. By stoichiometry, the gases released by early\nreactions from the gibbsite-containing batch (Table 1) are H\u2082O (1.20\u00d710\u00b3), CO\u2082 (94), NO (14), and O\u2082 (9);\nthe numbers in parentheses show m\u00b3 of gas per kg of glass at 630\u00b0C. Most of these gases are released\nfrom reactions between components rather than by simple decomposition of hydroxides and salts. The\npeak at ~130\u00b0C in the corundum-containing batch can possibly be attributed to borax, a reaction product\nof NaOH and HBO3; the peak is missing in the other two batches, where NaOH had possibly reacted\nwith gibbsite and boehmite to form sodium aluminate.\n\n\nBoth melts, ionic and covalent, attack quartz particles [16]. Alkali nitrates and carbonates react\nwith quartz to form alkali silicates that fuse with the borate melt. While the borate melt dissolves quartz\ndirectly, it also protects it from molten salts. The alumino-borate melt occupies more space (there is 24\nmass% Al2O3 in the final glass) and is more viscous, hence less aggressive, than the melt without alumina\nin the corundum-containing batch. This may explain why <10% of quartz was dissolved in the batches\ncontaining gibbsite or boehmite at 400\u00b0C, whereas more than twice as much quartz was dissolved in the\nbatch containing corundum.\n\n\nAs we argue below, a lower viscosity of the glass-forming melt in the corundum-containing batch\nmay also be responsible for the earlier closure of pores, i.e., the lower temperature at which the batch\n\n\nvolume reaches a minimum. The rapid expansion that follows the minimum volume indicates that the\nglass-forming melt has become connected, i.e., the open porosity, through which evolving gases were\nescaping freely to the atmosphere, has closed. The connected melt presents a barrier to gases trapped in\nbubbles, which cannot escape because the melt viscosity is high. If the porosity becomes closed at a\ntemperature at which the batch gases are still evolving, the growing bubbles change the bubbly melt into\nprimary foam. Primary foam expands to cellular foam and eventually collapses internally into large\ncavities that burst into the atmosphere because of the increasing internal pressure and decreasing melt\nviscosity as the temperature continues to increase [10, 15].\n\n\nThe temperature at which the glass-forming melt becomes connected determines whether primary\nfoam will occur in the batch. The batch with corundum shrank to a minimum volume\u2014the point at which\nglass melt became connected\u2014at a temperature of 660\u00b0C, whereas the batch with boehmite reached the\nminimum volume at 750\u00b0C and the batch with gibbsite at 810\u00b0C (Figure 3). In borosilicate waste glasses,\nthe gas-evolving batch reactions are complete by 700 to 800\u00b0C, typically by 750\u00b0C [18-20]. The crucial\npoint is that if the melt from the batch with corundum became connected while batch gases were still\nevolving, copious primary foam was the result; whereas the melts from the batches with gibbsite and\nboehmite released the batch gases without producing primary foam.\n\n\nAs hinted above, the cause of the early closure of open porosity in the batch with corundum\nappears associated with the delayed incorporation of alumina into the glass-forming melt. Factors that\ninfluence the temperature at which the glass-forming melt becomes connected are the volume fraction of\nthe melt (the melt-solid ratio) and the viscosity [10]\u2014a tricky concept because the melt is far from\nhomogeneous at this stage. We can assume that in the batches with gibbsite and boehmite, the melt with a\nhigher alumina content, though with a somewhat lower silica content, had a higher viscosity, and thus\nmoved more slowly, allowing the pores to close at a higher temperature.\n\n\nThe other source of bubbles, the redox or fining reactions (and also the decomposition of\n\n\nsodalite-\n\n\n-see Figure 6), produce secondary foam, to which the volume expansion in batches with gibbsite\n\n\nand boehmite can solely be attributed. In the melt from the batch with corundum, secondary foam was\n\n\nproduced before primary foam could collapse. Thus, in the batch with corundum, the transition from\nprimary foam to secondary foam was continuous.\n\n\nThough expansion from primary foam in the melt can be avoided by an appropriate choice of the\nchemical form of batch components, the presence of secondary foam is inevitable. In the glass melts\nunder study, secondary foam results from oxidation-reduction reactions, mainly involving ferric oxide\n[21]. These reactions are initially suppressed by nitrates, which are strong oxidizing agents. After nitrates\nare gone above 750\u00b0C, redox reactions reach equilibrium with oxygen in bubbles. The oxidation-\nreduction equilibria depend on the temperature and the melt basicity. As temperature increases, the\nequilibria shift to the reduced state and oxygen is liberated. Silica from dissolving quartz decreases melt\nbasicity [22], thus contributing to the secondary foam development by shifting iron redox equilibria in the\ndirection of oxygen release.\n\n\nOther contributing factors are dissolution and precipitation of iron-containing crystalline phases.\nAs has been suggested by Henager et al. [15], the dissolving hematite and the precipitating spinel may\nhelp accelerate the evolution of oxygen from the redox reaction 2 Fe2O3(m) \u2192 4 FeO(m) + O2(g), where\nm and g denote the melt and gas phase, respectively. Dissolving hematite supplies Fe2O3 to the melt by\nthe reaction Fe2O3(c) \u2192 Fe2O3(m), where c denotes the crystalline phase. At the same time, FeO reacts\nwith Fe2O3, precipitating magnetite by the reaction FeO(m) + Fe2O3(m) \u2192 Fe3O4(c), which removes FeO\nfrom the melt. Both reactions, dissolving hematite and precipitating magnetite, shift the oxidation-\nreduction reaction to the right, thus promoting oxygen evolution, and consequently secondary foaming.\n\n\nThe fact that the slow-melting batch with corundum (see Table 1) generated massive primary\nfoam, whereas the faster-melting batches with gibbsite did not, indicates that primary foam is likely a\nmajor cause of slow melting of glass batches. Secondary foam is less detrimental because it can be swept\naway by convection currents, especially by forced convection (bubbling), from the cold cap bottom to the\nfree surface of the melt, where it can collapse or accumulate. The faster melting of the batch with\n\n\nboehmite as compared to that with gibbsite remains unexplained. Attributing it to the smaller content of\n\n\nwater in boehmite seems somewhat speculative.\n\n\nThe differences in amounts and sizes of bubbles between the melts at 1100\u00b0C (Figure 8) are\nprobably associated with the melt homogeneity. In a less homogeneous melt, bubbles pass through less\nviscous portions that are likely to be connected, because high-viscosity inhomogeneities surround the\ndissolving silica and alumina grains. This may account for the presence of large bubbles in the melt from\nthe batch with gibbsite. Photographs of samples taken at 1200\u00b0C indicate that melts from all three batches\nwere nearly equally homogeneous at this temperature. The streaks, barely visible in the images, were\ncaused by convection. Two sources of natural convection operate in crucible melts. Convection driven by\nsurface tension gradients results from Na2O and B2O3 volatilization in the meniscus area. These currents\nbring the melt to the surface where the ferrous-ferric ratio is altered by the diffusion of oxygen from the\nmelt to the atmosphere, resulting in a change of phase equilibria between the melt and submicron iron-\ncontaining crystals (mostly magnetite). This change is reflected in the change of color that makes the\nstreaks visible. The other source of streaks is associated with silica-rich inhomogeneities stuck to the\ncrucible bottom. These inhomogeneities, some containing residues of quartz particles, have high viscosity\nand do not contain spinel crystals, which makes them transparent. Gas bubbles nucleate around them and\ndrag the inhomogeneities upward to the melt surface. A typical example of this effect is shown in Figure\n\n\n9.\n\n\nThe results of this study support our initial hypothesis that the slow melting of the batch with\ncorundum is caused by foam generation. Nevertheless, the batch-to-glass conversion process described in\nthis section, though plausible, is hypothetical in various aspects. Hopefully, it will inspire research into\ndiffusion processes that occur in the glass-forming melt. From a technological point of view, the set of\nsimple tests used in this study may allow preselecting batches for conducting more expensive scaling-up\nexperiments.\n\n\nConclusion\n\n\nThe alumina source in otherwise identical glass batches affects the reaction sequence during\nmelting, the kinetics of quartz dissolution, and, most notably, the batch expansion. The batch with\ncorundum exhibited ample primary foam, which was absent in batches prepared with gibbsite and\nboehmite. Excessive foaming of the batch with corundum was most likely caused by an early closure of\nopen pores by the glass-forming melt that had a low viscosity because most of the alumina was still\nundissolved when the pores were closing. The lower production rate when a batch containing corundum\nwas processed in a continuous electric melter can thus be attributed to the extensive foaming caused by\nthe late incorporation of alumina into the glass-forming melt.\n\n\nAcknowledgements\n\n\nThe authors are grateful to Albert Kruger for his assistance and guidance and to Dong-Sang Kim\nfor his help and insight in weekly meetings as well as Mark Steward for his contributions. Pacific\nNorthwest National Laboratory is operated for the U.S. Department of Energy by Battelle under Contract\n\n\nDE-AC05-76RL01830.\n\n\nReferences\n\n\n1. K. S Matlack., H. Gan, M. Chaudhuri, W. Kot, W Gong, T. Bardakci, I. Pegg, and J. Innocent,\nDM100 and DM1200 melter testing with high waste loading glass formulations for Hanford high-\naluminum HLW streams, VSL-10R1690-1, Vitreous State Laboratory, Washington DC, 2010.\n\n\n2. J. D. Vienna, Nuclear waste vitrification in the United States: Recent developments and future\n\n\noptions, Int. J. Appl. Glass Sci. 1 [3] 309-321 (2010).\n\n\n3. H. H. Russell III, W. R. Ott, \u201cThe effect of batch preparation on some zinc opal glasses,\" Glass\n\n\nTechnol. 21, 237-243 (1980).\n\n\n4. F. V. Tooley, The handbook of glass manufacture, Vol. I, p. 24, Books for Industry, New York, 1974.\n5. J. D. Vienna, A. Fluegel, D. S. Kim, P. Hrma, Glass Property Data and Models for Estimating High-\nLevel Waste Glass Volume, PNNL-18501, Pacific Northwest National Laboratory, Richland,\nWashington, 2009.\n\n\n6. P. Hrma., M. J. Schweiger, C. J. Humrickhouse, J. A. Moody, R. M. Tate, N. E. TeGrotenhuis, B. M.\nArrigoni, and C. P. Rodriguez, \u201cEffect of melter-feed-makeup on vitrification process,\u201d 2009ISRSM,\nProceedings of International Symposium on Radiation Safety Management, Daejeon, Korea, 280-290,\n\n\n2009.\n\n\n7. D. A. Pierce, P. Hrma, and M. J. Schweiger, \u201cEffect of alumina source on HLW-feed melting\nprocess,\u201d Proceedings of Materials Science and Technology 112th Ann. Meeting, CD-ROM, Amer.\n\n\nCeram. Soc. and ASM Intern. 2010.\n\n\n8. P. Hrma, \u201cMelting of foaming batches: nuclear waste glass,\" Glastech. Ber. 63K, 360-369 (1990).\n\n\n9. J. Klou\u017eek, M. Arkosinov\u00e1, L. N\u011bmec, and P. Cinibusov\u00e1 \u201cThe Role of Sulphur Compounds in Glass\nMelting,\" Eur. J. Glass Sci. Technol. A 48, 176 (2007).\n\n\n10. P. Hrma, M. J. Schweiger, C. J. Humrickhouse, J. A. Moody, R. M. Tate, T. T. Rainsdon, N. E.\nTeGrotenhuis, B. M. Arrigoni, J. Marcial, C. P. Rodriguez, and B. H. Tincher, \u201cEffect of glass-batch\nmakeup on the melting process,\u201d Ceramics-Silikaty 54, 193-211 (2010).\n\n\n11. R. Conradt, P. Suwannathada, and P. Pinkhaokham, \u201cLocal temperature distribution and primary melt\nformation in a melting batch heap,\" Glastech. Ber. 67, 103 (1994).\n\n\n12. A. H. Gerrard and I. H. Smith, \u201cLaboratory techniques for studying foam formation and stability in\nglass melting,\" Glastech. Ber. 56K, 13-18 (1983).\n\n\n13. J. Kappel, R. Conradt, and H. Scholze, \u201cFoaming behavior in glass melts.\" Glastech. Ber. 60, 189-\n\n\n201 (1978).\n\n\n14. M. J. Schweiger, P. Hrma, C. J. Humrickhouse, J. Marcial, B. J. Riley, and N. E. TeGrotenhuis,\n\n\n\"Cluster formation of silica particles in glass batches during melting,\" J. Non-Cryst. Solids 356 [25-\n27] 1359-1367 (2010).\n\n\n15. S. Henager, H., P. Hrma, K.J. Swearingen, M. J. Schweiger, J. Marcial, and N. E. TeGrotenhuis,\n\"Conversion of batch to molten glass, I: Volume expansion,\u201d accepted in J. Non-Cryst. Solids.\n\n\n16. P. Hrma, K. J. Swearingen, S. H. Henager, M. J. Schweiger, J. Marcial, N. E. TeGrotenhuis,\nConversion of batch to molten glass, II: Dissolution of quartz particles, J. Non-Cryst. Solids. 357,\n820-828 (2011).\n\n\n17. P. Hrma, B. M. Arrigoni, M. J. Schweiger, \u201cViscosity of many-component glasses\u201d, J. Non-Cryst.\nSolids 355 [14-15] 891-902 (2009).\n\n\n18. J. G. Darab, E. M. Meiers, and P. A. Smith, \u201cBehavior of Simulated Hanford Slurries During\nConversion to Glass,\" Mat. Res. Soc. Proc. 556, 215-222 (1999).\n\n\n19. P.A. Smith, J.D. Vienna, and P. Hrma, \u201cThe Effects of Melting Reactions on Laboratory-Scale Waste\nVitrification,\" J. Mat. Res., 10 [8] 2137-2149 (1995)\n\n\n20. P. Hrma, J. Maty\u00e1\u0161, and D.-S. Kim, \u201cThe Chemistry and Physics of Melter Cold Cap,\u201d 9th Biennial\nInt. Conf. On Nucl. And Hazardous Waste Management, Spectrum '02, American Nuclear Society,\nCD-ROM (2002).\n\n\n21. H. D. Schreiber, C. W. Schreiber, M. W. Riethmiller, J. S. Downey, \u201cThe Effect of Temperature on\nthe Redox Constraints for the Processing of High-Level Nuclear Waste into a Glass Waste Form,\"\nMaterials Research Society Symposium Proceeding, 176, 419\u2013426 (1990).\n\n\n22. J. A. Duffy, F. G. K. Baucke, \u201cEffect of glass composition and basicity on reduction of metal ions to\n\n\nthe metallic state in melts,\u201d Phys. Chem. Glasses 38 [1] 25-26 (1997).\n\n\nFigure captions\n\n\nFigure 1. Mass-loss rates and mass loss per unit mass of glass (inset) of batches as functions of\n\n\ntemperature.\n\n\nFigure 2. Photographic images of pressed pellets during heating.\n\n\nFigure 3. Normalized profile areas of pellets versus temperature.\n\n\nFigure 4. Fractions of undissolved quartz and corundum versus temperature. The coefficients of\ndetermination (R\u00b2) and the standard errors (s) are: R\u00b2 = 0.977 and s = 0.062 for quartz in batch with\ncorundum; R\u00b2 = 0.996 and s = 0.029 for quartz in batch with gibbsite; R\u00b2 = 0.991 and s = 0.047 for quartz\n\n\nin batch with boehmite; R\u00b2 = 0.983 and s = 0.067 for corundum.\n\n\nFigure 5. Dissolution rate of quartz and corundum versus temperature.\n\n\nFigure 6. Crystalline fractions of aluminum-containing crystalline phases in batches.\n\n\nFigure 7. SEM micrographs and EDS of batches with corundum at 600\u00b0C (A) and 700\u00b0C (B), and with gibbsite at\n700\u00b0C (C).\n\n\nFigure 8. Optical images of thin sections of melts from corundum- and gibbsite-containing batches\n\n\nheated in Pt crucibles at 5\u00b0C/min.\n\n\nFigure 9. Optical micrograph of a melt heated to 1200\u00b0C at 5\u00b0C/min in a Pt crucible. The streaks trace\ncirculation flow typical for the meniscus area on the left. Bright spots at the bottom are high-silica\ninhomogeneities stuck to the Pt surface. A rising bubble dragging the melt in its wake, while stretching\nand thinning the inhomogeneities, as can be seen on the right.\n\n\nTable 1. Steady state rate of melting in a small electric melter equipped with bubbling [1]\n\n\nRate of melting\ng/m\u00b2/s\n\n\nW(a)\n\n\nAl source\n\n\nFraction\n\n\n1150\u00b0C\n\n\n1200\u00b0C\n\n\nAl2O3\n\n\n0.436\n\n\n15\n\n\n11\n\n\n16\n\n\nAl(OH)3\n\n\n0.436\n\n\n11\n\n\nAl(OH)3\nAIO(OH)\n\n\n0.450\n\n\n17\n\n\n0.450\n\n\n19\n\n\n14\n\n\n(a) W is the mass fraction of glass components from the waste\n\n\nTable 2. Composition of batch with gibbsite in g per g of glass.\n\n\nCompound\n\n\nFraction\n\n\nAl(OH)3\n\n\n0.368\n\n\nSiO2\n\n\n0.305\n\n\nH3BO3\nNaOH\n\n\n0.270\n\n\n0.099\n\n\nLi2CO3\nFe(OH)3\n\n\n0.088\n\n\n0.074\n\n\nCaO\n\n\n0.061\n\n\nNaF\n\n\n0.015\n\n\nBi(OH)3\n\n\n0.013\n\n\nFe(H2PO2)3\n\n\n0.012\n\n\n0.011\n\n\nNa2CrO4\n\n\n0.006\n\n\nZr(OH)4.H\u2082O\n\n\nNiCO3\n\n\n0.006\n\n\nPb(NO3)2\n\n\n0.006\n\n\nNa2SO4\n\n\n0.004\n\n\nZn(NO3)2*4H2O\n\n\n0.003\n\n\n0.003\n\n\nKNO3\nNaNO2\n\n\n0.003\n\n\nMg(OH)2\nNa2C2O4\n\n\n0.002\n\n\n0.001\n\n\nTotal\n\n\n1.350\n\n\nTable 3. Batch-to-glass ratios.\n\n\nBatch-to-glass mass ratio\n\n\nDifference\n\n\nDried (b)\n\n\nAs batched (a)\n\n\nAl source\n\n\nAl2O3\n\n\n1.223\n1.350\n\n\n1.150\n\n\n0.073\n\n\n1.272\n\n\nAl(OH)3\nAlO(OH)\n\n\n0.078\n\n\n1.170\n\n\n1.260\n\n\n0.090\n\n\n(a) Mix of batch chemicals as weighed.\n\n\n(b) As determined by TGA.\n\n\n0.016\n\n\n1.30\n\n\n1.25\n\n\nMass Fraction per Glass\n\n\n0.014\n\n\n1.20\n\n\n1.15\n\n\n0.012\n\n\n1.10\n\n\nRate, Mass% s-1\n\n\n0.010\n\n\n1.05\n\n\n1.00\n\n\n0.008\n\n\n0\n\n\n200\n\n\n400\n\n\n600\n\n\n800\n\n\n1000\n\n\n1200\n\n\nTemperature, 'C\n\n\n0.006\n\n\nBatch with Al2O3\n\n\n0.004\n\n\nBatch with Al(OH)3\nBatch with AIO (OH)\n\n\n0.002\n\n\n0.000\n\n\n200\n\n\n800\n\n\n0\n\n\n400\n\n\n600\n\n\n1000\n\n\n1200\n\n\nTemperature, \u00b0C\n\n\nFigure 1.\n\n\n400\u00b0C\n\n\n800\u00b0C\n\n\n900\u00b0C\n\n\n1000\u00b0C\n\n\n100\u00b0C\n\n\n650\u00b0C\n\n\nBatch with\n\n\nAl2O3\n\n\nBatch with\n\n\nAl(OH)3\n\n\nBatch with\n\n\nAIO(OH)\n\n\nFigure 2.\n\n\n3.0\n\n\n2.8\n\n\n2.6\n\n\nBatch with Al2O3\nBatch with Al(OH)3\nBatch with AIO(OH)\n\n\nNormalized Area, A/A,\n\n\n2.4\n\n\n2.2\n\n\n2.0\n\n\n1.8\n\n\n1.6\n\n\n1.4\n\n\n1.2\n\n\n400\n\n\n0\n\n\n100\n\n\n200\n\n\n300\n\n\n500\n\n\n600\n\n\n700\n\n\n800\n\n\n900\n\n\n1000\n\n\nTemperature, \u00b0C\n\n\nFigure 3.\n\n\n1.0\n\n\nA Quartz in batch with Al2O3\nQuartz in batch with Al(OH)3\n\n\n\u25c6 Quartz in batch with AIO(OH)\nO Corundum in batch with Al2O3\n\n\n0.8\n\n\nUndissolved Fraction\n\n\n0.6\n\n\n0.4\n\n\n0.2\n12\n\n\n0.0\n\n\n1200\n\n\n200\n\n\n400\n\n\n600\n\n\n800\n\n\n1000\n\n\nTemperature, \u00b0C\n\n\nFigure 4.\n\n\n0.012\n\n\nd\n\n\nQuartz in Batch with Al2O3\n\n\na\n\n\n0.010\n\n\nQuartz in Batch with Al(OH)3\nQuartz in Batch with AIO(OH)\n\n\nb\n\n\nC\n\n\nDissolution Rate, K-1\n\n\nd Corundum in Batch with Al2O3\n\n\n0.008\n\n\n0.006\n\n\n0.004\n\n\nb\n\n\na\n\n\n0.002\n\n\n0.000\n\n\n200 300\n\n\n1100 1200\n\n\n400\n\n\n500\n\n\n600 700 800\n\n\n900\n\n\n1000\n\n\nTemperature (\u00b0C)\n\n\nFigure 5.\n\n\n0.40\n\n\nBatch with Al2O3\n\n\n0.35\n\n\nCrystalline Fraction\n\n\n0.30\n\n\n0.25\n\n\n0.20\n\n\nCorundum\n\n\n\u2015 Sodalite\n\n\n0.15\n\n\n0.10\n\n\n0.05\n\n\n0.00\n\n\n400\n\n\n500\n\n\n600\n\n\n700\n\n\n800\n\n\n900\n\n\n1000\n\n\nTemperature, \u00b0C\n\n\n0.40\n\n\nBatch with Al(OH) 3\n\n\n0.35\n\n\nCrystalline Fraction\n\n\n0.30\n\n\n0.25\n\n\nBoehmite\n\n\nNepheline\n\n\n0.20\n\n\n\u27a1. Sodalite\n\n\n0.15\n\n\n0.10\n\n\n0.05\n\n\n0.00\n\n\n600\n\n\n400\n\n\n500\n\n\n700\n\n\n800\n\n\n900\n\n\n1000\n\n\nTemperature, \u00b0C\n\n\n0.40\n\n\nBatch with AIO(OH)\n\n\n0.35\n\n\nCrystalline Fraction\n\n\n0.30\n\n\nBoehmite\n\n\n0.25\n\n\n- Nepheline\n\n\n0.20\n\n\nSodalite\n\n\n0.15\n\n\n0.10\n\n\n0.05\n\n\n0.00\n\n\n400\n\n\n700\n\n\n800\n\n\n1000\n\n\n600\n\n\n500\n\n\n900\n\n\nTemperature, \u00b0C\n\n\nFigure 6.\n\n\nA\n\n\n90\n80\n70\n60\n40\n30\n20\n10\n0\n822882222\u00b0\n100\n\n\n\u25a0a\n\n\nb\n\n\n\u0441\n\n\nMass%\n\n\nPbO\n\n\n\u03b5\u03bf\u03c4\u03b9\u03bd\n\n\nCaO\n\n\nCr203\n\n\nFe203\n\n\nNa20\n\n\nMgO\n\n\nSiO2\n\n\nZrO2\n\n\nZnO\n\n\nP205\n\n\nX1,000\n\n\n10\u03bcm\n\n\nZ0kU\n\n\n12 42 AUX\n\n\nCompounds\n\n\nB\n\n\n90\n80\n70\n60\n82822322 0\n100\n30\n20\n10\n\n\n\u041f\u0430\n\n\n\u25a0b\n\n\nMass %\n\n\na\n\n\nNa20\n\n\nSiO2\n\n\nZrO2\n\n\nPbO\n\n\nFe203\n\n\nMgO\n\n\nA1203\n\n\nP205\n\n\nCr203\n\n\nZnO\n\n\nCaO\n\n\nZOKU\n\n\nX750 20\u03bcm\n\n\n12 40 AUX\n\n\nCompounds\n\n\n0\n883 2 3 2 2 \u00b0\n100\n90\n80\n40\n30\n20\n10\n70\n60\n50\n\n\nC\n\n\na\n\n\n\u25a0a\n\n\n\u043f\u044c\n\n\n\u53e3\n\n\nMass %\n\n\n\u041f\u0441\n\n\nd\n\n\n\u25a0d\n\n\n\u03b5\u03bf\u03c4\u03b9\u03bd\n\n\nP205\n\n\nZrO2\n\n\nPbO\n\n\nBi203\n\n\nCr203\n\n\nFe203\n\n\nMgO\n\n\nSiO2\n\n\nZnO\n\n\nNa20\n\n\nCaO\n\n\n20kU\n\n\n12 40 AUX\n\n\nX500 50Mm\n\n\nCompounds\n\n\nFigure 7.\n\n\nBatch with Al(OH)3\n\n\nBatch with Al\u2082O3\n\n\n1100\u00b0C\n\n\nmm\n\n\n1 mm\n\n\n1200\u00b0C\n\n\n1 mm\n\n\n1 mm\n\n\nFigure 8.\n\n\n1 mm\n\n\nFigure 9.\n"}, "expected_output": {"claims": [{"unit": "g/g glass", "value": 0.15, "evidence": ["From TGA, the total mass losses on melting were -0.27 g/g glass for the gibbsite batch, ~0.17 g/g glass for the boehmite batch, and ~0.15 g/g glass for the corundum batch, roughly corresponding to the different content of water in the alumina sources, but were slightly smaller than those obtained from the stoichiometry of the batched chemicals", "~0.15 g/g glass"]}, {"unit": "g/g glass", "value": 0.17, "evidence": ["From TGA, the total mass losses on melting were -0.27 g/g glass for the gibbsite batch, ~0.17 g/g glass for the boehmite batch, and ~0.15 g/g glass for the corundum batch, roughly corresponding to the different content of water in the alumina sources, but were slightly smaller than those obtained from the stoichiometry of the batched chemicals", "~0.17 g/g\nglass"]}, {"unit": "g/g glass", "value": 0.27, "evidence": ["From TGA, the total mass losses on melting were -0.27 g/g glass for the gibbsite batch, ~0.17 g/g glass for the boehmite batch, and ~0.15 g/g glass for the corundum batch, roughly corresponding to the different content of water in the alumina sources, but were slightly smaller than those obtained from the stoichiometry of the batched chemicals", "-0.27 g/g glass"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_11986fccc0f1302c"}}
{"id": "4a48674514a874c59a8cb269", "input": {"query": "What are the tin loss rates to sludge in electrolytic tinning processes?", "source_url": "https://tenova.com/sites/default/files/2021-12/2004-Tinplate%20conference-Paper%20-%20Astengo%20and%20Ferrari.pdf", "document_text": "Centro Sviluppo\n\n\nTechint Technologies\n\n\nTHE Materiali S.p.A.\n\n\ntin lost vs. dissolved tin). The electrolyte is satu-\nrated with pure oxygen by means of a system for\noxygen dissolution designed and realized to this\npurpose.\n\n\nTECHINT INSOLUBLE\n\n\nANODES\n\n\nTECHNOLOGY WITH AN\n\n\nIM-\n\n\nPROVED TIN\n\n\nDISSOLUTION\n\n\nPROCESS.\n\n\nInteresting results were achieved on the PSA elec-\ntrolyte. After dedicated tests looking for the best\nprocess parameters, finally the sludge produced by\nthe process was less than 4 % as tin lost vs. dis-\nsolved tin. The tin dissolution rate can be easily\ncontrolled by setting the oxygen flow properly.\n\n\nG. Astengo); Ferrari(2)\n1. Techint Technologies, Torre Shipping, Via De Marini\n53 16149 Genova, Italy\n2. Centro Sviluppo Materiali, Via di Castel Romano 100-\n00128 Roma, Italy.\n\n\nFurther trials carried out on a pilot dissolution reac-\ntor at CSM laboratories showed that the process is\ninsensitive to the type of tinplating electrolyte. In\nfact, there were no significant differences between\nthe amounts of sludge generated with phenol-\nsulphonic acid (PSA) either with addition of DI-\nPHONE or ENSA, and methansulphonic acid\n(MSA).\n\n\nABSTRACT\n\n\nThe application of Techint insoluble anode technol-\nogy in new or existing tinning lines brings consider-\nable benefits in tinplating such as:\n\n\nminimizing the amount of sludge and hence\nthe loss of tin;\n\n\nThe pilot plant was able to produce more than 4\nkg/h of dissolved tin.\n\n\nreduction of manpower for anode handling\nimproved coating quality\n\n\nbetter process control\n\n\nThe trials were carried out at different dissolution\nrates and with different electrolytes, monitoring the\ntotal amount of generated sludge, the consumptions\nof the additives and the quality of tinplate by the\nHull cell test.\n\n\nelimination of phenol vapors inside the\nbuilding.\n\n\n-\n\n\nThe paper describes into details such improvements\nas well as the very satisfactory results reached in\ntinplate produced with insoluble anodes. In fact,\nwith conventional electroplating technology the\nhomogeneity of tin coating thickness decreases as\nthe coating weight decreases, due to the particular\ngeometry of the tin anodes and due to non-uniform\nconsumption of tin bars.\n\n\nThese successful tests demonstrate that Techint in-\nsoluble anodes technology is applicable either in\nnew or existing tinning lines whatever is the type of\nelectrolyte used.\n\n\nINTRODUCTION\n\n\nThe critical point of the process is the high produc-\ntion of sludge in the present tin dissolution systems\nused in the electrotinning lines with insoluble an-\nodes. The tin lost in sludge is higher than 10% of\nthe dissolved tin. This is the reason why the tinplate\nprocess with insoluble anodes is not yet used\nworldwide.\n\n\nTechint has recently developed and realised two\nmodern and fast tinning lines. The Basic design pa-\nrameters are summarized in Table 1.\n\n\nThe ETL/TFSL are capable of tin coating from 1.12\ng/m\u00b2 to 11.2 g/m\u00b2. All combinations of differential\ncoatings from 1.12 / 2.24 g/m\u00b2 to 8.4 / 11.2 g/m\u00b2 are\nproducible.\n\n\nBecause of that, Techint Technologies, a division of\nTECHINT Group, and its connected Research Cen-\nter, Centro Sviluppo Materiali (CSM) have devel-\noped a new low-sludge tin dissolution process suc-\ncessfully proven through extensive tests on Siderar's\ntinning line (Argentina).\n\n\nThe most common product of the equal coating per\nside is 2.8 / 2.8 g/m\u00b2. The most common differential\ncoated strip is 2.8/8.4 g/m\u00b2. Lower tin coatings\n<1.12 g/m\u00b2 are also feasible.\n\n\nThe process is based on the oxidation of metallic tin\ngranules by the tinning electrolyte flowing in a dis-\nsolution reactor, which has well defined process and\ntechnological settings to reduce sludge under 4% (as\n\n\nThe process is based on the use of soluble anodes\nand either PSA or MSA baths. Anyway, Techint and\nCSM are making considerable efforts to develop the\ninsoluble-anode technology provided with an im-\n\n\nPag. 1/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\n\n\nTechint Technologies\n\n\nT Materiali S.p.A.\n\n\nproved tin dissolution process able to produce very\nlow amount of sludge.\n\n\nincrease of tin concentration in the plating solution\nis unavoidable with the use of tin anodes, and dilu-\ntion of the solution generates overflow and dis-\ncharge, with loss of expensive material and possible\nwater pollution, unless adequately treated.\n\n\nBASIC DESIGN PARAMETERS\n\n\nElectrolyte: either\n\n\nPhenol sulphonic acid\n(PSA)\nMethane\nAcid (MSA)\n\n\nThe conventional equipment used worldwide for the\nremoval of excess tin is the insoluble anode and\nmany customers today have one half cell equipped\nwith insoluble anodes, but the control of the solu-\ntion is very difficult and concerns about a rapid de-\ncrease of tin and an increase of free acid in the plat-\ning solution make the use of insoluble anodes in this\nway less practical. Other disadvantages with the\nFerrostan process are the fumes exiting the plating\ntanks, the labour requirements for handling the tin\nanodes and the low productivity.\n\n\nSulphonic\n\n\nor\n\n\nProduction Capacity\nTinplate:\n\n\n200,000 t/y and over\ndepending on the mix\n\n\nTFS:\n\n\n50,000 t/y\n250,000 t/y\n\n\nTotal:\n\n\nLine Reference Length\n\n\n140 m\n\n\nMaterial Specification\n\n\nAdditionally, market demand is towards tinplate\nwith thinner coatings; indeed for some uses tin coat-\nings down to 0.2-0.4 g/m\u00b2 are required, causing\nproduction problems.\n\n\n0.10 to 0.60 mm\n\n\nThickness:\n\n\n508 to 1250 mm\n25,000 kg maximum\nup to 2100 mm\n\n\nWidth:\n\n\nCoil Weight:\n\n\nCoil OD:\n\n\nWith conventional electroplating technology the\nhomogeneity of tin coating thickness decreases as\nthe coating weight decreases, due to the particular\ngeometry of the tin anodes, which do not present a\ncontinuous surface. In fact, each anode is formed by\na series of vertical bars drawn against each other so\nas to leave only a minimum space between the bars,\nwhich may produce a lower tin thickness.\n\n\nLine Speeds\n\n\nEntry & Exit Sections:\nProcess Section:\n\n\n700 m/min maximum\n\n\n550 m/min maximum\n450 m/min maximum\nfor TFS\n\n\n30 m/min\n\n\nThreading Speed:\n\n\nExit Section Shear Speed: 120 m/min maximum\n\n\nCoating Thickness Range (per side)\n\n\n11.2 g/m\u00b2 maximum\n1.12 g/m\u00b2 minimum\n100 mg/m\u00b2 maximum\n30 mg/m\u00b2 minimum\n25 mg/m\u00b2 maximum\n10 mg/m\u00b2 minimum\n\n\nTin Coating:\n\n\nAnother cause of irregular tin coating derives from\nnon-uniform consumption of tin bars, which in turn\ngives rise to preferential current distribution. For\nthicker coatings such situations are alleviated since\nmore cells are employed in the sequence of electro-\nplating steps, thus allowing the tin coating to grow\nmore uniformly. In thin coatings this will not be al-\nlowed anymore; for instance, in tinplate for the fab-\nrication of two-piece DWI cans where, owing to\nwall ironing, the tin coating is reduced and could\ndisappear in those areas where the coating thickness\nis not homogeneous.\n\n\nChrome Coating as Metal\n\n\nChrome as Oxide:\n\n\nOil Type\n\n\nDOS and ATBC\n2.5-20 mg/m\u00b2/side\n\n\nTable 1 - Design parameters of ETL/TFS line\n\n\nThe Ferrostan process, based on the use of soluble\ntin anodes in Phenolsulphonate solution, is well\nconsolidated in electro tinning lines. The use of\nsoluble anodes is advantageous because the tin\nplated-out on the strip can be automatically pro-\nduced by the dissolution of tin from the anodes; but\nthere are disadvantages.\n\n\nOne solution to these problems is to equip an entire\ntinning line with insoluble anodes, the anode then\nhas a continuous surface and very thin tin coating\nlayers can be produced with high thickness homo-\ngeneity.\n\n\nThe main advantages of tin-plating with insoluble\nanodes are summarized in table 2.\n\n\nThe most important one is the necessity to drain off\nthe plating solution because of the different electro-\nchemical efficiency in plating and dissolution. An\n\n\nPag. 2/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\nTH Materiali S.p.A.\n\n\nTechint Technologies\n\n\nTECHINT\n\n\nNevertheless, the use of non-soluble anodes still re-\nquires a tin dissolution reactor, connected to the\nelectro tinning line to replenish the amount of tin\ndeposited on the strip.\n\n\nMECHANISM OF DISSOLUTION DEPOSI-\n\n\nTION PROCESS\n\n\nThe process is based on the oxidation of metallic tin\ngranules in the tinning electrolyte saturated with\npure oxygen.\n\n\nConsiderable research was done to dissolve tin\n(chemically or electrochemically) in the plating so-\nlution [1-10]. An industrial process was set up many\nyears ago and a few plants, mainly in Japan, are\nworking with insoluble anodes.\n\n\nThe main reactions involved in the metallic tin dis-\nsolution process are:\n\n\n2Sn + O2 + 4H \u2192 2Sn\u00b2+ + 2H\u2082O\n\n\n1.\n\n\nSn + O2 + 4H* \u2192 Sn4+ + 2H\u2082O\n\n\nPARAMETER\n\n\nBENEFIT\n\n\n2.\n\n\n2Sn\u00b2 + O2 + 4H* \u2192 2Sn** + 2H\u2082O\n\n\n4+\n\n\nConstant tin cov-\n\n\nLess tin consumed\n\n\n3.\n\n\nering on strip\nBetter edges\n\n\nSn\u00b2+2PSA \u2192 Sn(PSA)2\n\n\n4.\n\n\nBetter strip quality, particularly\nimportant for thin coatings\nReduced labour costs, higher\nproductivity and flexibility,\nsafer and better working envi-\nronment\n\n\nSn+ Sn 4+ 2Sn\u00b2+\n\n\n5.\n\n\nNo anode\n\n\nhandling\n\n\nOxygen is necessary to enhance the reaction rate of\nthe oxidation of tin by the acidity of the bath formed\nat the insoluble anode.\n\n\nNo anode melting\nplant\nCovering on tanks\nElectrolyte always\nunder control\nAnodes closer to\nstrip\n\n\nReduced labour\n\n\nThe electrochemical reactions occurring at the elec-\ntroplating site with insoluble anodes are:\n\n\nLess fumes\n\n\nLower electrolyte discharge,\nconsumption and pollution\nReduced electricity consump-\ntion\n\n\n2Sn24e2Sn (cathode)\n\n\n6.\n\n\n2H\u2082O \u2192 O2 + 4e\u00af + 4H* (anode)\n\n\n7.\n\n\nTable 2- Advantages of insoluble-anode process\n\n\nThe four moles of hydrogen ions formed at the an-\node (reaction 7) restore the as many moles of hy-\ndrogen ions consumed in the dissolution reactor to\ndissolve two moles of metallic tin (reaction 1),\nwhich, in turn, restore the two moles of tin depos-\nited on the strip (reaction 6). Thus the mass balance\nis assured.\n\n\nThe process of chemical tin dissolution, using oxy-\ngen to accelerate the oxidation reaction of metallic\ntin to ionic tin, cannot avoid the formation of a\nSn(IV)-based sludge.\n\n\nThe critical point of the process is the high produc-\ntion of sludge in the present tin dissolution systems\nused in the electrotinning lines with insoluble an-\nodes.\n\n\nPILOT LABORATORY LINE\n\n\nThe tin lost in sludge is higher than 10% of the dis-\nsolved tin. This is the reason why the tinplate proc-\ness with insoluble anodes is not yet used worldwide.\nTechint, in cooperation with CSM, developed an in-\nnovative process for dissolving tin.\n\n\nAfter a first set of laboratory investigations, a 30 1\ncapacity pilot reactor for tin dissolution was initially\nrealised at CSM laboratories (figure 1).\n\n\nThe plant comprises a tin dissolution section and a\ntin electroplating section. The dissolution section\ncontains the chemical reactor, the pressurizing\npump, the oxygen feeder and the tank for the prepa-\nration and storage of the solution.\n\n\nThis process is able to minimize the amount of\nsludge and loss of Sn and has been successfully\nproven through extensive tests on Siderar's tinning\nline.\n\n\nPag. 3/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\nTE Materiali S.p.A.\n\n\nTechint Technologies\n\n\nTECHINT\n\n\nElectrolytic solution flow rate\n\n\nTin particle size\n\n\nAn additional test was performed at high tempera-\nture (~60\u00b0C) to evaluate the influence of this pa-\nrameter on the dissolution kinetics and sludge for-\nmation.\n\n\nThe effect of the above parameters on the tin disso-\nlution rate and on the sludge production allowed for\nthe definition of the criteria for designing an indus-\ntrial tin dissolution plant able to generate a low\namount of sludge.\n\n\nPILOT PLANT CONNECTED TO AN INDUS-\nTRIAL LINE\n\n\nFigure 1 \u2013 CSM pilot plant\n\n\n-\n\n\nIn a further phase of the work an industrial-scale re-\nactor was realised to feed one insoluble anodes elec-\ntrolytic cell of the electro tinning line of Siderar\nWorks in S. Nicolas, Argentina (figure 2).\n\n\nThe PSA electrolyte flows through the reactor by\nmeans of a pump operating at a pressure of up to 8\nbar. Oxygen is fed into the depleted electrolyte by a\nspecial feeder designed to minimize the size of oxy-\ngen bubbles and promote their immediate dissolu-\ntion. The electroplating section has a vertical elec-\ntrolytic cell with flat parallel electrodes, a pump for\nsolution movement and a recirculation tank.\n\n\nSince the tin dissolution rate is proportional to oxy-\ngen activity, oxygen-saturated electrolyte at pres-\nsures higher than 1 bar, are used. A nozzle for feed-\ning oxygen into the solution was specifically devel-\noped for this purpose.\n\n\nThe dissolution of metallic tin achieved through\noxidation in acidic environment by dissolved oxy-\ngen involves two main technological problems: i) to\nget the maximum solution of gaseous oxygen, pos-\nsibly up to its saturation, even when working under\npressure, ii) to maximize the mass transport coeffi-\ncient using a fluidized bed reactor, which also\nminimizes the non-reactive volume of the tin\ncharge.\n\n\nA\n\n\nPreliminary pilot tests\n\n\nA first series of tests were carried out by varying the\nfollowing parameters:\n\n\nOxygen flow rate\n\n\nHeight of the fluidized tin bed\n\n\nFigure 2 - Tin Dissolving Pilot Plant at Siderar\n\n\nSurface area of metallic tin\n\n\nPag. 4/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\n\n\nTechint Technologies\n\n\nT Materiali S.p.A.\n\n\nTechint Technologies decided to design the pre-\nwetting tank so that it could be used either as a\nnormal pre-wetting tank or as an insoluble anode\nplating tank connected to a dedicated tin dissolution\nplant for replenishing the plated-out tin.\n\n\ndepending on the tin dissolution rate set in the plant,\nand the call for charging is performed by the plant\nautomatic system when the quantity inside the reac-\ntor has reached a minimum value.\n\n\nThe reactor is a vertical cylindrical vessel, divided\nin two parts: the upper part has a large diameter to\navoid small tin particles to escape from the reactor.\nThe total volume is about 1 m\u00b3.\n\n\nThe tin dissolution plant was erected, connected to\nthe revamped line and began operation at the begin-\nning of 2001. By the end of the same year the re-\nsearch activity connected to this first industrial\ncampaign was completed.\n\n\nThe plating solution is fed from the bottom into the\nreactor through a distributor that supports the metal-\nlic tin particles and distributes the solution. Before\nentering the reactor the solution is enriched in dis-\nsolved oxygen.\n\n\nThe schematic flow diagram of the plant is here be-\nlow shown in figure 3.\n\n\nSuitable heat exchangers are installed in the circuits\nin order to maintain the plating solution at a con-\nstant temperature and, because of the drag-out from\nthe plating tanks in order to keep the level constant\nin the two recirculating tanks, a refilling of solution\nfrom the normal line is available. The plating tank is\nequipped with insoluble anodes located in the tank\nas shown in figure 4.\n\n\nSKETCH OF THE INDUSTRIAL PILOT PLANT at SIDERAR WORK\n\n\nFLOOR LEVEL\n\n\n2nd Step\n1st Plating Tank\n\n\n1st Step\nPre-Wetting\n\n\nTK 1-7 MC\n\n\nREACTOR\nCIRCUIT\n\n\nFILTER\n\n\nBASEMENT\n\n\nPLATING\nCIRCUIT\n\n\nTK 2-3 MC\n\n\nFILTER\nCIRCUIT\n\n\nFigure 3 \u2013 Scheme of the industrial-scale reactor\n\n\nThe plant consists of a reactor into which irregular\nsized tin pellets of about 2 mm diameter, are\ncharged into the metallic tin bed from the top of the\nreactor, four recirculating circuits and two recircu-\nlating tanks, one of which was already existing in\nthe old line.\n\n\nThe first circuit feeds the plating tank from which\nthe low tin ion content solution is returned because\nof the deposition of tin on the running strip; and a\nsecond circuit connects the two recirculation tanks.\nA third, and most import circuit, feeds the tin disso-\nlution reactor. This circuit works at high pressure\n(5-6 bar) in order to provide high solubility of oxy-\ngen in the solution. The fourth circuit removes the\nsludge from the solution. The tin dissolution plant\nhas a design capacity of 30 kg/h of dissolved tin.\nThe charging system was designed with proper\nautomatic valves to maintain the recirculation inside\nthe reactor during the charging operation. The tin is\ncharged in the reactor once every two to three hours\n\n\nFigure 4 - Insoluble anodes in industrial plant\n\n\nThe Iridium Mixed Metal Oxide covered titanium\nanodes are connected to two rectifiers each having a\ncapacity of 4,000 A. The replenishing rate of the\nplant is controlled by the flow rate of the oxygen.\nThe plant can work in three modes:\n\n\nPag. 5/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\n\n\nTechint Technologies\n\n\nTHE Materiali S.p.A.\n\n\nThe results were very good (see figure 5); the plant\nwas able to reach a dissolution rate higher than 40\nkg/h without loss in efficiency and without higher\ngeneration of sludge.\n\n\nat constant oxygen flow rate (manual)\nat constant tin dissolution rate (auto 1 mode)\n\n\nat a rate proportional to the current given by the\nrectifiers (auto 2 mode).\n\n\nIn the Autumn campaign the sludge percentage was\nvery low with a maximum rate of 4%, expressed as\ntin lost in sludge vs. dissolved tin; the only excep-\ntion being in the fifth test when the plant was set\noutside the standard conditions for a final check.\n\n\nThis last mode permits working at constant tin con-\ncentration in the solution because the replenishing\nrate is calculated taking into consideration the tin\nplated out at any given moment.\n\n\nA simulation model was developed to providing a\ncontinuous on-line calculated value of actual disso-\nlution rate, based on the following main process pa-\nrameters: the tin charge, pressure, temperature, in-\njected oxygen and recirculating solution flow rate.\n\n\n% OF TIN LOST IN SLUDGE\nRESULTS FROM PLANT OPERATION\n\n\nTechint Results\n\n\n15%\n14%\n\n\n12-13 %: Actual average from competitors\n\n\n13%\n12%\n\n\n10%: Actual minimum from competitors\n\n\n11%\n\n\n10%\n\n\n9%\n8%\n\n\nINDUSTRIAL TIN DISSOLUTION\nAND DEPOSITION TESTS\n\n\nFIRST\n\n\n5%: Declared minimum from competitors\n\n\n7%\n\n\n6%\n\n\n5%\n\n\nI\n\n\n4%\n\n\n3%\n2%\n\n\nContinuous weekly tests with simultaneous dissolu-\ntion and tin-plating were performed. No problems\nwere encountered with the line operation or product\nquality and during the test operation the plant\nmainly worked at constant tin dissolution rate for\ncalibration purposes.\n\n\n1%\n\n\n0%\n\n\nS3\n\n\nA3 A4\nAUTUMN CAMPAIGN\n\n\nS1\n\n\nS2\n\n\nS4\nSUMMER CAMPAIGN\n\n\nS5\n\n\nS6\n\n\nA1\n\n\nA2\n\n\nA5\n\n\nA6\n\n\nFigure 5 \u2013 Tin losses during summer and autumn\ncampaigns\n\n\nThe dissolution rate was found to always depend on\nthe oxygen flow rate while working in conditions of\nfull solubility of oxygen according to Henry's law.\nThis was demonstrated at reactor pressures of 2, 3\nand 4 bar.\n\n\nDuring the autumn tests, tinplate with a coating\nweight of 2 g/m\u00b2 for each side was produced us-\ning only the first electrolytic cell equipped with in-\nsoluble anodes and fed with the electrolyte coming\nfrom the tin dissolution reactor.\n\n\nAt each test the quantity of generated sludge was\nweighed and correlated with the dissolved tin. The\nsludge can be generated both in the dissolution and\nplating circuits. The weighed amount is the total\nfrom the two circuits.\n\n\nThen, tinplate with the same tin weight was pro-\nduced, for comparison, using only the second elec-\ntrolytic cell, equipped with soluble anodes and fed\nwith classic electrolyte coming from the storage\ntanks.\n\n\nDepending on the operating conditions (pressure,\ninjection of oxygen, tin charge, solution flow rates,\nand so on) different quantities of tin lost in the\nsludge were measured.\n\n\nNo difference was seen as far the visual appearance\nof the two products. SEM (Scanning Electron Mi-\ncroscope) and GDOES (Glow Discharge Optical\nEmission Spectroscopy) investigations (figures 6-9)\nconfirmed that the tinplate produced with insoluble\nanodes is quite equal to that produced with soluble\nanodes.\n\n\nTwo main test campaigns were performed (in sum-\nmer and in autumn). In the first one, performed dur-\ning summer 2001, the behavior of the plant was\ntested in different operating conditions by varying\none parameter at a time and keeping the others con-\n\n\nstant.\n\n\nThen the plant was set up in the best operating con-\ndition working out from the results of the first cam-\npaign. A second industrial campaign was carried out\nin the autumn 2001.\n\n\nPag. 6/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\nT Materiali S.p.A.\n\n\nTechint Technologies\n\n\nTECHINT\n\n\nTinplate produced with insoluble anodes\n\n\n100%\n\n\n90%\n\n\nConcentration (weight %)\n\n\n80%\n\n\n70%\n\n\nIron\nTin\n\n\n60%\n\n\n50%\n\n\nChromium\nOxygen\n\n\n40%\n\n\n30%\n\n\n20%\n\n\n10%\n\n\n0%\n\n\n0 50 100 150 200 250 300 350 400 450 500\nThickness (nm)\n\n\nMAR\n\n\nFigure 9 - Element profile, tinplate insoluble anodes\n\n\n100\u03bcm\nX100 21 mm\n\n\n00006 20KV\n\n\nFigure 6 \u2013 Tinplate produced with soluble anodes\n\n\nPILOT TESTS WITH DIFFERENT ELEC-\nTROLYTES\n\n\n1B\n\n\nAll the activities and the results described above\nwere performed with PSA-based electrolyte\ncontaining Diphone V as additive.\n\n\nTo test the low-sludge tin-dissolution process also\nwith baths having different additives (Diphone VI,\nENSA) and with new ecological baths based on\nMSA system, a new series of tests was carried out\non the pilot dissolution plant at CSM laboratories.\nThe trials were carried out at different dissolution\nrates and with different electrolytes, monitoring the\ntotal amount of generated sludge, the consumptions\nof the additives and the quality of tinplate by the\nHull cell test.\n\n\n100\u03bcm\n\n\n00006 20KV\n\n\nX100 21mm\n\n\nFigure 7 - Tinplate produced with insoluble anodes\n\n\nFour types of electrolytes were used in the tests:\n\n\nPSA + Diphone 5\nPSA + Diphone 6\n\n\n-\n\n\nPSA ENSA 6\n\n\nTinplate produced with soluble anodes\n\n\nRonastan system (MSA + proprietary addi-\ntives)\n\n\n100%\n90%\n\n\nConcentration (weight %)\n\n\n80%\n70%\n60%\n\n\nThe electrolyte volume was 500 1. The tests were\nmade with a flow rate of 5 m\u00b3/h and a pressure of 5\n\n\nIron\nTin\n\n\n50%\n\n\nata.\n\n\nChromium\nOxygen\n\n\n40%\n\n\nAlso electrolytes containing up to 15 g/l Fe were\ntested to check the effect of such ion on the loss of\ntin in sludge. Fe ion is generally present in tin\nelectrolyte owing to the dissolution of steel strip in\nthe first half cell (pre-dip).\n\n\n30%\n\n\n20%\n10%-\n\n\n++\n\n\n0%\n\n\n400\n\n\n200\n\n\n500\n\n\n300\nThickness (nm)\n\n\n0\n\n\n100\n\n\nFigure 8 \u2013 Element profile, tinplate soluble anodes\n\n\nThe results were really good with all the electro-\nlytes. Very low percentages of tin loss were de-\n\n\nPag. 7/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\n\n\nTechint Technologies\n\n\nTH Materiali S.p.A.\n\n\nrate, according to the current applied in the two\nelectrolytic cells equipped with insoluble anodes.\n\n\ntected, with a maximum rate of 4,5 %, no matter the\nelectrolyte used (figure 6).\n\n\nThus, the amount of tin dissolved at any time is\nequal to that of tin deposited on the strip, plus the\namount of tin to compensate the drag-out.\n\n\nTin lost in sludge\n\n\n5,0\n\n\n4,5\n\n\n\u05d5\u05d5\u05d5\u05d5\u05d9\nPSA+Diphone 6\n\n\nTin in sludge vs. tin dissolved (%)\n\n\n0,0\n\n\nPSA+Diphone 5\n\n\nPSA+ENSA 6\n\n\nMSA+SG+TP\n\n\nPSA+Diphone 5\n\n\nFigure 6 \u2013 Tin loss with various electrolytes\n\n\nAs low percentage of tin loss were found even with\nthe same electrolytes containing up to 15 g/l iron\n(figure 7).\n\n\nFigure 8\npaign.\n\n\nInsoluble Anodes in 2004 industrial cam-\n\n\n-\n\n\nTwo long-term running tests were carried out. The\namount of sludge was measured at the end of each\ntest and correlated to the dissolved tin. The total\namount of tin lost resulted less than 4 %.\n\n\nTin lost in sludge\n\n\n5,0\n4,5\n4,0\n3,5\n3,0\nCNN w w A A S\n\n\nTin in sludge vs. tin dissolved (%)\n\n\n13,5gl Fe\n\n\nThe insoluble anodes were connected to four rectifi-\ners for a total capacity of 22,000 A. Figure 9 shows\nthe amount of current provided, during normal op-\nerating conditions of the line, with insoluble anodes,\nwhich is proportional to the tin plated with insoluble\nanodes. In general more than 30 % is given with in-\nsoluble anodes, up to a maximum of 45 % in case of\nlow coating (2.2 g/m\u00b2).\n\n\n15,1 gl Fe\n\n\n12,7 gl Fe\n\n\n4,5g/l Fe\n\n\n0,5\n\n\n0,0\n\n\nPSA+Diphone 6\n\n\nPSA+ENSA 6\n\n\nPSA+ENSA 6\n\n\nMSA+SG+TP\n\n\nTwo commercial coils were also produced only with\ninsoluble anodes.\n\n\nFigure 7 \u2013 Tin loss with electrolytes containing Fe\n\n\nFigure 10 shows the progressive amount of tin dis-\nsolved and plated out during one test.\n\n\nSECOND INDUSTRIAL TIN DISSOLUTION\nAND DEPOSITION TESTS\n\n\nThe very good results obtained in the previous in-\ndustrial campaign with only the pre-dip cell\nequipped with insoluble anodes, were confirmed\neven in this second campaign with one cell and a\nhalf equipped with insoluble anodes.\n\n\nIn May and June 2004 a second industrial campaign\nwas carried out at Siderar electrotinning line. Both\nthe pre-dip half cell and the second electroplating\ncell were equipped with insoluble anodes.\n\n\nAn automatic procedure was adopted to set the oxy-\ngen flow rate and, consequently, the tin dissolution\n\n\nPag. 8/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\n\n\nTechint Technologies\n\n\nTH Materiali S.p.A.\n\n\nTECHINT\n\n\n95% Operating time with\nINSOLUBLE ANODES\n525.5 tons with INS. ANODES\n\n\nTin plated with insoluble anodes / Total Tin deposit\non 18/5/04 (8\u00b0 day)\n\n\n50%\n\n\n45%\n\n\n40%\n\n\n35%\n\n\n30%\n\n\n25%\n\n\n20%\n\n\n15%\n\n\n10%\n\n\n5%\n\n\nTin coating g/m\u00b2\n\n\n0%\n\n\n* 3 3 5 3 3 3 3 3 3 3 3\n\n\n23.28\n\n\nTime [h.min]\n\n\nFig. 9 - Amount of current given with Insoluble Anodes during one typical day\n\n\nCONCLUSIONS\n\n\nLast trials carried out on a pilot laboratory plant at\nCSM showed that the process is insensitive to the\ntype of tinplating electrolyte. In fact, there were no\nsignificant differences between the amounts of\nsludge generated with phenol-sulphonic acid (PSA)\neither with addition of DIPHONE or ENSA, and\nmethanesulphonic acid (MSA). Moreover also the\npresence of Fe ions in solutions has been checked.\nThe results showed no influence on the amount of\nproduced sludge.\n\n\nTIN DISSOLVED AND TIN PLATED OUT\n\n\n3500.0\n\n\n3000.0\n\n\nDISS.\nPLATED\n\n\n2500.0\n\n\n2000.0\n\n\nkg\n\n\n1500.0\n\n\n1000.0\n\n\n500.0\n\n\n0.0\n\n\n11/05/2004 12/05/2004 13/05/2004 14/05/2004 15/05/2004 16/05/2004 17/05/2004 18/05/2004\n\n\nThe second industrial campaign confirmed that the\nTechint-CSM tin dissolution process is able to pro-\nduce a tin loss in sludge as great as that of the clas-\nsic electrotinning process based on the use of solu-\nble anodes.\n\n\nDay\n\n\nFigure 10 \u2013 Tin dissolved and plated out during a test\n\n\nThanks to such results, the tinplate process with in-\nsoluble anodes is ready for full industrialisation,\nbringing all the advantages connected to the use of\ninsoluble anodes, which were not exploited so far\nfor the absence of a low sludge dissolution system.\nA further step in the direction of a full insoluble an-\nodes line is scheduled for the tinning line at Siderar\nwork, where a third plating tank is going to be\nequipped with insoluble anodes.\n\n\nElectrolytic El tinplate was produced by using only\nthe two cells with insoluble anodes. No visual dif-\nference in the aspect was seen in comparison to El\ntinplate produced with soluble anodes. Samples of\nsuch tinplates are under investigation by SEM and\nGDOES techniques.\n\n\nPag. 9/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n\n\nCentro Sviluppo\n\n\nTechint Technologies\n\n\nTH Materiali S.p.A.\n\n\nTECHINT\n\n\nREFERENCES\n\n\n1. J. Lampereur et al.; \u201cElectrotinning at high cur-\nrent densities\"; Proceedings 4th International\nTinplate Conference, (1988)\n2. B. Paramanathan, E. N. Soepenberg; \"A pilot\nplant for steel strip coating by radial jet elec-\ntrolysis at high current densities\", ECSC Final\nReport EUR 17817 EN, (1997)\n3. B. Colinel; \u201cElectrolyte and apparatus for elec-\ntroplating at high current density\"; Proceedings\n6th International Tinplate Conference, (1996)\n\n\nY. Suganuma et al.; Method for electrolytic tin\nplating of steel plate\"; U.S. Patent No.\n5194141, (1993)\n\n\n4.\n\n\nB. Paramanathan, \u201cMethod and apparatus for\nthe electrolytic coating of one side of a moving\nmetal strip\"; EP No. 0 364 013 B1, (1994)\n6. M. Kitayama et al.; \u00abProcess for electro-tin\nplating \u00bb ; U.S. Patent No. 4181580, (1980)\nH. Bunk et al.; \"Method of electrolytic tinning\nusing an insoluble anode\"; EP No. 0 268 823\nB1, (1987)\n\n\n5.\n\n\n7.\n\n\nD. L. De Respiris et al.; \u201cProcess for replenish-\ning metals in aqueous electrolyte solutions\";\nU.S. Patent No. 5082538, (1992)\nT. Kikuchi et al.; \"Tinning apparatus and tin-\nning method\"; JP No. 8-20889, (1996)\n\n\n8.\n\n\n9.\n\n\n10. D. Thomson; \u201cElectrolytic tin plating method\";\nU.S. Patent No. 5312539, (1994).\n\n\nPag. 10/10\n\n\nFile Name: Paper_rev2.docPaper_rev2.doc\n"}, "expected_output": {"claims": [{"unit": "% of dissolved tin", "value": 10, "evidence": ["The tin lost in sludge is higher than 10% of the dissolved tin. This is the reason why the tinplate process with insoluble anodes is not yet used worldwide.", "10%: Actual minimum from competitors"]}, {"unit": "% of dissolved tin", "value": 12, "evidence": ["The tin lost in sludge is higher than 10% of the dissolved tin. This is the reason why the tinplate process with insoluble anodes is not yet used worldwide.", "10%: Actual minimum from competitors"]}, {"unit": "% of dissolved tin", "value": 13, "evidence": ["The tin lost in sludge is higher than 10% of the dissolved tin. This is the reason why the tinplate process with insoluble anodes is not yet used worldwide.", "10%: Actual minimum from competitors"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_59489338f88848d4"}}
{"id": "f2a8fe424609c612ae561e1d", "input": {"query": "What are the specific temperature parameters for T73 and T76 tempers for 7xxx aluminum alloys? I need the two-stage aging process details including temperatures in Celsius. Extract for alloys like 7075, 7050, 7150.", "source_url": "https://materialsdata.nist.gov/bitstream/handle/11115/192/Heat%20Treating%20of%20Aluminum%20Alloys.pdf?sequence=3&isAllowed=y", "document_text": "Copyright \u00a9 1991 ASM International\u24c7\n\n\nASM Handbook, Volume 4: Heat Treating\nASM Handbook Committee, p 841-879\nDOI: 10.1361/asmhba0001205\n\n\nAll rights reserved.\nwww.asminternational.org\n\n\nHeat Treating of Aluminum Alloys\n\n\nHEAT TREATING in its broadest sense,\nrefers to any of the heating and cooling\noperations that are performed for the pur-\npose of changing the mechanical properties,\nthe metallurgical structure, or the residual\nstress state of a metal product. When the\nterm is applied to aluminum alloys, howev-\ner, its use frequently is restricted to the\nspecific operations employed to increase\nstrength and hardness of the precipitation-\nhardenable wrought and cast alloys. These\nusually are referred to as the \"heat-treat-\nable\" alloys to distinguish them from those\nalloys in which no significant strengthening\ncan be achieved by heating and cooling. The\nlatter, generally referred to as \"non-heat-\ntreatable\" alloys, depend primarily on cold\nwork to increase strength. Heating to de-\ncrease strength and increase ductility (an-\nnealing) is used with alloys of both types;\nmetallurgical reactions may vary with type\nof alloy and with degree of softening desired.\nExcept for the low-temperature stabilization\ntreatment sometimes given for 5xxx series\nalloys (which is a mill treatment and not\ndiscussed in this article), complete or partial\nannealing treatments are the only ones used\nfor non-heat-treatable alloys. A general\noverview of these heat treatments is covered\nin the article \"Principles of Heat Treating of\nNonferrous Alloys\" in this Volume.\n\n\nAluminum-copper-magnesium\n(magnesium intensifies precipitation)\n\n\nThe mechanism of strengthening from\nprecipitation involves the formation of co-\nherent clusters of solute atoms (that is, the\nsolute atoms have collected into a cluster\nbut still have the same crystal structure as\nthe solvent phase). This causes a great deal\nof strain because of mismatch in size be-\ntween the solvent and solute atoms. Conse-\nquently, the presence of the precipitate par-\nticles, and even more importantly the strain\nfields in the matrix surrounding the coher-\nent particles, provide higher strength by\nobstructing and retarding the movement of\ndislocations. The characteristic that deter-\nmines whether a precipitate phase is coher-\nent or noncoherent is the closeness of\nmatch or degree of disregistry between\natomic spacings on the lattice of the matrix\nand on that of the precipitate. These\nchanges in properties result from the forma-\ntion of solute-rich microstructural domains,\nor GP zones.\nThe exact size, shape, and distribution of\nGP zones depend on the alloy in which they\nform and on the thermal and mechanical\nhistory of the specimen. Their shape can\n\n\nsystems\n\n\n\u2022\n\n\n\u2022 Aluminum-magnesium-silicon\nwith strengthening from Mg\u2082Si\nAluminum-zinc-magnesium systems with\nstrengthening from MgZn2\n\n\nsystems\n\n\n\u2022\n\n\n\u2022 Aluminum-zinc-magnesium-copper\n\n\nsys-\n\n\ntems\n\n\nThe general requirement for precipitation\nstrengthening of supersaturated solid solu-\ntions involves the formation of finely dis-\npersed precipitates during aging heat treat-\nments (which may include either natural aging\nor artificial aging). The aging must be accom-\nplished not only below the equilibrium solvus\ntemperature, but below a metastable miscibil-\nity gap called the Guinier-Preston (GP) zone\nsolvus line. The supersaturation of vacancies\nallows diffusion, and thus zone formation, to\noccur much faster than expected from equi-\nlibrium diffusion coefficients. In the precipi-\ntation process, the saturated solid solution\nfirst develops solute clusters, which then be-\ncome involved in the formation of transitional\n(nonequilibrium) precipitates.\n\n\n800\n\n\n1400\n\n\n(b)\n\n\nL\n\n\nPrecipitation from Solid Solution\n\n\n600\n\n\nOne essential attribute of a precipitation-\nhardening alloy system is a temperature-\ndependent equilibrium solid solubility char-\nacterized by increasing solubility with\nincreasing temperature (see, for example,\nthe phase diagrams in Fig 1 and 2). Al-\nthough this condition is met by most of the\nbinary aluminum alloy systems, many ex-\nhibit very little precipitation hardening, and\nthese alloys ordinarily are not considered\nheat treatable. Alloys of the binary alumi-\nnum-silicon and aluminum-manganese sys-\ntems, for example, exhibit relatively insig-\nnificant changes in mechanical properties as\na result of heat treatments that produce\nconsiderable precipitation. The major alu-\nminum alloy systems with precipitation\nhardening include:\n\n\nA1 + L\n\n\n1000\n\n\nTemperature range for\nsolution heat treating\n\n\nTemperature, \u00b0C\n\n\nTemperature, \u00b0F\n\n\n\u0391\u0399\n\n\n400\n\n\nTemperature range\n\n\nfor annealing\n\n\n600\n\n\nTemperature range for\n\n\nprecipitation heat\ntreating\n\n\n200\n\n\n200\n\n\nAl + CuAl2\n\n\n2\n\n\n6\n\n\n4\n\n\n8\n\n\n10\n\n\n12\n\n\nCopper.%\n\n\nAluminum-copper systems with strength- Fig 1 Portion of aluminum-copper binary phase diagram. Temperature ranges for annealing, precipitation heat\n\n\n\u2022\n\n\ntreating, and solution heat treating are indicated. The range for solution treating is below the eutectic\nmelting point of 548 \u00b0C (1018 \u00b0F) at 5.65 wt% Cu.\n\n\nening from CuAl2\n\n\n842 / Heat Treating of Nonferrous Alloys\n\n\n700\n\n\nnar aggregates (GP zones), which form on\nparticular crystallographic planes of the alu-\nminum matrix. These aggregates create co-\nherency strain fields that increase resis-\ntance to deformation, and their formation is\nresponsible for the changes in mechanical\nproperties that occur during natural aging.\nAt higher temperatures, transition forms of\napproximate composition Al\u2082Cu develop\nand further increase strength. In the highest\nstrength condition, both the 0\" and 0' tran-\nsition precipitates may be present. When\ntime and temperature are increased suffi-\nciently to form high proportions of the equi-\nlibrium 0, the alloy softens and is said to be\n\"overaged.\"\nThe commercial heat-treatable aluminum\nalloys are, with few exceptions, based on\nternary or quaternary systems with respect\nto the solutes involved in developing\nstrength by precipitation. Commercial al-\nloys whose strength and hardness can be\nsignificantly increased by heat treatment\ninclude 2xxx, 6xxx, and 7xxx series wrought\nalloys (except 7072) and 2xx.0, 3xx.0, and\n7xx.0 series casting alloys. Some of these\ncontain only copper, or copper and silicon,\nas the primary strengthening alloy addi-\ntion(s). Most of the heat-treatable alloys,\nhowever, contain combinations of magne-\nsium with one or more of the elements\ncopper, silicon, and zinc. Characteristical-\nly, even small amounts of magnesium in\nconcert with these elements accelerate and\naccentuate precipitation hardening, while\nalloys in the 6xxx series contain silicon and\nmagnesium approximately in the propor-\ntions required for formulation of magnesium\nsilicide (Mg2Si). Although not as strong as\nmost 2xxx and 7xxx alloys, 6xxx series al-\nloys have good formability, weldability, ma-\nchinability, and corrosion resistance, with\nmedium strength.\nIn the heat-treatable wrought alloys, with\nsome notable exceptions (2024, 2219, and\n7178), such solute elements are present in\namounts that are within the limits of mutual\nsolid solubility at temperatures below the\neutectic temperature (lowest melting tem-\nperature). In contrast, some of the casting\nalloys of the 2xx.0 series and all of the 3xx.0\nseries alloys contain amounts of soluble\nelements that far exceed solid-solubility\nlimits. In these alloys, the phase formed by\ncombination of the excess soluble elements\nwith the aluminum will never be dissolved,\nalthough the shapes of the undissolved par-\nticles may be changed by partial solution.\nMost of the heat-treatable aluminum alloy\nsystems exhibit multistage precipitation and\nundergo accompanying strength changes\nanalogous to those of the aluminum-copper\nsystem. Multiple alloying additions of both\nmajor solute elements and supplementary\nelements employed in commercial alloys are\nstrictly functional and serve with different\nheat treatments to provide the many differ-\nent combinations of properties\u2014physical,\n\n\nThe GP zones are characteristically meta-\nstable and thus dissolve in the presence of a\nmore stable precipitate. This dissolution\ncauses a precipitate-free, visibly denuded\nregion to form around the stable precipitate\nparticles. The final structure consists of\nequilibrium precipitates, which do not con-\ntribute as significantly to hardening. More\ndetailed information about preprecipitation\nphenomena can be found in the article\n\"Structures Resulting From Precipitation\nFrom Solid Solution\" in Volume 9 of the\n9th Edition of Metals Handbook.\nPrecipitation in Aluminum-Copper Alloys.\nFigure 1, which illustrates the required sol-\nubility-temperature relationship needed in\nprecipitation strengthening, shows the tem-\nperature ranges required for solution treat-\nment and subsequent precipitate hardening\nin the aluminum-copper system. The equi-\nlibrium solid solubility of copper in alumi-\nnum increases as temperature increases-\nfrom about 0.20% at 250 \u00b0C (480 \u00b0F) to a\nmaximum of 5.65% at the eutectic melting\ntemperature of 548 \u00b0C (1018 \u00b0F). (It is con-\nsiderably lower than 0.20% at temperatures\nbelow 250 \u00b0C.) For aluminum-copper alloys\ncontaining from 0.2 to 5.6% Cu, two distinct\nequilibrium solid states are possible. At\ntemperatures above the lower curve in Fig 1\n(solvus), the copper is completely soluble,\nand when the alloy is held at such temper-\natures for sufficient time to permit needed\ndiffusion, the copper will be taken com-\npletely into solid solution. At temperatures\nbelow the solvus, the equilibrium state con-\nsists of two solid phases: solid solution, a,\nplus an intermetallic-compound phase 0\n(Al\u2082Cu). When such an alloy is converted to\nall solid solution by holding above the sol-\nvus temperature and then the temperature is\ndecreased to below the solvus, the solid\nsolution becomes supersaturated and the\nalloy seeks the equilibrium two-phase con-\ndition; the second phase tends to form by\nsolid-state precipitation.\nThe preceding description is a gross over-\nsimplification of the actual changes that\noccur under different conditions even in\nsimple binary aluminum-copper alloys. A\nvariety of different nonequilibrium precipi-\ntate structures is formed at temperatures\nbelow solvus. In alloys of the aluminum-\ncopper system, a succession of precipitates\nis developed from a rapidly cooled super-\nsaturated solid solution (SSS). These pre-\ncipitates develop sequentially either with\nincreasing temperature or with increasing\ntime at temperature between room temper-\nature and the solvus. The several stages are\nidentified by the following notation:\n\n\n1200\n\n\n600\n\n\nSolidus\n\n\n1000\n\n\n500\n\n\n595 \u00b0C\nat 1.85%\nMg2Si\n\n\nTemperature, \u00b0C\n\n\nTemperature, \u00b0F\n\n\n800\n\n\n400\n\n\nSolvus\n\n\n600\n\n\n300\n\n\n400\n\n\n200\n\n\n100\n\n\n200\n\n\nMg-Si ratio of 1.73:1\n\n\n1\n\n\n0\n\n\n0.5\n\n\n1.0\n\n\n1.5\n\n\n2.0\n\n\nMg2Si, %\n\n\n(a)\n\n\nTemperature, \u00b0F\n\n\n660 750 840 930 1020 1110\n\n\n570\n\n\n1.0\n\n\n1.4\n\n\n1.2\n\n\nSolvus with silicon\nand Mg2Si present\n\n\n0.8\n\n\nMagnesium, wt%\n\n\n1.0\n\n\nSilicon, wt%\n\n\n0.6\n\n\n0.8\n\n\n0.6\n\n\n0.4\n\n\n0.4\n\n\n0.2\n\n\n0.2\n\n\n0\n\n\n0\n\n\n300 350 400 450 500 550 600\nTemperature, \u00b0C\n\n\n(b)\n\n\nFig 2 Equilibrium solubility as function of tempera-\nture for (a) Mg2Si in aluminum with an Mg-Si\nratio of 1.73-to-1 and (b) magnesium and silicon in solid\naluminum when both Mg2Si and silicon are present\n\n\nsometimes be deduced by refined studies of\ndiffuse x-ray scattering. Under favorable\nconditions, GP zones can be seen in trans-\nmission electron micrographs. Spherical\nsolute-rich zones usually form when the\nsizes of the solvent and solute atoms are\nnearly equal, as in the aluminum-silver and\naluminum-zinc systems. If there is a large\ndifference in atom sizes, as in the alumi-\nnum-copper system, the GP zones usually\nform as disks whose planes lie parallel with\nsome low-index plane of the matrix lattice.\nSometimes, the solute atoms occupy pre-\nferred lattice sites within the GP zone, and\nthus form a small region of lattice order.\nThe GP zones are of the size range of tens\nof angstroms in diameter. They are essential-\nly distorted regions of the matrix lattice, rath-\ner than discrete particles of a new phase\nhaving a different lattice. As such, they are\ncompletely coherent with the matrix, impos-\ning local but often large strains on it. These\nmechanical strains, as well as the presence of\na locally solute-rich, sometimes ordered lat-\ntice, can account for large changes in me-\nchanical properties of the alloy before any\nlong-range microstructural changes occur.\n\n\nSSSGP zones \u21920\" \u2192\n0'\u21920 (Al\u2082Cu)\n\n\nAt temperatures in the natural aging\nrange (about 20 to 60 \u00b0C, or 0 to 140 \u00b0F),\nthe distribution of copper atoms changes\nwith time from random to the disklike pla-\n\n\nHeat Treating of Aluminum Alloys / 843\n\n\nminum-lithium alloys aimed at optimizing\nmechanical properties after artificial aging.\nThe age hardening of aluminum-lithium\nalloys involves the continuous precipitation\nof d' (AlLi) from a supersaturated solid\nsolution. The aluminum and lithium in the 8'\nprecipitates are positioned at specific loca-\ntions. The eight shared corner sites are\noccupied by lithium, and the six shared\nfaces are occupied by aluminum. This gives\nrise to the aluminum-lithium composition of\n8' precipitates. The geometrical similarity\nbetween the lattice of the precipitates and\nthe face-centered cubic lattices of the solid\nsolution facilitates the observed cube/cube\norientation. The lattice parameters of the\nprecipitate are also closely matched to\nthose of the matrix. Consequently, the mi-\ncrostructure of an aluminum-lithium alloy\nsolution heat treated and aged for short\ntimes below the d' solvus is characterized\nby a homogeneous distribution of coherent,\nspherical d' precipitates.\nAluminum-lithium-base alloys are micro-\nstructurally unique. They differ from most\nof the aluminum alloys in that once the\nmajor strengthening precipitate (8') is ho-\nmogeneously precipitated, it remains coher-\nent even after extensive aging. In addition,\nextensive aging at high temperatures (>190\n\u00b0C, or 375 \u00b0F) can result in the precipitation\nof icosahedral grain-boundary precipitates\nwith five-fold symmetry. Although the qua-\nsi-crystalline structure and the composition\nof these grain-boundary precipitates are not\nyet exactly known, it has been suggested\nthat both the precipitates and the precipi-\ntate-free zones (PFZs) near the grain bound-\naries might play a major role in the fracture\nprocess.\nThe low ductility and toughness of binary\naluminum-lithium alloys can be traced, at\nleast in part, to the inhomogeneous nature\nof their slip, resulting from coherent-parti-\ncle hardening of spherical d' precipitates.\nThe presence of equilibrium 6 (aluminum-\nlithium) precipitates at grain boundaries can\nalso cause PFZs, which can induce further\nstrain localization and promote intergranu-\nlar failure. Consequently, for the develop-\nment of commercial alloys, slip has been\nhomogenized by introducing dispersoids\n(manganese, zirconium) and semicoherent/\nincoherent precipitates, such as\n(Al\u2082CuLi), 0' (Al\u2082Cu), or S (Al\u2082LiMg),\nthrough copper or magnesium additions.\nMagnesium and copper improve the\nstrength of aluminum-lithium alloys through\nsolid-solution and precipitate strengthening,\nand they can minimize the formation of\nPFZs near grain boundaries. Zirconium,\nwhich forms the cubic Al3Zr coherent dis-\npersoid, stabilizes the subgrain structure\nand suppresses recrystallization.\nDevelopment of commercially available\naluminum-lithium-base alloys was started\nby adding lithium to aluminum-copper, alu-\nminum-magnesium, and aluminum-copper-\n\n\nmagnesium alloys. These alloys were cho-\nsen to superimpose the precipitation-\nhardening characteristics of aluminum-\ncopper-, aluminum-copper-magnesium-,\nand aluminum-magnesium-base precipitates\nto the hardening of lithium-containing pre-\ncipitates. Proceeding in this manner, alloys\n2020 (Al-Cu-Li-Cd), 01429 (Al-Mg-Li), 2090\n(Al-Cu-Li), and 2091 and 8090 (Al-Cu-Mg-\nLi) evolved. Besides these registered al-\nloys, other commercial aluminum-lithium\nalloys include Weldalite 049 and CP276.\nProperties and applications of these alloys\nare discussed in the article \"Aluminum-\nLithium Alloys\" in Volume 2 of the 10th\nEdition of Metals Handbook.\nIn terms of d' precipitation, the only\neffect of magnesium appears to be a reduc-\ntion in the solubility of lithium. The micro-\nstructure of an aluminum-magnesium-lithi-\num alloy in the early stages of aging is\nsimilar to that of an aluminum-lithium alloy.\nPrecipitation in the aluminum-copper-lithi-\num system is more complicated than that in\neither the aluminum-lithium or aluminum-\nmagnesium-lithium systems.\nEffects on Physical and Electrochemical\nProperties. The above description of the\nprecipitation processes in commercial heat-\ntreatable aluminum alloys (as well as the\nheat-treatable binary alloys, none of which\nis used commercially in wrought form) af-\nfect not only mechanical properties but also\nphysical properties (density and electrical\nand thermal conductivities) and electro-\nchemical properties (solution potential). On\nthe microstructural and submicroscopic\nscales, the electrochemical properties de-\nvelop point-to-point nonuniformities that\naccount for changes in corrosion resistance.\nMeasurements of changes in physical and\nelectrochemical properties have played an\nimportant role in completely describing pre-\ncipitation reactions and are very useful in\nanalyzing or diagnosing whether heat-treat-\nable products have been properly or im-\nproperly heat treated. Although they may\nbe indicative of the strength levels of prod-\nucts, they cannot be relied upon to deter-\nmine whether or not the product meets\nspecified mechanical-property limits. Since\nelements in solid solution are always more\nharmful to electrical conductivity than the\nsame elements combined with others as\nintermetallic compounds, thermal treat-\nments are applied to ingots used for fabri-\ncation of electrical conductor parts. These\nthermal treatments are intended to precipi-\ntate as much as possible of the dissolved\nimpurities. Iron is the principal element\ninvolved, and although the amount precip-\nitated is only a few hundredths of a percent,\nthe effect on electrical conductivity of the\nwire, cable, or other product made from the\ningot is of considerable practical impor-\ntance. These alloys may or may not be heat\ntreatable with respect to mechanical prop-\nerties. Electrical conductor alloys 6101 and\n\n\nmechanical, and electrochemical-that are\nrequired for different applications. Some al-\nloys, particularly those for foundry produc-\ntion of castings, contain amounts of silicon\nfar in excess of the amount that is soluble or\nneeded for strengthening alone. The function\nhere is chiefly to improve casting soundness\nand freedom from cracking, but the excess\nsilicon also serves to increase wear resis-\ntance, as do other microstructural constitu-\nents formed by manganese, nickel, and iron.\nParts made of such alloys are commonly\nused in gasoline and diesel engines (pistons,\ncylinder blocks, and so forth).\nAlloys containing the elements silver,\nlithium, and germanium are also capable of\nproviding high strength with heat treatment,\nand in the case of lithium, both increased\nelastic modulus and lower density, which\nare highly advantageous-particularly for\naerospace applications (see the following\nsection \"Aluminum-Lithium Alloys\" in this\narticle). Commercial use of alloys contain-\ning these elements has been restricted either\nby cost or by difficulties encountered in\nproducing them. Such alloys are used to\nsome extent, however, and research is be-\ning directed toward overcoming their disad-\nvantages.\nIn the case of alloys having copper as the\nprincipal alloying ingredient and no magne-\nsium, strengthening by precipitation can be\ngreatly increased by adding small fractional\npercentages of tin, cadmium, or indium, or\ncombinations of these elements. Alloys\nbased on these effects have been produced\ncommercially but not in large volumes be-\ncause of costly special practices and limita-\ntions required in processing, and in the case\nof cadmium, the need for special facilities to\navoid health hazards from formation and\nrelease of cadmium vapor during alloying.\nSuch alloys, as well as those containing\nsilver, lithium, or other particle-forming el-\nements, may be used on a selective basis in\nthe future.\nAluminum-Lithium Alloys. Like other age-\nhardened aluminum alloys, aluminum-lithi-\num alloys achieve precipitation strengthen-\ning by thermal aging after a solution heat\ntreatment. The precipitate structure is sen-\nsitive to a number of processing variables,\nincluding, but not limited to, the quenching\nrate following the solution heat treatment,\nthe degree of cold deformation prior to\naging, and the aging time and temperature.\nMinor alloying elements can also have a\nsignificant effect on the aging process by\nchanging the interface energy of the precip-\nitate, by increasing the vacancy concentra-\ntion, and/or by raising the critical tempera-\nture for homogeneous precipitation. Like\nsome other age-hardened 2xxx aluminum\nalloys, aluminum-lithium-base alloys also\ngain increased strength and toughness from\ndeformation prior to aging. This unusual\nphenomenon has given rise to a number of\nthermomechanical processing steps for alu-\n\n\nT\u2081\n\n\n844 / Heat Treating of Nonferrous Alloys\n\n\n6201 are heat treatable. These alloys are\nused in tempers in which their strengthening\nprecipitate, the transition form of Mg\u2082Si, is\nlargely out of solid solution to optimize both\nstrength and conductivity.\n\n\nshould, when possible, be above the temper-\nature at which complete solution occurs (sol-\nvus). In the alloy represented by line (a) in Fig\n1, these temperatures would be about 575 and\n515 \u00b0C (1065 and 960 \u00b0F), respectively. How-\never, under production conditions, the tem-\nperature interval for solution treatment\n(shown in Fig 1 for typical 2xxx or 2xx.x)\nalloys provides a margin to safeguard against\neutectic melting and a cushion on the low side\nfor increased solution and diffusion rates.\nFor alloys containing more than 5.65%\nCu, complete solution can never occur. For\nthese alloys, such as alloy 2219 (which has\n5.8 to 6.8% Cu), the minimum solution heat-\ntreating temperature is established so that it\nis as close as practical to the eutectic tem-\nperature while providing a margin of safety\ncommensurate with the capability of the\nequipment. Line (b) in Fig 1 is another\nexample of a composition above 5.65% Cu\nthat does not allow complete dissolution of\naluminum-copper precipitates.\nFor more complex ternary and quaterna-\nry systems, solution treatments are modi-\nfied according to the effect of new elements\non the solid solubility and/or the eutectic\nmelting points of the basic binary system. In\naluminum-lithium alloys, for example, mag-\nnesium reduces the solubility of lithium in\naluminum. In the aluminum-copper system,\nmagnesium also lowers the eutectic melting\npoint. The proximity of typical solution-\ntreating temperature ranges to eutectic\nmelting temperatures for three common alu-\nminum-copper-magnesium alloys is shown\nin the following table:\n\n\naluminum and particles of Al\u2082Cu. When this\nproduct is heated slowly, the Al\u2082Cu begins to\ndissolve, and if heating is slow enough, all of\nthe Al,Cu is dissolved when temperatures\nabove the solvus (500 \u00b0C, or 932 \u00b0F) are\nreached. When the heating rate is high, how-\never, much of the Al\u2082Cu remains undis-\nsolved. If a material with this microstructure\nis heated at or above the eutectic temperature\nof 548 \u00b0C (1018 \u00b0F), melting will begin at the\ninterface between the Al\u2082Cu and the matrix.\nWith sufficient time above the eutectic tem-\nperature, this metastable liquid will dissolve\nto form a solid solution and will leave no trace\nprovided that hydrogen gas has not con-\ndensed at the interface to form a void. If the\nproduct is quenched before the liquid has\ntime to equilibrate, however, it will solidify\nand form fine eutectic rosettes. This nonequi-\nlibrium melting should not be confused with\ntrue equilibrium melting, which would occur\nin any alloy containing more than 5.65% Cu.\nIn such an alloy, eutectic melting is equilibri-\num melting. No matter how long such an\nalloy is held above the eutectic temperature,\nthe liquid will never solidify. In commercial\nalloys, which usually are ternaries or quater-\nnaries of the major alloying elements, the\nsituation is more complex. Different phases\nhave different solvus temperatures, and non-\nequilibrium melting may occur at different\ntemperatures depending on composition, size\nof precipitates, and rate of heating. When\nnew solution heat-treating equipment (which\nprovides higher heating rates) is employed,\ncareful examination of alloy microstructures\nshould be included as part of the certification\nprocess.\nUnderheating. When the temperatures at-\ntained by the parts or pieces being heat treat-\ned are appreciably below the normal range,\nsolution is incomplete, and strength some-\nwhat lower than normal is expected. In the\naluminum-copper system (Fig 1), the shallow\nslope of the solvus at its intersection with the\ncomposition line indicates that a slight de-\ncrease in temperature will result in a large\nreduction in the concentration of the solid\nsolution and a correspondingly significant de-\ncrease in final strength. The effect of solution-\ntreating temperature on the strength of two\naluminum alloys is illustrated by the following\ndata:\n\n\nStrengthening by Heat Treatment\n\n\nHeat treatment to increase strength of\naluminum alloys is a three-step process:\n\n\nSolution heat treatment: dissolution of\nsoluble phases\nQuenching: development of supersatura-\ntion\n\n\n\u2022\n\n\nAge hardening: precipitation of solute at-\noms either at room temperature (natural\naging) or elevated temperature (artificial\naging or precipitation heat treatment)\nEach of these steps and the use of quench-\nfactor analysis are described in the follow-\ning four sections. Typical solution and pre-\ncipitation heat treatments for mill products\nare given in Tables 1(a, b, and c) and 2, and\ntreatments for castings are given in Table 3.\nTemper designations are defined at the end\nof this article.\n\n\n\u2022\n\n\nSolution Heat Treating\n\n\nTo take advantage of the precipitation-\nhardening reaction, it is necessary first to\nproduce a solid solution. The process by\nwhich this is accomplished is called solution\nheat treating, and its objective is to take into\nsolid solution the maximum practical\namounts of the soluble hardening elements\nin the alloy. The process consists of soaking\nthe alloy at a temperature sufficiently high\nand for a time long enough to achieve a\nnearly homogeneous solid solution.\nNominal commercial solution heat-treat-\ning temperature is determined by the com-\nposition limits of the alloy and an allowance\nfor unintentional temperature variations.\nAlthough ranges normally listed allow vari-\nations of \u00b16 \u00b0C (\u00b110 \u00b0F) from the nominal,\nsome highly alloyed, controlled-toughness,\nhigh-strength alloys require that tempera-\nture be controlled within more restrictive\nlimits. Broader ranges may be allowable for\nalloys with greater intervals of temperature\nbetween their solvus and eutectic melting\ntemperatures.\nOverheating. Care must be exercised to\navoid exceeding the initial eutectic melting\ntemperature. If appreciable eutectic melting\noccurs as a result of overheating, properties\nsuch as tensile strength, ductility, and frac-\nture toughness may be degraded. Materials\nthat exhibit microstructural evidence of\noverheating are generally categorized as\nunacceptable by specification. Evidence of\ngrain-boundary melting that occurs above\nthe eutectic melting temperature of the alloy\nusually is not detectable by either visual\nexamination or nondestructive testing.\nAlthough maximum temperature must be\nrestricted to avoid melting, the lower limit\n\n\nSolution-treating\ntemperature\n\n\nEutectic melting\ntemperature\n\n\nAlloy\n\n\n\u00b0C\n\n\n\u00b0F\n\n\n\u00b0C\n\n\n\u00b0F\n\n\n2014\n2017\n2024\n\n\n496-507\n496-507\n488-499\n\n\n925-945\n925-945\n910-930\n\n\n950\n955\n935\n\n\n510\n513\n502\n\n\nSimilar considerations apply to other age-\nhardenable alloy systems such as aluminum-\nmagnesium-silicon alloys. For example, ac-\ncording to Fig 2(a), a 1.08% Mg\u2082Si alloy\nwould be soaked at a temperature in excess of\n500 \u00b0C (930 \u00b0F) but below the solidus of 595 \u00b0C\n(1100 \u00b0F) to avoid incipient melting. Howev-\ner, because some alloy constituents may form\ncomplex eutectics that melt at temperatures\nbelow the equilibrium eutectic temperature,\nthe upper limit for solution treatment of alu-\nminum-magnesium-silicon alloys is in the\nrange of 515 to 540 \u00b0C (960 to 1000 \u00b0F). At 540\n\u00b0C (1000 \u00b0F), about 0.6% Mg can be placed in\nsolution (Fig 2b).\nNonequilibrium Melting. When high heat-\ning rates are employed, the phenomenon of\nnonequilibrium melting must be considered.\nThis phenomenon can also be explained with\nthe help of the aluminum-copper phase dia-\ngram (Fig 1). The room-temperature micro-\nstructure of an F-temper product containing\n4% Cu consists of a solid solution of copper in\n\n\nSolution-\ntreating\ntemperature\n\n\nTensile strength\nksi\n\n\nYield strength\n\n\n\u00b0C\n\n\n\u00b0F\n\n\nksi\n\n\nMPa\n\n\nMPa\n\n\n6061-T6 sheet 1.6 mm (0.064 in.) thick\n\n\n272\n288\n305\n315\n\n\n39.4\n41.7\n44.3\n45.7\n\n\n493\n504\n516\n527\n\n\n920\n940\n960\n980\n\n\n43.7\n45.8\n48.3\n50.5\n\n\n301\n316\n333\n348\n\n\n2024-T4 sheet 0.8 mm (0.032 in.) thick\n\n\n488\n491\n493\n496\n\n\n419\n422\n433\n441\n\n\n255\n259\n269\n271\n\n\n37.0\n37.5\n39.0\n39.3\n\n\n910\n915\n920\n925\n\n\n60.8\n61.2\n62.8\n63.9\n\n\nHeat Treating of Aluminum Alloys / 845\n\n\nTable 1(a)\nalloying\n\n\nTypical solution and precipitation heat treatments for commercial heat-treatable aluminum alloy mill products with copper\n\n\nSolution heat treatment(a)\n\n\nPrecipitation heat treatment\n\n\nMetal temperature(b)\n\n\nMetal temperature(b)\n\n\nTemper\ndesignation\n\n\nTemper\ndesignation\n\n\nTime(c),\nh\n\n\nAlloy\n\n\nProduct form\n\n\n\u00b0C\n\n\n\u00b0F\n\n\n\u00b0F\n\n\n\u00b0C\n\n\nAl-Cu alloys without magnesium alloying\n\n\n525\n\n\n2011\n\n\nRolled or cold finished rod and bar\n\n\n975\n\n\nT3(d)\n\n\nT8(d)\n\n\n160\n\n\n320\n\n\n14\n\n\nT4\n\n\nT451(e)\n\n\n00000000000000\n10\n18\n24\n36\n18\n18\n18\n36\n18\n18\n18\n18\n36\n26\n26\n18\n\n\nDie forgings\n\n\n2025\n2219(f)\n\n\n515\n535\n\n\n960\n995\n\n\nT4\n\n\n170\n175\n165\n190\n175\n175\n175\n190\n190\n190\n190\n190\n190\n190\n190\n175\n\n\n340\n350\n325\n375\n350\n350\n350\n375\n375\n375\n375\n375\n375\n375\n375\n350\n\n\nT6\n\n\nFlat sheet\n\n\nT31(d)\nT37(d)\n\n\nT81(d)\nT87(d)\n\n\nT42\n\n\nT62\n\n\nT31(d)\nT37(d)\nT351(e)\n\n\nPlate\n\n\n535\n\n\nT81(d)\nT87(d)\nT851(e)\n\n\n995\n\n\nT62\n\n\nT42\n\n\nRolled or cold finished wire, rod, and bar\nExtruded rod, bar, shapes, and tube\n\n\nT351(e)\nT31(d)\n\n\nT851(e)\nT81(d)\n\n\n535\n535\n\n\n995\n995\n\n\n2219(f)\n\n\nT3510(e)\nT3511(e)\n\n\nT8510(e)\nT8511(e)\nT62\n\n\nT42\nT4\nT4\n\n\nDie forgings and rolled rings\n\n\n535\n535\n\n\n995\n995\n\n\nT6\nT6\nT852f)\n\n\nHand forgings\n\n\nT352(f)\n\n\nAl-Cu-Mg alloys\n\n\nDie forgings\n\n\n950(g)\n\n\n10\n02826:26282228622261 206\n9\n9\n9\n\n\n2018\n2024(h)\n\n\n510(g)\n\n\nT4\n\n\n170\n190\n190\n190\n190\n\n\n340\n375\n375\n375\n375\n\n\nT61\nT81(d)\nT861(d)\n\n\nFlat sheet\n\n\n495\n\n\n920\n\n\nT3(d)\n\n\nT361(d)\n\n\nT42\n\n\nT62\nT72\n\n\n2024(h)\n\n\n495\n\n\nCoiled sheet\n\n\n920\n\n\nT4\nT42\n\n\n375\n375\n375\n375\n375\n375\n375\n375\n375\n375\n375\n375\n375\n\n\nT62\nT72\n\n\n190\n190\n190\n190\n190\n190\n190\n190\n190\n190\n190\n190\n190\n\n\n920\n920\n22\n\n\nPlate\n\n\nT851(e)\nT861(d)\n\n\n495\n\n\nT351(e)\nT361(d)\n\n\nT42\nT4\n\n\nT62\nT6\n\n\nRolled or cold finished wire, rod, and bar\n\n\n495\n\n\nT851(e)\nT86(d)\n\n\nT351(e)\nT36(d)\n\n\nT42\nT3\n\n\nT62\nT81\n\n\nExtruded rod, bar, shapes, and tube\n\n\n495\n\n\n920\n\n\nT3510(e)\nT3511(e)\n\n\nT8510(e)\nT8511(e)\n\n\nT42\nT3(d)\nT42\nT4\nT4\nT4\nT41\n\n\nT62\n\n\nDrawn tube\n\n\n495\n\n\n920\n\n\nSheet\nSheet\n\n\n2036\n2038\n2218\n\n\n500\n540\n\n\n930\n1000\n\n\n205\n170\n240\n\n\n400\n340\n460\n\n\nT6\nT61\nT72\n\n\nDie forgings\n\n\n510(g)\n510(i)\n\n\n950(g)\n950(i)\n\n\nAl-Cu-Mg-Si alloys\n\n\nT4(d)(j)\nT3(d)\nT42\nT4\nT42\nT42\n\n\n510\n500\n\n\nT62(e)\n\n\n2008\n\n\nSheet\n\n\n950\n935\n\n\n205\n160\n160\n160\n160\n160\n160\n\n\n400\n320\n320\n320\n320\n320\n320\n\n\n1\n18\n18\n18\n18\n18\n18\n18\n18\n18\n18\n18\n18\n18\n18\n10\n\n\nFlat sheet\n\n\n2014(h)\n\n\nT62\nT6\nT6\nT62\nT62\n\n\nCoiled sheet\n\n\n935\n\n\n500\n\n\nPlate\n\n\n500\n\n\n935\n\n\nT451(e)\n\n\nT651(e)\n\n\nRolled or cold finished wire, rod, and bar\n\n\n935\n\n\n320(k)\n320(k)\n320(k)\n320(k)\n320(k)\n320(k)\n320(k)\n320(k)\n\n\n500\n\n\nT4\nT42\n\n\n160(k)\n160(k)\n160(k)\n160(k)\n160(k)\n160(k)\n160(k)\n160(k)\n\n\nT6\n\n\nT62\nT651(e)\n\n\nT451(e)\n\n\nExtruded rod, bar, shapes, and tube\n\n\n500\n\n\n935\n\n\nT4\nT42\n\n\nT6\n\n\nT62\nT6510(e)\n\n\nT4510(e)\n\n\n935\n\n\nDrawn tube\n\n\n500\n\n\nT4\nT42\nT4\nT4\nT42\nT4\nT42\nT4\nT4\n\n\nT6\nT62\nT6\n\n\nDie forgings\n\n\n935(1)\n\n\n500(1)\n\n\n170\n\n\n340\n\n\nRolled or cold finished wire, rod, and bar\n\n\n2017\n\n\n500\n\n\n935\n\n\n2117\n\n\nRolled or cold finished wire and rod\n\n\n500\n\n\n935\n\n\nForgings and rolled rings\n\n\n2618\n4032\n\n\n530\n\n\n985\n\n\n200\n170\n\n\n390\n340\n\n\nT61\nT6\n\n\n20\n10\n\n\nDie forgings\n\n\n510(h)\n\n\n950(h)\n\n\nAl-Cu-Li alloys\n\n\n2090\n2091\n\n\nSheet\nSheet\n\n\n540\n530\n530\n530\n540\n\n\nT3(d)\nT3(d)\nT3(d)\nT3(d)\nT3(d)\n\n\n165\n120\n190\n190\n190\n\n\n1000\n990\n990\n990\n1000\n\n\n325\n250\n375\n375\n375\n\n\nT83(d)\nT84(d)\n\n\n24\n24\n12\n12\n12-15\n\n\nPeak aged(d)\nPeak aged(d)\nPeak aged(d)\n\n\nExtruded bar\nExtruded bar\nExtruded bar\n\n\n8090\nCP276\n\n\n(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water.\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) Cold working subsequent to solution heat treatment and prior to any precipitation heat treatment is necessary\nto attain the specified properties for this temper. (e) Stress relieved by stretching to produce a specified amount of permanent set subsequent to solution heat treatment and prior to any precipitation heat\ntreatment. (f) Stress relieved by 1 to 5% cold reduction after solution treatment and prior to precipitation heat treatment. (g) Quenched in water at 100 \u00b0C (212 \u00b0F). (h) These heat treatments also apply to\nalclad sheet and plate of these alloys. (i) Quenched with room-temperature air blast. (j) See U.S. Patent 4,840,852. (k) An alternative heat treatment of 8 h at 177 \u00b0C (350 \u00b0F) may also be used. (1) Quenched\nin water at 60 to 80 \u00b0C (140 to 180 \u00b0F).\n\n\n846 / Heat Treating of Nonferrous Alloys\n\n\nTable 1(b) Typical solution and precipitation heat treatments for Mg-Si aluminum alloys (6xxx series alloys)\n\n\nSolution heat treatment(a)\n\n\nPrecipitation heat treatment\n\n\nMetal temperature(b)\n\n\nMetal temperature(b)\n\n\nTemper\ndesignation\n\n\nTemper\ndesignation\n\n\nProduct form\n\n\nTime(c), h\n\n\n\u00b0F\n\n\n\u00b0C\n\n\n\u00b0F\n\n\nAlloy\n6005\n\n\n\u00b0C\n\n\n530(d)\n\n\nExtruded rod, bar, shapes,\n\n\n175\n\n\nT5\n\n\n985(d)\n\n\nTI\n\n\n8\n\n\n350\n\n\nand tube\n\n\nSheet\nSheet\n\n\nT6(e)\nT6(e)\n\n\n1\n1\n10\n18\n18\n18\n18\n18\n18\n\n\n6009(e)\n\n\n555\n565\n520\n530\n\n\n1030\n1050\n970\n985\n\n\nT4\nT4\nT4\nT4\nT42\nT4(g)\nT42\n\n\n205\n205\n170\n160\n160\n160\n160\n160\n\n\n400\n400\n340\n320\n320\n320\n320\n320\n320(i)\n\n\n6010\n6053\n\n\nT6\nT6\nT62\nT6(g)\nT62\n\n\nDie forgings\n\n\n6061(f)\n\n\nSheet\n\n\nPlate\n\n\n530\n\n\n985\n\n\nT451(h)\n\n\nT651(h)\n\n\n160(i)\n\n\nT6\n\n\nRolled or cold finished\nwire, rod, and bar\n\n\n530\n\n\n985\n\n\nT4\n\n\n160(i)\n160(i)\n160(i)\n160(i)\n160(i)\n160(i)\n175\n\n\nT89(j)\nT93(k)\nT913(k)\nT94(k)\n\n\n320(i)\n320(i)\n320(i)\n320(i)\n320(i)\n320(i)\n\n\n18\n18\n18\n18\n18\n18\n8\n\n\nT62\n\n\nT42\n\n\nT451(h)\n\n\nT651(h)\n\n\nExtruded rod, bar, shapes,\nand tube\n\n\n530(d)\n\n\n985(d)\n\n\nT6\n\n\nT4\n\n\n350\n\n\nT6510(h)\nT6511(h)\n\n\nT4510(h)\nT4511(h)\n\n\n175\n175\n175\n\n\n350\n350\n350\n\n\n8\n8\n8\n18\n18\n8\n8\n8\n1\n\n\nT62\nT6\nT62\nT6\n\n\nT42\nT4\nT42\nT4\nT4\n\n\n530\n530\n\n\n985\n985\n\n\n160(i)\n160(i)\n175\n175\n175\n\n\n6061(f)\n\n\nDrawn tube\n\n\n320(i)\n320(i)\n350\n350\n350\n\n\nDie and hand forgings\n\n\n530\n530\n\n\n985\n985\n\n\nRolled rings\n\n\nT6\nT652(1)\n\n\nT452(1)\n\n\nExtruded rod, bar, shapes,\n\n\nTI\n\n\nT5\n\n\n(d)\n\n\n205(m)\n\n\n400(m)\n\n\n6063\n\n\n(d)\n\n\nand tube\n\n\n\u221e \u221e \u221e \u221e \u221e \u221e \u221e \u221e\n8\n8\n8\n8\n8\n8\n8\n4\n4\n8\n\n\n175(n)\n175(n)\n\n\n350(n)\n350(n)\n\n\n970(d)\n\n\nT6\nT62\nT6\n\n\nT4\nT42\nT4\n\n\n520(d)\n\n\n970\n970\n\n\n520\n520\n\n\nDrawn tube\n\n\n175\n175\n175\n175\n175\n190\n190\n175\n\n\n350\n350\n350\n350\n350\n375\n375\n350\n\n\nT83(j)(d)\nT831(j)(d)\nT832(j)(d)\n\n\nT62\nT6\n\n\nT42\nW(p)\nW(p)\n\n\n6013(0)\n\n\nSheet\nPlate\n\n\n570\n570\n530\n\n\n1055\n1055\n990\n\n\nT651\nT6\n\n\nExtruded rod, bar, shapes,\n\n\n6066\n\n\nT4\n\n\nand tube\n\n\nao ao ao ao \u221e \u221e \u221e\n8\n8\n8\n8\n8\n8\n18\n\n\nT42\nT4\nT42\nT4\nT4\nT42\nT4\nT4\nT4\nT4\nT42\nT4\nT42\nT4\nT42\nTI\nT4\nT42\nE = FEE FEE FERRE PREFECE\nT451\nT4\nT42\n\n\nT62\n\n\n175\n175\n175\n175\n175\n175\n160\n\n\n350\n350\n350\n350\n350\n350\n320\n\n\nT6510(h)\nT6511(h)\n\n\nT4510(h)\nT4511(h)\n\n\nDrawn tube\n\n\n990\n\n\nT6\nT62\nT6\nT6\n\n\n530\n\n\nDie forgings\n\n\n990\n1015(d)\n\n\n530\n\n\nExtruded rod, bar, shapes,\n\n\n545(d)\n\n\n6070\n\n\nand tube\n\n\n18\n8\n10\n10\n10\n8\n\n\nT62\nT6(q)\n\n\n160\n175\n170\n170\n170\n170\n\n\n320\n350\n340\n340\n340\n340\n\n\nSheet\n\n\n6111\n6151\n\n\n560\n515\n515\n\n\n1040\n960\n960\n\n\nDie forgings\nRolled rings\n\n\nT6\nT6\n\n\nT452(1)\n\n\nT652(1)\n\n\nRolled or cold finished\nwire, rod, and bar\n\n\n6262\n\n\n540\n\n\nT6\n\n\n1000\n\n\n12\n12\n2882\n\n\nT9(k)\nT651(h)\n\n\n170\n170\n170\n175\n\n\n340\n340\n340\n350\n\n\nT62\nT6\n\n\n540(d)\n\n\nExtruded rod, bar, shapes,\nand tube\n\n\n6262\n\n\n1000(d)\n\n\nT6510(h)\nT62\n\n\nT4510(h)\n\n\n175\n175\n170\n170\n170\n\n\n350\n350\n340\n340\n340\n\n\n12\n12\n8\n8\n8\n1\n\n\n540\n540\n\n\n1000\n1000\n\n\nT6\n\n\nDrawn tube\n\n\nT9(k)\nT62\nT5\n\n\nExtruded rod, bar, shapes,\nand tube\n\n\n6463\n\n\n(d)\n\n\n(d)\n\n\n205(m)\n\n\n400(m)\n\n\n350(n)\n350(n)\n\n\nT6\nT62\nT6\nT62\n\n\n970(d)\n970\n985\n\n\n8\n8\n18\n18\n\n\n175(n)\n175(n)\n\n\n520(d)\n\n\n520\n530\n\n\n320\n320\n\n\nSheet\n\n\n160\n160\n\n\n6951\n\n\n(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water,\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) By suitable control of extrusion temperature, product may be quenched directly from extrusion press to\nprovide specified properties for this temper. Some products may be adequately quenched in room-temperature air blast. (e) Alternate heat treatments of 4 h at 190 \u00b0C (375 \u00b0F) or 8 h at 175 \u00b0C (350 \u00b0F) may\nalso be used. See U.S. Patent 4,082,578. (f) These heat treatments also apply to alclad sheet and plate in these alloys. (g) Applicable to tread plate only. (h) Stress relieved by stretching to produce a specified\namount of permanent set prior to precipitation heat treatment. (i) An alternative heat treatment of 8 h at 170 \u00b0C (340 \u00b0F) also may be used. (j) Cold working after solution treatment is necessary to attain\nspecified properties during precipitation heat treatments. (k) Cold working after precipitation heat treatment is necessary to attain specified properties. (1) Stress relieved by 1 to 5% cold reduction subsequent\nto solution heat treatment and prior to precipitation heat treatment. (m) An alternative treatment of 3 h at 182 \u00b0C (360 \u00b0F) also may be used. (n) An alternative treatment of 6 h at 182 \u00b0C (360 \u00b0F) also may\nbe used. (o) See U.S. Patent 4,589,932. (p) Two weeks of natural aging to a T4 condition. (q) Artificially aged in laboratory from T4 to T6.\n\n\nHeat Treating of Aluminum Alloys / 847\n\n\nTable 1(c)\n\n\nTypical solution and precipitation heat treatments for heat-treatable Zn-Mg aluminum alloys from the 7xxx series\n\n\nSolution heat treatment(a)\n\n\nPrecipitation heat treatment\n\n\nMetal temperature(b)\n\n\nMetal temperature(b)\n\n\nTime(c),\nh\n\n\nTemper\ndesignation\n\n\nTemper\ndesignation\n\n\nAlloy\n\n\n\u00b0C\n\n\nProduct form\n\n\n\u00b0F\n\n\n\u00b0C\n\n\n\u00b0F\n\n\n7001\n\n\nExtruded rod, bar, shapes,\nand tube\n\n\n465\n\n\n870\n\n\n24\n\n\nW\n\n\n120\n\n\n250\n\n\nT6\n\n\n120\n120\n120\n\n\nT62\n\n\n250\n250\n250\n\n\n24\n\n\nW510(d)\nW511(d)\n\n\nT6510(d)\nT6511(d)\n\n\n24\n24\n\n\n7005\n\n\nExtruded rod, bar, and\n\n\nT53(e)\n\n\n:\n\n\nshapes\n\n\n7050\n\n\nW51(d)\n\n\n(f)\n(h)\n(f)\n(f)\n(h)\n(h)\n120(j)\n120(j)\n(f)\n(h)(k)\n120(j)\n(h)(k)\n120(j)\n(f)\n120\n\n\n(f)\n(h)\n(f)\n(f)\n(h)\n(h)\n250(j)\n250(j)\n(f)\n(h)(k)\n250(j)\n(h)(k)\n250(j)\n(f)\n250\n\n\nPlate\n\n\n475\n\n\n(f)\n(h)\n(f)\n(f)\n(h)\n(h)\n24\n24\n(f)\n(h)(k)\n24\n(h)(k)\n24\n(f)\n24\n\n\n890\n\n\nT7651(g)\nT7451(g)\nT76510(g)\nT76511(g)\nT74(g)\nT7452(g)\n\n\n475\n\n\nExtrusions\n\n\nW510(d)\nW511(d)\n\n\n890\n\n\nDie and hand forgings\n\n\n475\n\n\n890\n\n\nW\n\n\nW52(d)\n\n\n7075(i)\n\n\nSheet\n\n\nT6\nT62\n\n\n480\n\n\n900\n\n\nW\n\n\nT76(g)\nT73(g)\n\n\nPlate\n\n\n480\n\n\n900\n\n\nW\n\u0791\u0791\n\n\nT62\n\n\nT7351(d)(g)\nT651(d)\nT7651(g)\n\n\nW51(d)\n\n\n7075(i)\n\n\nRolled or cold finished\nwire, rod, and bar\n\n\n490\n\n\n915\n\n\nW\n\n\nT6\n\n\n120\n(h)(k)\n120\n(h)(k)\n120(1)\n\n\n250\n(h)(k)\n250\n(h)(k)\n250(1)\n\n\n24\n(h)(k)\n24\n(h)(k)\n24\n\n\nT62\nT73(g)\nT651(d)\nT7351(d)(g)\n\n\nW51(d)\n\n\nExtruded rod, bar, shapes,\nand tube\n\n\n465\n\n\n870\n\n\nW\n\n\nT6\n\n\n120(1)\n(h)(k)\n(f)\n120(1)\n(h)(k)\n(f)\n120(1)\n(h)(k)\n(f)\n120\n120\n(h)(k)\n120\n(h)\n(h)\n120\n(h)\n120\n(h)\n120\n(o)\n(o)\n(o)\n\n\n250(1)\n(h)(k)\n(f)\n250(1)\n(h)(k)\n(f)\n250(1)\n(h)(k)\n(f)\n250\n250\n(h)(k)\n250\n(h)\n(h)\n250\n(h)\n250\n(h)\n250\n(0)\n(o)\n(o)\n(0)\n(0)\n250\n315\n(f)\n250\n(f)\n(h)\n250\n315\n(f)\n\n\n24\n(h)(k)\n(f)\n24\n(h)(k)\n(f)\n24\n(h)(k)\n(f)\n24\n24\n(h)(k)\n24\n(h)\n(h)\n24\n(h)\n24\n(h)\n24\n(o)\n(o)\n\n\nT62\n\n\nT73(g)\n\n\nT76(g)\nT6510(d)\nT73510(d)(g)\nT76510(g)\nT6511(d)\nT73511(d)(g)\nT76511(g)\n\n\nW510(d)\n\n\nW511(d)\n\n\n=\nW\n\n\n870\n\n\nDrawn tube\n\n\n465\n\n\nT6\n\n\nT62\nT73(g)\n\n\n3\n\n\nDie forgings\n\n\n470(m)\n\n\n880(h)\n\n\nT6\nT73(g)\nT7352(n)(g)\nT6\nT73(g)\nT652(n)\nT7352(n)(g)\n\n\nW52(n)\n\n\nHand forgings\n\n\n880(h)\n\n\n470(m)\n\n\nW\n\n\nW52(n)\n\n\nRolled rings\nDie forgings\n\n\nBBBBBB\nW\nW\nW\nW\nW\n\n\n880\n(o)\n(o)\n(0)\n(0)\n(o)\n960(p)\n\n\n470\n(o)\n(o)\n\n\nT6\n\n\n7175\n\n\nT66(0)\nT74(g)(0)\nT7452(n)(g)(0)\nT74(g)(0)\nT7452(n)(g)(0)\n\n\nW52(n)\n\n\nHand forgings\n\n\n(o)\n(0)\n515(p)\n\n\nW52(n)\n\n\n(0)\n120\nplus 155\n(f)\n120\n(f)\n(h)\n120\nplus 155\n(f)\n\n\n(o)\n3\n3\n(f)\n24\n\n\n7475\n\n\nSheet\n\n\nT61(p)\nT761(g)(p)\nT651(p)\nT7651(g)(p)\nT7351(g)(p)\n\n\n510(p)\n\n\n950(p)\n\n\nPlate\n\n\nW51(d)\n\n\n(h)\n3\n3\n(f)\n\n\nAlclad\n\n\nSheet\n\n\n495\n\n\n920\n\n\nW\n\n\n7475\n\n\nT61(p)\nT761(g)(p)\n\n\n(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water,\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) Stress relieved by stretching to produce a specified amount of permanent set after solution treatment and\nprior to precipitation heat treatment. (e) No solution heat treatment; 72 h at room temperature following press quench, followed by two-stage precipitation heat treatment comprised of 8 h at 107 \u00b0C (225 \u00b0F)\nplus 16 h at 149 \u00b0C (300 \u00b0F). (f) Aging practice varies with product, size, nature of equipment, loading procedures, and furnace-control capabilities. The optimum practice for a specific item can be ascertained\nonly by actual trial treatment of the item under specific conditions. Typical procedures involve a two-stage treatment comprised of 3 to 30 h at 121 \u00b0C (250 \u00b0F) followed by 15 to 18 h at 163 \u00b0C (325 \u00b0F) for\nextrusions. An alternative two-stage treatment of 8 h at 99 \u00b0C (210 \u00b0F) followed by 24 to 28 h at 163 \u00b0C (325 \u00b0F) also may be used. (g) Aging of aluminum alloys 7050, 7075, 7175, and 7475 from any temper\nto the T73 or T76 temper series requires closer-than-normal controls on aging variables such as time, temperature, heatup rate, and so forth, for any given item. In addition, when material in a T6-type temper\nis reaged to a T73- or T76-type temper, the specific condition of the T6 material (such as property levels and other effects of processing variables) is extremely important and will affect the capability of the\nreaged material to conform to the requirements specified for the applicable T73- or T76-type temper. (h) Two-stage treatment comprised of 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by: 24 to 30 h at 163 \u00b0C (325 \u00b0F)\nfor sheet and plate: 8 to 10 h at 177 \u00b0C (350 \u00b0F) for rolled or cold finished rod and bar; 6 to 8 h at 177 \u00b0C (350 \u00b0F) for extrusions and tube; 8 to 10 h at 177 \u00b0C (350 \u00b0F) for forgings in the T73 temper; and 6\nto 8 h at 177 \u00b0C (350 \u00b0F) for forgings in the T7352 temper. (i) These heat treatments also apply to alclad sheet and plate of these alloys. (j) An alternative two-stage treatment comprised of 4 h at 96 \u00b0C (205 \u00b0F)\nfollowed by 8 h at 157 \u00b0C (315 \u00b0F) also may be used. (k) For sheet, plate, tube, and extrusions, an alternative two-stage treatment comprised for 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by 14 to 18 h at 168 \u00b0C\n(335 \u00b0F) may be used, provided that a heatup rate of approximately 14 \u00b0C/h (25 \u00b0F/h) is employed. For rolled or cold finished rod and bar, the alternative treatment is 10 h at 177 \u00b0C (350 \u00b0F). (I) An alternative\nthree-stage treatment comprised of 5 h at 99 \u00b0C (210 \u00b0F), 4 h at 121 \u00b0C (250 \u00b0F), and then 4 h at 149 \u00b0C (300 \u00b0F) may also be used. (m) Quenched in water at 60 to 80 \u00b0C (140 to 180 \u00b0F). (n) Stress relieved\nby 1 to 5% cold reduction after solution treatment and prior to precipitation heat treatments. (o) 7175-T74 and -T7452 heat treatments are directed to specific results, may vary from supplier to supplier and\nare either proprietary or patented. (p) Must be preceded by soak at 466 to 477 \u00b0C (870 to 890 \u00b0F). See U.S. Patent 3,791,880.\n\n\n848 / Heat Treating of Nonferrous Alloys\n\n\nTable 2 Soak times and maximum quench delays for solution treatment of wrought\naluminum alloys\n\n\nmore economical in the long run than heavi-\ner loading, because with lighter loads heat-\ning rates are higher and fewer rejections and\nservice failures are encountered.\nHigh-Temperature Oxidation. There is a\ncondition, commonly but erroneously\nknown as HTO or high-temperature oxida-\ntion, which can lead to deterioration of\nproperties in aluminum alloys. High-tem-\nperature oxidation is a misnamed condition\nof hydrogen diffusion that affects surface\nlayers during elevated-temperature treat-\nment. This condition can result from mois-\nture contamination in the furnace atmo-\nsphere and is sometimes aggravated by\nsulfur (as in heat-treatment furnaces also\nused for magnesium alloy castings) or other\nfurnace refractory contamination.\nMoisture in contact with aluminum at\nhigh temperatures serves as a source of\nnascent hydrogen, which diffuses into the\nmetal. Foreign materials, such as sulfur\ncompounds, function as decomposers of the\nnatural oxide surface film, eliminating it as a\nbarrier either between the moisture and the\naluminum or between the nascent hydrogen\nand the aluminum. The most common man-\nifestation of high-temperature oxidation is\nsurface blistering, but occasionally the only\nmanifestations are internal discontinuities\nor voids, which can be detected only by\ncareful ultrasonic inspection or by metallo-\ngraphic techniques.\nIt is important to recognize that the symp-\ntoms of high-temperature oxidation are\nidentical to those of unsoundness or high\ngas content in the original ingot or of other\nimproper mill practice. Blisters resulting\nfrom ingot defects, improper extrusion or\nimproper rolling may be lined up in the\ndirection of working. However, it usually is\nimpossible to distinguish among defect\nsources, and therefore the possibility that a\ncontaminated atmosphere is the cause of\nthe defects must be checked.\nNot all alloys and product forms are\nequally vulnerable to this type of attack.\nThe 7xxx series alloys are most susceptible,\nfollowed by the 2xxx alloys. Extrusions\nundoubtedly are the most susceptible form;\nforgings are probably second. Low-strength\nalloys and alclad sheet and plate are rela-\ntively immune to high-temperature oxida-\ntion. (Blistering of alclad material as a result\nof inadequate bonding is not the same as the\nblistering caused by high-temperature oxi-\ndation.)\nIf the protective oxide film formed during\nmill operations is removed from the mill\nproduct by a subsequent mechanical condi-\ntioning operation such as sanding, the con-\nditioned surface will be more susceptible to\nhigh-temperature oxidation than those from\nwhich the film was not removed.\nMoisture can be minimized by thoroughly\ndrying parts and racks before they are\ncharged. Drain holes often are needed in\nracks of tubular construction to avoid en-\n\n\nSee Table 1 for solution-treating temperatures.\n\n\nSoak time, minutes\n\n\nMaximum\nquench\ndelay, s\n\n\nAir furnace(b)\n\n\nSalt bath(c)\n\n\nThickness(a), mm (in.)\n\n\nmin\n\n\nmin\n\n\nmax(d)\n\n\nmax(d)\n\n\n20\n20\n25\n25\n30\n30\n30\n35\n35\n35\n40\n40\n50\n50\n55\n65\n222222222222222328\n+30\n60\n\n\n15\n20\n30\n30\n30\n55\n+20\n2822288333444445\n\n\n25\n30\n35\n35\n40\n40\n40\n45\n45\n45\n55\n55\n60\n60\n65\n75\n+30\n28339999972288678\n\n\n10\n10\n15\n15\n20\n20\n20\n25\n25\n25\n30\n30\n35\n35\n35\n45\n+20\n30\n22222222222222292\n\n\n\u22640.41 (0.016)\n\n\n7\n5\n7\n7\n10\n10\n10\n10\n10\n10\n15\n15\n15\n15\n15\n15\n(e)\n5\n\n\n0.51 (0.020)\n0.64 (0.025)\n0.81 (0.032)\n1.02 (0.040)\n1.27 (0.050)\n1.35 (0.053)\n1.80 (0.071)\n2.03 (0.080)\n2.29 (0.090)\n2.54 (0.100)\n3.18 (0.125)\n4.06 (0.160)\n4.57 (0.180)\n6.35 (0.250)\n\n\n>6.35 (0.250)-12.7 (0.500)\n\n\nFor each additional 12.7 (12) or fraction\nRivets (all)\n\n\n(a) Minimum dimension of thickest section. (b) Soak time begins when all pyrometer instruments recover to original operating\ntemperature. (c) Soak time begins at time of immersion except when a heavy charge causes bath temperature to drop below specified\nminimum, in which case soak time begins when bath regains minimum temperature. (d) Applicable to alclad materials only. (e) Increases\nin thickness above 12.7 mm (12 in.) do not affect maximum quench delay, which remains constant at 15 s.\n\n\nIn the tabulation above, note especially the\neffects of small increments of temperature,\nwithin the normal range, on the properties\nof 0.8 mm (0.032 in.) 2024-T4 sheet.\nSolution-Treating Time. The time at the\nnominal solution heat-treating temperature\n(soak time) required to effect a satisfactory\ndegree of solution of the undissolved or\nprecipitated soluble phase constituents and\nto achieve good homogeneity of the solid\nsolution is a function of microstructure be-\nfore heat treatment. This time requirement\ncan vary from less than a minute for thin\nsheet to as much as 20 h for large sand or\nplaster-mold castings. Guideline informa-\ntion for soak times required for wrought\nproducts of various section thicknesses is\ngiven in Table 2. Similar guidelines for\ncastings are presented in Table 3. The time\nrequired to heat a load to the treatment\ntemperature in furnace heat treatment also\nincreases with section thickness and fur-\nnace loading, and thus total cycle time\nincreases with these factors.\nSoak time for alclad sheet and for parts\nmade from alclad sheet must be held to a\nminimum, because excessive diffusion of\nalloying elements from the core into the\ncladding reduces corrosion protection. For\nthe same reason, reheat treatment of alclad\nsheet less than 0.75 mm (0.030 in.) thick\ngenerally is prohibited, and the number of\nreheat treatments permitted for thicker al-\nclad sheet is limited.\nThe soak times for wrought alloys take\ninto account the normal thermal lag be-\ntween furnace and part and the difference\nbetween surface and center temperatures\nfor commercial equipment qualified to the\nstandards of MIL-H-6088. The rapid heat-\ning rates of salt baths permit all immersion\n\n\ntime to be counted as soak time unless the\nbath temperature drops below the minimum\nof the range. Even then, soak time begins as\nsoon as the bath temperature returns to the\nminimum. In air furnaces, soak time does\nnot begin until all furnace instruments re-\nturn to their original set temperature-that\nis, the temperature reading before insertion\nof the load.\nIn air furnaces, thermocouples may also\nbe attracted to, or buried in, parts located in\nthe load in such a manner as to represent\nthe hottest and coldest temperatures in each\nzone. In this way, it is possible to ensure\nthat adequate soaking is obtained.\nSpecial consideration is given also to es-\ntablishing soak times for hand and die forg-\nings; soak time in some specifications is\nextended to complete solution and homog-\nenization in areas that received marginal\nreduction during forging. Considerable vari-\nation exists in the amount of soak time\nadded; some specifications call for an arbi-\ntrary addition, such as one hour, and others\nrequire one hour per inch of thickness of the\noriginal forging.\nIn air furnaces, careful attention should\nbe given to arrangement of the load. Air\nflow and natural temperature distribution\nwithin the furnace should be arranged to:\n\n\nOffer minimum resistance to air flow\nProduce the least disturbance in the natu-\nral temperature distribution\n\n\n\u2022\n\n\n\u2022\n\n\n\u2022 Afford constant replenishment of the en-\nvelope of air around each part\n\n\nIt is common practice to specify a minimum\nspacing of 50 mm (2 in.) between parts, but\nlarge complex shapes may require consider-\nably greater spacing. Many operators have\nfound conservative loading practices to be\n\n\nHeat Treating of Aluminum Alloys / 849\n\n\nTable 3 Typical heat treatments for aluminum alloy sand and permanent mold castings\n\n\nSolution heat treatment(b)\n\n\nAging treatment\n\n\nTemperature(c)\n\n\nTemperature(c)\n\n\nType of\ncasting(a)\n\n\nAlloy\n\n\nTime, h\n\n\nTemper\n\n\n\u00b0C\n\n\n\u00b0F\n\n\n\u00b0C\n\n\n\u00b0F\n\n\nTime, h\n\n\n201.0(d)\n\n\nS or P\n\n\n490-500(e)\n+525-530\n510-515(e)\n+525-530\n510-515(e)\n+525-530\n525\n490-500(e)\n+525-530\n530\n520\n530\n490-500(e)\n+525-530\n490-500(e)\n+525-530\n490-500(e)\n+525-530\n490-500(e)\n+525-530\n\n\nT4\n\n\n910-930(e)\n+980-990\n950-960(e)\n+980-990\n950-960(e)\n+980-990\n980\n910-930(e)\n\n\n2\n14-20\n2\n14-20\n2\n14-20\n20\n2\n14-20\n12\n10\n12\n2\n14-20\n2\n14-20\n2\n14-20\n2\n14-20\n\n\nMinimum of 5 days at room temperature\n\n\nS\n\n\nT6\n\n\n20\n\n\n155\n\n\n310\n\n\nT7\n\n\nS\n\n\n370\n\n\n5\n\n\n190\n\n\nT43(f)\n\n\n24 h at room temperature + \u00bd to 1 h at 160 \u00b0C\n\n\nT71\n\n\n+980-990\n\n\n390\nMinimum of 5 days at room temperature\n\n\n200\n\n\n4\n\n\nS or P\nS or P\nS or P\nS or P\n\n\n204.0(d)\n\n\nT4\nT4\n\n\n985\n970\n985\n910-930(e)\n+980-990\n910-930(e)\n+980-990\n910-930(e)\n\n\nT6(g)\n\n\n(g)\n\n\n(g)\n\n\n206.0(d)\n\n\nT4\n\n\nMinimum of 5 days at room temperature\n\n\nS or P\n\n\nT6\n\n\n12-24\n\n\n155\n\n\n310\n\n\nS or P\n\n\nT7\n\n\n+980-990\n\n\n200\n\n\n390\n\n\n4\n\n\nS or P\n\n\nT72\n\n\n910-930(e)\n\n\n+980-990\n\n\n470-480\n310\n600\n310\n340\n340\n650\n400\n330-340\n625-675\n400-450\n\n\n243-248\n155\n315\n155\n170\n170\n345\n205\n165-170\n330-355\n205-230\n\n\n208.0\n222.0\n\n\nSSSP\n\n\nT55\nO(h)\nT61\nT551\nT65\nO(i)\nT571\n\n\n16\n3\n11\n16-22\n7-9\n3\n8\n22-26\n2 (minimum)\n3-5\n\n\n510\n\n\n950\n\n\n12\n\n\n510\n\n\n950\n\n\n4-12\n\n\n242.0\n\n\nT77\nT61\nT4\nT6\nT62\nT7\nT4\nT6\nT7\nT5\nT6\n\n\nS\n\n\n515\n515\n515\n515\n515\n515\n510\n510\n510\n\n\n960\n960\n960\n960\n\n\n5(j)\n4-12(j)\n\u00a3sts\u00a3stts:E;g;sEs:\u0f68EsES\n12\n12\n12\n12\n8\n8\n8\n12\n4-12\n12\n6-12\n6-12\n8\n10-12\n12\n4-12\n4-12\n12\n4-12\n12\n4-12\n12\n6-12\n\n\nS or P\n\n\n295.0\n\n\n310\n310\n500\n\n\n3-6\n12-24\n46\n\n\n155\n155\n260\n\n\n960\n960\n950\n950\n950\n\n\n296.0\n\n\n...\n\n\n155\n260\n205\n155\n155\n155\n205\n205\n155\n260\n205\n205\n(h)\n225\n155\n155\n170\n225\n225\n245\n245\n155\n\n\n1-8\n46\n8\n2-5\n2-5\n2-5\n7-9\n7-9\n2-5\n46\n7-9\n7-9\n(1)\n7-9\n3-5\n2-5\n14-18\n3-5\n3-9\n46\n3-6\n3-5\n8 (minimum)\n10-12\n7-9\n3-5\n2-5\n3-5\n7-9\n3\n3-6\n3-5\n8 (minimum)\n6-12\n\n\n310\n500\n400\n310\n310\n310\n400\n400\n310\n500\n400\n400\n(h)\n440\n\n\n319.0\n\n\n505\n505\n515\n\n\n940\n940\n960\n\n\n328.0\n332.0\n333.0\n\n\nT6\nT5\nT5\nT6\nT7\nT551\nT65\n\n\nP\nP\nP\nP\nP\n(k)\n\n\n505\n505\n\n\n950\n940\n\n\n336.0\n\n\n515\n525-535\n\n\n960\n980-995\n\n\n354.0\n355.0\n\n\nS or P\n\n\nT51\nT6\n\n\nS\nP\n\n\n525\n525\n525\n525\n525\n525\n525\n525\n525\n\n\n980\n980\n980\n980\n980\n980\n980\n980\n980\n\n\n310\n310\n340\n440\n440\n475\n475\n310\n\n\nT62\nT7\n\n\nT71\n\n\nC355.0\n\n\nT6\nT61\n\n\nRoom temperature\n\n\n155\n225\n155\n155\n205\n225\n245\n245\n155\n\n\n310\n440\n310\n310\n400\n440\n475\n475\n310\n\n\n356.0\n\n\nSSPSPSesa\n\n\nS or P\n\n\nT51\nT6\n\n\n540\n540\n540\n540\n540\n540\n540\n540\n\n\n1000\n1000\n1000\n1000\n1000\n1000\n1000\n1000\n\n\n12\n4-12\n12\n4-12\n10-12\n4-12\n12\n6-12\n\n\nT7\n\n\nT71\n\n\nA356.0\n\n\nT6\nT61\n\n\nRoom temperature\n\n\n155\n\n\n310\n\n\n(continued)\n\n\n(a) S, sand; P, permanent mold. (b) Unless otherwise indicated, solution treating is followed by quenching in water at 65-100 \u00b0C (150-212 \u00b0F). (c) Except where ranges are given, listed temperatures are \u00b16 \u00b0C\nor \u00b110 \u00b0F. (d) Casting wall thickness, solidification rate, and grain refinement affect the solution heat-treatment cycle in alloys 201.0, 204.0, and 206.0, and care must be taken in approaching the final solution\ntemperature. Too rapid an approach can result in the occurrence of incipient melting. (e) For castings with thick or other slowly solidified sections, a pre-solution heat treatment ranging from about 490 to\n515 \u00b0C (910 to 960 \u00b0F) may be needed to avoid too rapid a temperature rise to the solution temperature and the melting of CuAl2. (f) Temper T43 for 201.0 was developed for improved impact resistance with\nsome decrease in other mechanical properties. Typical Charpy value is 20 J (15 ft lb). (g) The French precipitation treatment technology for the heat treatment of 204.0 alloy requires 12 h at temperature.\nThe aging temperatures of 140, 160, or 180 \u00b0C (285, 320, or 355 \u00b0F) are selected to meet the required combination of properties. (h) Stress relieve for dimensional stability as follows: hold 5 h at 413 \u00b1 14 \u00b0C\n(775 \u00b1 25 \u00b0F): furnace cool to 345 \u00b0C (650 \u00b0F) over a period of 2 h or more: furnace cool to 230 \u00b0C (450 \u00b0F) over a period of not more than 2 h; furnace cool to 120 \u00b0C (250 \u00b0F) over a period of approximately\n2 h; cool to room temperature in still air outside the furnace. (i) No quench required: cool in still air outside the furnace. (j) Air-blast quench from solution-treating temperature. (k) Casting process varies\n(sand, permanent mold, or composite) depending on desired mechanical properties. (1) Solution heat treat as indicated, then artificially age by heating uniformly at the temperature and for the time necessary\nto develop the desired mechanical properties. (m) Quench in water at 65-100 \u00b0C (150-212 \u00b0F) for 10-20 s only. (n) Cool to room temperature in still air outside the furnace.\n\n\n850 / Heat Treating of Nonferrous Alloys\n\n\nTable 3 (continued)\n\n\nSolution heat treatment(b)\n\n\nAging treatment\n\n\nTemperature(c)\n\n\nTemperature(c)\n\n\nType of\ncasting(a)\n\n\nAlloy\n\n\nTemper\n\n\nTime, h\n\n\n\u00b0C\n\n\n\u00b0F\n\n\nTime, h\n\n\n\u00b0C\n\n\n\u00b0F\n\n\nP\nS\nP\nS\nS\nandzannn\n(k)\n(k)\n\n\n357.0\n\n\nT6\nT61\n\n\n540\n540\n540\n540\n540\n430\n400\n\n\n1000\n1000\n1000\n1000\n1000\n810\n750\n\n\n8\n10-12\n8-12\n10-14\n8-12\n18(m)\n5\n\n\n175\n155\n(h)\n(h)\n\n\n350\n310\n(h)\n(h)\n\n\n6\n10-12\n(h)\n(h)\n\n\nA357.0\n\n\n359.0\n\n\nA444.0\n520.0\n535.0\n705.0\n\n\nT4\nT4\n\n\nT5(h)\n\n\nRoom temperature\n\n\n21 days\n8\n21 days\n10\n3-5\n21 days\n8\n4-10\n4-10\n21 days\n21 days\n21 days\n6-8\n21 days\n16\n4(n)\n3-5(n)\n6\n(h)\n3\n15\n7-9\n7-9\n4\n7-9\n\n\nT5\n\n\n210\n\n\n100\n\n\nRoom temperature\n\n\nP\n\n\n:\n\n\n100\n155\n\n\n210\n310\n\n\n707.0\n\n\nSP\n\u0420\n\n\nT5\n\n\nRoom temperature, or\n\n\n100\n175\n175\n\n\n210\n350\n350\n\n\nSPSPS\n\n\n8-16\n4-8\n\n\nT7\n\n\n530\n530\n\n\n990\n990\n\n\nRE\nT5\nT1\nT5\n\n\n710.0\n711.0\n712.0\n\n\nRoom temperature\nRoom temperature\nRoom temperature, or\n155\nRoom temperature, or\n120\n\n\n315\n\n\nT5\n\n\n713.0\n\n\nS or P\n\n\n250\n360(n)\n355(n)\n405\n(h)\n265\n285\n430\n430\n430\n430\n\n\n771.0\n\n\n415(n)\n\n\nT53(h)\n\n\n775(n)\n\n\n5(n)\n\n\nSSSSSSSS\n\n\n180(n)\n180(n)\n205\n(h)\n130\n140\n220\n220\n220\n220\n\n\nT5\nT51\nT52\nT6\nT71\nT5\nT5\nT6\nT5\n\n\n6(n)\n6(i)\n\n\n1090(n)\n1090(i)\n\n\n590(n)\n590(i)\n\n\nS or P\nS or P\nP\nS or P\n\n\n850.0\n851.0\n\n\n480\n\n\n900\n\n\n6\n\n\n852.0\n\n\n(a) S, sand; P, permanent mold. (b) Unless otherwise indicated, solution treating is followed by quenching in water at 65-100 \u00b0C (150-212 \u00b0F). (c) Except where ranges are given, listed temperatures are \u00b16 \u00b0C\nor \u00b110 \u00b0F. (d) Casting wall thickness, solidification rate, and grain refinement affect the solution heat-treatment cycle in alloys 201.0, 204.0, and 206.0, and care must be taken in approaching the final solution\ntemperature. Too rapid an approach can result in the occurrence of incipient melting. (e) For castings with thick or other slowly solidified sections, a pre-solution heat treatment ranging from about 490 to\n515 \u00b0C (910 to 960 \u00b0F) may be needed to avoid too rapid a temperature rise to the solution temperature and the melting of CuAl2. (f) Temper T43 for 201.0 was developed for improved impact resistance with\nsome decrease in other mechanical properties. Typical Charpy value is 20 J (15 ft lb). (g) The French precipitation treatment technology for the heat treatment of 204.0 alloy requires 12 h at temperature.\nThe aging temperatures of 140, 160, or 180 \u00b0C (285, 320, or 355 \u00b0F) are selected to meet the required combination of properties. (h) Stress relieve for dimensional stability as follows: hold 5 h at 413 \u00b1 14 \u00b0C\n(775 +25 \u00b0F); furnace cool to 345 \u00b0C (650 \u00b0F) over a period of 2 h or more; furnace cool to 230 \u00b0C (450 \u00b0F) over a period of not more than 2 h; furnace cool to 120 \u00b0C (250 \u00b0F) over a period of approximately\n2 h; cool to room temperature in still air outside the furnace. (i) No quench required; cool in still air outside the furnace. (j) Air-blast quench from solution-treating temperature. (k) Casting process varies\n(sand, permanent mold, or composite) depending on desired mechanical properties. (j) Solution heat treat as indicated, then artificially age by heating uniformly at the temperature and for the time necessary\nto develop the desired mechanical properties. (m) Quench in water at 65-100 \u00b0C (150-212 \u00b0F) for 10-20 s only. (n) Cool to room temperature in still air outside the furnace.\n\n\ntrapment of water. Another common re-\nquirement is adjustment of the position of\nthe quench tank with respect to furnace\ndoors and air intake. Because it is unlikely\nthat all moisture can be eliminated from the\natmosphere in a production heat-treating\nfurnace, it is extremely important to elimi-\nnate all traces of other contaminants from\nboth the parts and the furnace atmosphere.\nThe most virulent contaminants in attack-\ning aluminum are sulfur compounds. Resi-\ndues from forming or machining lubricants, or\nfrom a sulfur dioxide protective atmosphere\nused in prior heat treatment of magnesium,\nare potential sources of sulfur contamination.\nIn one plant, surface contamination resulted\nfrom sulfur-containing materials in tote boxes\nused to transport parts. In another, an epi-\ndemic of blistering was cured by rectifying a\n\"sour\" degreaser. In a third instance, it was\nfound that a vapor-degreasing operation was\nnot completely removing a thin, hard waxy\nresidue, and an alkaline cleaning operation\nwas added.\nVery often, the source of contamination\nis obscure and difficult to detect, and the\nproblem must be combated in another way.\nThe most common of the alternative meth-\n\n\nods is use of a protective fluoborate com-\npound in the furnace. Such a compound\nusually is effective in minimizing the harm-\nful effects of moisture and other undesirable\ncontaminants because it forms a barrier\nlayer or film on the aluminum surface. The\nadditive is not a universal solution; in some\napplications, high-temperature oxidation\nhas occurred even though a fluoborate com-\npound was employed. Also, the use of such\ncompounds, particularly ammonium fluo-\nborate, may present a hazard to personnel if\nused in poorly sealed furnaces or in furnac-\nes that discharge their atmospheres into\nenclosed areas.\nProtective fluoborate compounds accen-\ntuate staining or darkening of the parts\nbeing treated. (At times, this attack, partic-\nularly on parts located near the protective-\ncompound container during heat treatment,\nhas been severe enough to be termed \"cor-\nrosion.\") Although this minor nuisance\nmight be considered a small price to pay for\nsolution of a problem of high-temperature\noxidation, the residual compound in the\nfurnace dissipates slowly. Therefore, subse-\nquent loads of alloys and product forms\nwhose end uses require bright surfaces, and\n\n\nthat are not susceptible to high-temperature\noxidation, may be detrimentally affected.\nSuccessful use of fluoborate protective\ncompounds appears to depend on specify-\ning the right amount for each furnace; this\nmust be established on a trial-and-error\nbasis. One aircraft manufacturer adds 4\ng/m\u00b3 (0.004 oz/ft\u00b3) of furnace chamber to\neach load. Another adds 0.45 kg (1 lb) per\nshift to a metal container hung on the fur-\nnace chamber wall, thus avoiding loss of the\ncompound during quenching.\nA second method of combating high-tem-\nperature oxidation is to anodize the work\nbefore it is heat treated. The resultant alu-\nminum oxide film prevents attack by con-\ntaminants in the furnace atmosphere. The\nonly deterrents to the use of anodizing are\nits cost (in money and time) and the slight\nsurface frostiness which results from the\nsubsequent stripping operation.\nThe usual objection to the blistered sur-\nface produced by high-temperature oxida-\ntion is its unsightly appearance. This often\ncan be improved (for salvage purposes) by\napplying local pressure to flatten each blis-\nter and then finishing by a mechanical pro-\ncess such as polishing, buffing, sanding, or\n\n\nHeat Treating of Aluminum Alloys / 851\n\n\nabrasive blasting. In general, the effect of\nHTO on static properties and fatigue\nstrength is slight. However, if a void result-\ning from HTO is located close to another\nstress concentration, such as a hole, much\ngreater degradation of fatigue strength is\nlikely. In critical aluminum alloy forgings,\nany blistering must be evaluated carefully\nfor its effect on the integrity of the part. Any\n\"cosmetic\" salvage should be performed\nonly after it has been established that the\nblisters are superficial and will not remain in\nthe finished product.\nPrecipitation Heat Treating without Prior\nSolution Heat Treatment. Certain alloys that\nare relatively insensitive to cooling rate\nduring quenching can be either air cooled or\nwater quenched directly from a final hot-\nworking operation. In either condition,\nthese alloys respond strongly to precipita-\ntion heat treatment. This practice is widely\nused in producing thin extruded shapes of\nalloys 6061, 6063, 6463, and 7005. Upon\nprecipitation heat treating after quenching\nat the extrusion press, these alloys develop\nstrengths nearly equal to those obtained by\nadding a separate solution heat treating op-\neration. Changes in properties occurring\nduring the precipitation treatment follow\nthe principles outlined in the discussion of\nsolution heat-treated alloys.\n\n\nprimary fabricating mills, by progressive\nflooding or high-velocity spraying with cold\nwater. However, parts of complex shape,\noften with both thin and thick sections (such\nas die forgings, most castings, impact extru-\nsions, and components formed from sheet)\nare commonly quenched in a medium that\nprovides somewhat slower cooling. This\nmedium may be water at 65 to 80 \u00b0C (150 to\n180 \u00b0F), boiling water, an aqueous solution\nof polyalkylene glycol, or some other fluid\nmedium such as forced air or mist.\nIf appreciable precipitation during cool-\ning is to be avoided, two requirements must\nbe satisfied. First, the time required for\ntransfer of the load from the furnace to the\nquenching medium must be short enough to\npreclude slow precooling into the tempera-\nture range where very rapid precipitation\ntakes place. For alloy 7075, this range was\ndetermined to be 400 to 290 \u00b0C (750 to 550\n\u00b0F), and some sources quote this range (or a\nslightly different range) as the most critical\nrange for quenching of any aluminum alloy.\nLater work has shown that the most critical\nrange is alloy-dependent, and as will be\ndiscussed in detail under \"Quench-Factor\nAnalysis,\u201d significant errors can result from\nthe assumption that precipitation is negligi-\nble outside of a so-called \"critical range.'\nThe second requirement for avoidance of\nappreciable precipitation during quenching\nis that the volume, heat-absorption capaci-\nty, and rate of flow of the quenching medi-\num be such that little or no precipitation\noccurs during cooling. Any interruption of\nthe quench that might allow reheating into a\ntemperature range where rapid precipitation\ncan occur must be prohibited.\nFor maximum dimensional stability,\nsome forgings and castings are fan cooled or\nstill-air cooled. In such instances, precipita-\ntion-hardening response is limited, but sat-\nisfactory values of strength and hardness\nare obtained. Extrusions produced without\nseparate solution heat treatment can be air\nor mist quenched, but thicker sections may\nrequire water quenching by immersion or\nspraying. Alloys that are relatively dilute,\nsuch as 6063 and 7005, are particularly well\nsuited to air quenching, and their mechani-\ncal properties are not greatly affected by its\nlow cooling rate. Lower quenching rates are\nalso employed for forgings, castings, and\ncomplex shapes to minimize warpage or\nother distortion and the magnitude of resid-\nual stresses developed as a consequence of\ntemperature nonuniformity from surface to\ninterior.\nEffect of Quench Rate on Properties. As a\nbroad generalization, the highest strengths\nattainable and the best combinations of\nstrength and toughness are those associated\nwith the most rapid quenching rates. Resis-\ntance to corrosion and stress-corrosion\ncracking are other characteristics that are\ngenerally improved by maximum rapidity of\nquenching. Some of the alloys used in arti-\n\n\nficially aged tempers, and in particular the\ncopper-free 7xxx alloys, are exceptions to\nthis rule. The effect of quench rate on\nmechanical properties may also depend on\nthe desired temper. In the underaged con-\ndition, for example, a slow quench rate is\nmore detrimental on ductility and fracture\ntoughness. Strength would be more affected\nafter near-to-peak aging.\nBecause of these effects, much work has\nbeen done over the years to understand and\npredict how quenching conditions and prod-\nuct form influence properties. The relative\neffects of quench methods can be compared\nin terms of average quench rates. In Fig 3,\nfor example, the effects of quenching on the\nyield strength of four alloys are compared in\nterms of average quenching rates through\nthe range from 400 to 290 \u00b0C (750 to 550 \u00b0F).\nFor alloys relatively high in sensitivity to\nquenching rate, such as 7075, rates of about\n300 \u00b0C/s (540 \u00b0F/s) or higher are required in\norder to obtain near-maximum strength af-\nter precipitation heat treatment. The other\nalloys in Fig 3 maintain their strengths at\ncooling rates as low as about 100 \u00b0C/s (180\n\u00b0F/s). Similar comparisons in terms of aver-\nage quench rates are shown in Tables 4 and\n5.\nAverage quench rates are useful in com-\nparing experimental results from various\nquench methods. In Table 4, for example, a\nsevere reduction in strength occurred at the\naverage quench rate of 36 \u00b0C/s (65 \u00b0F/s).\nHowever, average quench rates only com-\npare results in a \"critical\" temperature\nrange, where precipitation is most likely to\noccur. This method is not entirely accurate,\nbecause significant precipitation can also\noccur outside the specified critical temper-\nature range of average quench rates. More-\nover, for high-strength alloys, toughness\nand corrosion resistance may be impaired\nwithout significant loss of tensile strength.\nTherefore, a more sophisticated compar-\nison, known as quench-factor analysis, is\nneeded for quantitative property prediction\nor property optimization. Quench-factor\nanalysis, as discussed in a later section, is\nuseful when cooling rates are nonuniform.\nDelay in Quenching. Whether the transfer\nof parts from the furnace to the quench is\nperformed manually or mechanically, it\nmust be completed in less than the specified\nmaximum time. The maximum allowable\ntransfer time or \u201cquench delay\" varies with\nthe temperature and velocity of the ambient\nair and the mass and emissivity of the parts.\nFrom cooling curves such as those illustrat-\ned in Fig 4, maximum quench delays (see\ntable accompanying Fig 4) can be deter-\nmined that will ensure complete immersion\nbefore the parts cool below 400 \u00b0C (750 \u00b0F).\nMIL-H-6088 specifies maximum quench de-\nlays for high-strength alloys of 5, 7, 10, and\n15 s for thickness ranges of up to 0.016 in.\n(0.41 mm), 0.017 to 0.031 in. (0.43 to 0.79\nmm), 0.032 to 0.090 in. (0.81 to 2.29 mm),\n\n\nQuenching\n\n\nQuenching is in many ways the most\ncritical step in the sequence of heat-treating\noperations. The objective of quenching is to\npreserve the solid solution formed at the\nsolution heat-treating temperature, by rap-\nidly cooling to some lower temperature,\nusually near room temperature. From the\npreceding general discussion, this statement\napplies not only to retaining solute atoms in\nsolution, but also to maintaining a certain\nminimum number of vacant lattice sites to\nassist in promoting the low-temperature dif-\nfusion required for zone formation. The\nsolute atoms that precipitate either on grain\nboundaries, dispersoids, or other particles,\nas well as the vacancies that migrate (with\nextreme rapidity) to disordered regions, are\nirretrievably lost for practical purposes and\nfail to contribute to the subsequent\nstrengthening.\nIn most instances, to avoid those types of\nprecipitation that are detrimental to me-\nchanical properties or to corrosion resis-\ntance, the solid solution formed during so-\nlution heat treatment must be quenched\nrapidly enough (and without interruption) to\nproduce a supersaturated solution at room\ntemperature the optimum condition for\nprecipitation hardening. The resistance to\nstress-corrosion cracking of certain copper-\nfree aluminum-zinc-magnesium alloys,\nhowever, is improved by slow quenching.\nMost frequently, parts are quenched by\nimmersion in cold water or, in continuous\nheat treating of sheet, plate, or extrusions in\n\n\n852 / Heat Treating of Nonferrous Alloys\n\n\nAverage cooling rate\nfrom 750-550 \u00b0F, \u00b0F/s\n\n\nstopwatch or, if necessary, by attaching\nthermocouples to parts. However, although\nthe cooling rate between 400 and 260 \u00b0C (750\nand 500 \u00b0F) is most critical and must be\nextremely high for many high-strength al-\nloys, it cannot be directly measured in pro-\nduction operations. It is usual to rely on\nstandardized practices, augmented by re-\nsults of tension tests and tests of suscepti-\nbility to intergranular corrosion.\nWater-immersion quenching normally is\ncontrolled in practice by stipulating maxi-\nmum quench-delay time and maximum wa-\nter temperature. The first requirement con-\ntrols the cooling rate during transfer and,\nfor high-strength alloys, often is based on\nthe criterion of complete immersion before\nthe metal cools below 415 \u00b0C (775 \u00b0F). This\nspecification of 415 \u00b0C (775 \u00b0F) is based on a\ncritical temperature for alloy 7075, which\nhas one of the more severe C-curves (Fig 5).\nTherefore, the criterion for complete im-\nmersion of other alloys might be based on a\ntemperature lower than the 415 \u00b0C (775 \u00b0F)\nspecification, depending on the characteris-\ntics of the particular C-curve.\nThe second requirement controls the\ncooling rate during immersion. MIL-H-6088\nspecifies that for water-immersion quench-\ning, except quenching of forgings and cast-\nings, the temperature of the water shall not\nexceed 38 \u00b0C (100 \u00b0F) upon completion of\nquenching. This requirement controls both\nthe temperature of the quench water prior\nto immersion and the ratio of the combined\nmass of load and rack to the volume of\nwater. However, to ensure adequate\nquenching effectiveness, it is necessary also\nthat the cooling fluid flow past all surfaces\nof each part during the first few seconds\nafter immersion. Before parts enter the fur-\nnace, their placement in racks or baskets\nshould be compatible with this requirement.\nDuring the first few seconds of quenching,\nagitation of the parts or the water should be\nsufficient to prevent local increases in tem-\nperature due to the formation of steam\npockets.\nIn one application, it was found that\n2024-T4 plates 13 by 760 by 760 mm (1/2 by\n30 by 30 in.), quenched singly into a large\nvolume of still water, were quite susceptible\nto intergranular corrosion. This susceptibil-\nity disappeared completely when the\nquenching practice was modified by adding\nsufficient agitation to break up the insulat-\ning blanket of steam that formed on the\nsurface of the hot metal. Quenching prac-\ntices for small parts such as fasteners and\nhydraulic fittings have been modified for the\nsame reason. Dumping in bulk from baskets\nhas been replaced by methods, such as the\nuse of shaker hearth furnaces or special\nracking, which permit parts to be quenched\nsingly.\nSpray Quenching. For spray quenching,\nthe quench rate is controlled by the velocity\nof the water and by volume of water per unit\n\n\nAverage quenching rate\nfrom 750-550 \u00b0F, \u00b0F/s\n\n\n102 103 104 105\n\n\n10\n\n\n102\n\n\n103\n\n\n104\n\n\n10\n\n\n700\n\n\nYield strength, MPa\n\n\n100\n\n\n600\n\n\nYield strength, ksi\n\n\n80\n\n\nTensile strength, MPa\n\n\n7050-T6\n\n\n7178-T6\n7075-T6\n7050-T73\n7075-T73\n2014-T6\n2024-T4\n\n\n500 -7050-T736\n\n\n600\n\n\nTensile strength, ksi\n\n\n60\n\n\n400\n\n\n80\n\n\n7050-T73\n\n\n500\n\n\n\u30ad\u30fc\u30ad\n-2024-T4\n\n\n300\n\n\n40\n\n\n200\n\n\n60\n\n\n400\n\n\n6061-T6\n\n\n6070-T6\n1\n6061-T6\n\n\n20\n20\n\n\n100\n\n\n103\n\n\n102\n\n\n104\n\n\n1\n\n\n10\n\n\n300\n\n\n40\n\n\nAverage quenching rate\nfrom 400-290 \u00b0C, \u00b0C/s\n\n\n200\n\n\n102 103 104 105\n\n\n1\n\n\n10\n\n\n(a)\n\n\nAverage cooling rate\nfrom 400-290 \u00b0C, \u00b0C/s\n\n\n(b)\n\n\n100\n\n\nFraction of WQ yield strength, %\n\n\nFurnace cooling\n\n\n80\n\n\nForced air cooling\n\n\n60\n50\n\n\nAir cooled\n\n\n40\n\n\nAlloy and condition (Source: Ref 1)\n\n\nO 8090, peak aged\n\u26ab 2090, peak aged\n\n\n20\n20\n\n\n\u25b2 7150, aged 24 h at 120 \u00b0C\n\u25b2 7475, aged 24 h at 120 \u00b0C\n\u043d\u043d\n\n\n0\n0.01\n\n\n0.1\n\n\n1\n\n\n10\n\n\nWQ\n\n\nAverage cooling rate, \u00b0C/s\n\n\n(c)\n\n\nFig 3 Quench sensitivity of various aluminum alloys as a function of average quench rates. (a) Yield strength\nafter aging of four wrought alloys. (b) Tensile strength after aging of eight wrought alloys. (c) Relative\nquench sensitivity of two aluminum-lithium alloys (2090 and 8090, both solution treated for 1 h at 520 \u00b0C, or 970\n\u00b0F) and two Zn-Mg-Cu aluminum alloys (7150 and 7475, both solution treated for 40 min at 480 \u00b0C, or 895 \u00b0F)\n\n\nand over 0.090 in., respectively. Quench de-\nlay is conservatively defined as commencing\n\"when the furnace door begins to open or the\nfirst corner of a load emerges from a salt\nbath\" and ending \"when the last corner of the\nload is immersed in the water quench tank.\"\nRecommended maximum quench-delay times\nare listed in Table 2. However, exceeding the\n\n\nmaximum delay time is permitted if tempera-\nture measurements of the load prove that all\nparts are above 415 \u00b0C (775 \u00b0F) when\nquenched. The C-curves used in quench-fac-\ntor analysis can also assist in determining a\nmaximum allowable delay.\nIt is relatively easy to control quench\ndelay in day-to-day operations by using a\n\n\nTable 4 Effect of average quench rate on tensile properties of aluminum-lithium alloy\n2090\n\n\nTensile\nstrength(a)\n\n\nYield\nstrength(a)\n\n\nAverage quench rate at center of plate\n0.5 \u00b0C/s (13 mm plate, air\ncooled)\n\n\nElongation(a), %\n\n\nCondition\n\n\nMPa\n\n\nksi\n\n\nMPa\n\n\nksi\n\n\nAs-quenched\n\n\n334\n\n\n162\n\n\n2\n\n\n6% stretch + aged\n\n\n513\n\n\n448\n\n\n5\n\n\n8 h at 190 \u00b0C\n\n\n36 \u00b0C/s (38 mm plate, quenched\nin room-temperature water)\n\n\n128\n\n\nAs-quenched\n\n\n312\n\n\n12\n\n\n6% stretch + aged\n\n\n338\n\n\n476\n\n\n6\n\n\n8 h at 190 \u00b0C\n\n\n46 \u00b0C/s (13 mm plate, quenched\nin boiling water)\n\n\nAs-quenched\n\n\n331\n\n\n138\n\n\n16\n\n\n6% stretch aged\n\n\n530\n\n\n570\n\n\n9\n\n\n8 h at 190 \u00b0C\n\n\n48 \u00b0C/s (13 mm plate, quenched\nin room-temperature water)\n\n\n139\n\n\n331\n\n\nAs-quenched\n\n\n17\n\n\n6% stretch aged\n\n\n526\n\n\n570\n\n\n7\n\n\n8 h at 190 \u00b0C\n\n\n85 \u00b0C/s (13 mm plate, quenched\nin ice brine)\n\n\nAs-quenched\n\n\n349\n\n\n135\n\n\n19\n\n\n6% stretch aged\n\n\n8 h at 190 \u00b0C\n\n\n575\n\n\n535\n\n\n7\n\n\n(a) Data are averages from 4 specimens.\n\n\nHeat Treating of Aluminum Alloys / 853\n\n\nTable 5 The effect of quench rate on the mechanical properties of age-hardened aluminum-lithium alloy 8090\n\n\nUltimate tensile\nstrength(b)\n\n\nYield strength(b)\n\n\nCooling from\nsolution treatment(a)\n\n\nElongation in 50 mm\n(2 in.)(b), %\n\n\nStretch, %\n\n\nAging treatment\n\n\nAlloy composition\nAl-2.28Li-0.86Cu-\n0.90Mg-0.13Zr-\n0.13Fe-0.06Si\n\n\nMPa\n\n\nksi\n\n\nksi\n\n\nMPa\n\n\nAir cool\n\n\n190 \u00b0C for 16 h\n\n\n64.5\n\n\n380\n\n\n7.7\n\n\n2\n\n\n55\n\n\n446\n\n\n(-0.25 \u00b0C/s)\n\n\n170 \u00b0C for 24 h\n190 \u00b0C for 16 h\n\n\n67.5\n70\n\n\n4\n2\n\n\n401\n415\n\n\n88 28 80\n58\n60\n60\n62\n60\n60\n\n\n6.0\n8.0\n\n\n465\n481\n\n\nPolymer quench\n(-18 \u00b0C/s)\n\n\n170 \u00b0C for 24 h\n190 \u00b0C for 16 h\n\n\n7.2\n8.1\n\n\n4\n2\n\n\n415\n428\n\n\n481\n492\n\n\n70\n71.4\n\n\nWater quench\n(~120 \u00b0C)\n\n\n170 \u00b0C for 24 h\n190 \u00b0C for 16 h\n\n\n7.5\n6.5\n\n\n4\n2\n\n\n417\n417\n\n\n483\n485\n\n\n70\n70.3\n\n\nAir cool\n\n\nAl-2.58Li-1.36Cu-\n0.89Mg-0.13Zr-\n0.17Fe-0.04Si\n\n\n(-0.25 \u00b0C/s)\n\n\n170 \u00b0C for 24 h\n190 \u00b0C for 16 h\n\n\n4\n2\n\n\n2 22 28\n64\n65\n65\n67\n65\n\n\n73\n76\n\n\n4.5\n5.0\n23 24\n6.8\n8.2\n\n\n442\n\n\n503\n524\n\n\nPolymer quench\n(-18 \u00b0C/s)\n\n\n448\n\n\n170 \u00b0C for 24 h\n190 \u00b0C for 16 h\n\n\n448\n464\n\n\n519\n535\n\n\n75\n\n\nWater quench\n(~120 \u00b0C/s)\n\n\n2\n\n\n77.5\n\n\n170 \u00b0C for 24 h\n\n\n6.3\n\n\n448\n\n\n517\n\n\n75\n\n\n(a) Solution treatment of 550 \u00b0C (1020 \u00b0F) for 1 h. (b) Data are averages from two specimens.\n\n\narea per unit time of impingement of the\nwater on the workpiece. Rate of travel of\nthe workpiece through the sprays is an\nimportant variable.\nLocal increases in temperature that occur\nwithin the first few seconds of quenching,\ncaused by a phenomenon such as plugged\nspray nozzles, are particularly deleterious.\nThe remaining \u201cinternal heat\u201d may be suf-\nficient to reheat the surface region. When\nthis happens, a large loss in strength occurs\nat the previously quenched surface. The\nloss of strength in the affected area of a\nheavy part is much more severe than that\ncaused by an inadequate quenching rate\nalone. This is illustrated for 75 mm (3 in.)\nthick 7075-T62 plate in Fig 6, which com-\npares, at various depths, the properties of a\nplate for which quenching was interrupted\non one side after 3 s with those of a plate\nthat was quenched from one side only.\nQuench Severity and Quenchant Selec-\ntion. Quench severity is commonly ex-\npressed in terms of an H-value (or Gross-\nmann number), where the H-value is related\nto the thermal conductivity (k) of the part(s)\nand the coefficient of heat transfer (C) be-\n\n\ntween the quenchant and the part. These\nquantities are related by the equation H =\nC/2k, where the coefficient of heat transfer\n(C) is affected by the quenchant velocity at\nthe surface of the part and several inherent\ncharacteristics of the quenchant (such as\nquenchant boiling point, viscosity, density,\nthermal conductivity, and specific heat).\nWater, which is the most widely used and\neffective quenching medium, can obtain\ncooling rates up to about 200 \u00b0C/s (400 \u00b0F/s)\nat the midplane of 25 mm (1 in.) thick\naluminum alloy plate (see the dashed line in\nFig 7). No rates higher than those defined\nby this line have been observed, although\nrates approaching them were measured with\nimpinging spray quenches. Lower cooling\nrates are achieved by immersion in heated\nwater (Fig 7) or by reducing the velocity of\nthe quenchant around the part (Table 6).\nCooling rates can also be reduced by low-\nering surface tension or by increasing the\nstability of the vapor film around the part.\nPolymer quenchants, which retard cool-\ning rates by the formation of films around\nthe part, are compared with water in Table\n6. The effective film coefficient is essential-\n\n\nly the heat transfer coefficient (C), which is\nrelated to the Grossmann number (H). The\napplication of polymer quenchants is cov-\nered in AMS specifications 3025 and 2770,\nalthough many aluminum and aerospace\ncompanies have developed internal specifi-\ncations that differ from AMS-2770. Typical\nparameters for quenching wrought products\n(other than forgings) in glycol-water solu-\ntions are presented in Table 7.\nOther Factors Affecting Quench Rate.\nQuenching rates are very sensitive to the\nsurface condition of the parts. Lowest rates\nare observed with products having freshly\nmachined or bright-etched, clean surfaces,\nor products that have been coated with\nmaterials that decrease heat transfer. The\npresence of oxide films or stains increases\ncooling rates. Further marked changes can\nbe effected through the application of non-\nreflective coatings, which also accelerate\nheating (Fig 8). Surface roughness exerts a\nsimilar effect; this appears related to vapor\nfilm stability. The manner in which complex\nproducts, such as engineered castings and\ndie forgings, enter the quenching medium\ncan significantly alter the relative cooling\n\n\nTime per unit thickness, s/in\n\n\n0\n\n\n2000\n\n\n3000\n\n\n1000\n\n\n600\n\n\n1000\n\n\nMaximum quench delay, s\n\n\nThickness\n\n\nTemperature, \u00b0C\n\n\nTemperature, \u00b0F\n\n\n800\n\n\nAlclad\n\n\nNonclad\n\n\nin.\n\n\nmm\n\n\n400\n\n\nAlclad\n\n\n0.016\n0.020\n0.025\n0.032\n0.040\n\n\n0.41\n0.51\n0.64\n0.81\n1.02\n\n\n4.4\n5.5\n6.8\n8.8\n11.0\n\n\n6.4\n8.0\n10.0\n12.8\n20.0\n\n\n600\n\n\n400\n\n\n200\n\n\nNonclad\n\n\n200\n\n\n0\n\n\n100\n\n\n150\n\n\n0\n\n\n50\n\n\nTime per unit thickness, s/mm\n\n\nFig 4 Cooling curves for alclad and nonclad aluminum products cooled from 495 \u00b0C (920 \u00b0F) in forced air. Air temperature, 25 \u00b0C (80 \u00b0F); air velocity, 2.3 m/s (450 ft/min).\n\n\nTabulated values of quench delay (maximum delay before the material being quenched has cooled below 400 \u00b0C, or 750 \u00b0F) were determined from cooling curves\n\n\nshown.\n\n\n854 / Heat Treating of Nonferrous Alloys\n\n\n600\n\n\nflow between parts, jets should not impinge\ndirectly and cause rapid localized cooling.\nQuenching to Minimize Residual Stress\nand Warpage. Although cold-water immer-\nsion or flushing is most common, because it\nproduces the most effective quench (and\nhas been required by MIL-H-6088 for 2014,\n2017, 2024, 2117, 7075, and 7178 alloys\nexcept forgings), it presents problems in-\nvolving residual stress and warpage.\nResidual stresses in heavy sections of\naluminum alloys originate from differential\nthermal expansion during quenching \u2015that\nis, the still-warm central material contracts,\npulling in the already cooled outer shell.\nThe magnitude of stresses increases with\nsection size, as shown in Fig 9.\nThe distribution pattern of residual\nstresses in as-quenched parts (compression\nin the outer layers and tension in the central\nportion) is usually desirable in service.\nCompressive stresses inhibit failure by fa-\ntigue and stress corrosion-two mecha-\nnisms that initiate in the outer fibers. Un-\nfortunately, metal-removal operations\nrequired after heat treating often expose\nmaterial that is stressed in tension. Also,\nmetal-removal operations that are asym-\nmetrical (with respect to residual stresses)\ncause distortion by redistributing residual\nstresses. When close-tolerance parts are\nbeing fabricated, the resulting warpage can\nbe costly and difficult to correct.\nAlthough service performance is some-\ntimes a factor, the major incentive for re-\nducing residual stress differentials has been\na reduction in warpage during machining or\nan improvement in shape before machining.\nOne approach to reducing the cooling-\nrate differential between surface and center\nis the use of a milder quenching medium-\nwater that is hotter than that normally used\nor water-glycol solutions. Boiling water,\nwhich is the slowest quenching medium\nused for thick sections, is sometimes em-\nployed for quenching wrought products\neven though it lowers mechanical properties\nand corrosion resistance. Quenching of\ncastings in boiling water, however, is stan-\ndard practice, and is reflected in design\nallowables.\nAnother approach to the minimization of\nresidual stresses that is generally successful\nconsists of rough machining to within 3.2\nmm (0.125 in.) or less of finish dimensions,\nheat treating, and then finish machining.\nThis procedure is intended to reduce the\ncooling-rate differential between surface\nand center by reducing thickness; other\nbenefits that accrue if this technique is used\nto reduce or reverse surface tension stress-\nes in finished parts are improvements in\nstrength, fatigue life, corrosion resistance,\nand reduced probability of stress-corrosion\ncracking.\nSeveral factors (especially quenching\nwarpage) sometimes preclude general use of\nthis procedure. The thinner and less sym-\n\n\n1110\n\n\n500\n\n\n930\n\n\n750\n\n\n400\n\n\nTemperature, \u00b0C\n\n\nTemperature, \u00b0F\n\n\nA\n\n\n570\n\n\n300\n\n\n200\n\n\n390\n\n\nA: 7075\nB: 2017\nC: 6061\nD: 6063\n\n\n100\n\n\n212\n\n\n32\n\n\n0\n\n\n1\n\n\n103\n\n\n10\n\n\n100\n\n\nTime, s\n\n\nTime-temperature-property curves at 95% of maximum tensile stress for various alloys. See the section\n\"Quench-Factor Analysis\" for discussion. Source: Ref 2\n\n\nFig 5\n\n\nrates at various points, thereby affecting\nmechanical properties and residual stresses\nestablished during quenching. Similarly,\nquenching complex extruded shapes whose\nwall thicknesses differ widely poses special\nproblems if distortion and stresses are to be\nminimized. In batch heat-treating opera-\n\n\ntions, placement and spacing of parts on the\nracks can be a major factor in determining\nthe quenching rates. In immersion quench-\ning, adequate volumes of the quenching\nmedium must be provided to prevent an\nexcessive temperature rise in the medium.\nWhen jet agitation is used to induce water\n\n\nTable 6 Grossmann numbers and heat transfer coefficients (C) of quenchant-to-part films\n\n\nQuenchant\n\n\nEffective film heat transfer\ncoefficient (C)\n\n\nGrossmann\nNumber\n(H = C/2k)\n\n\nVelocity\n\n\nTemperature\n\n\nBtu/ft2 h \u00b0F\n\n\n\u0422\u0443\u0440\u0435\n\n\nW/cm\u00b2. K\n\n\n\u00b0C\n\n\n\u00b0F\n\n\nm/s\n\n\nft/min\n\n\n27\n\n\nWater\n\n\n80\n\n\n3.55\n4.78\n5.14\n3.28\n4.01\n4.91\n3.65\n4.29\n5.31\n2.85\n3.62\n4.41\n0.70\n1.89\n2.62\n0.36\n0.69\n0.89\n0.20\n0.27\n0.30\n0.13\n0.13\n0.13\n\n\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n0.00\n0.25\n0.50\n\n\n0\n50\n100\n\n\n2460\n3105\n3565\n2275\n2785\n3400\n2530\n2970\n3680\n1980\n2510\n3060\n485\n1310\n1815\n255\n485\n620\n138\n207\n92\n92\n184\n92\n\n\n1.07\n1.35\n1.55\n0.99\n1.21\n1.48\n1.10\n1.29\n1.60\n0.86\n1.09\n1.33\n0.21\n0.57\n0.79\n0.11\n0.21\n0.27\n0.06\n0.08\n0.09\n0.04\n0.04\n0.04\n\n\nWater\n\n\n38\n\n\n100\n\n\n0\n50\n100\n0\n50\n100\n\n\nWater\n\n\n49\n\n\n120\n\n\nWater\n\n\n60\n\n\n140\n\n\n0\n50\n100\n\n\nWater\n\n\n71\n\n\n160\n\n\n0\n50\n100\n\n\n180\n80\n\n\nWater\n\n\n82\n\n\n0\n50\n100\n\n\nWater\n\n\n93\n\n\n200\n\n\n0\n50\n100\n50\n100\n88088\n\n\n212\n\n\nWater\n\n\n100\n\n\nPolyalkylene glycol\n(UCON A)(a)\n\n\n50\n100\n088\n\n\n30\n\n\n85\n\n\n0.00\n0.25\n0.50\n\n\n0.63\n0.70\n0.77\n\n\n0.19\n0.21\n0.23\n\n\n429\n475\n529\n\n\nPolyvinyl\n\n\npyrrolidone\n(PVP90)(a)\n\n\n30\n\n\n0.00\n0.25\n0.50\n\n\n1.49\n1.34\n1.41\n\n\n85\n\n\n0\n50\n100\n\n\n0.44\n0.40\n0.42\n\n\n1012\n912\n966\n\n\n(a) Polymer quenchants with concentrations of 25%. K is equal to the thermal conductivity of the aluminum alloy (7075). Source: Ref 4\n\n\nHeat Treating of Aluminum Alloys / 855\n\n\nDepth, in.\n\n\nBecause of the difficulties encountered\nwith quenching in cold water, milder quen-\nchants have been employed. Indiscriminate\nuse of milder quenchants can have cata-\nstrophic effects; however, when their use is\nbased on sound engineering judgment and a\nmetallurgical knowledge of the effects on\nthe specific alloy, significant cost savings or\nperformance improvements can be realized.\nThe most frequent advantage is the re-\nduction in costly straightening operations\nand in resultant uncontrolled residual\nstresses. For example, one aircraft manu-\nfacturer utilizes water-spray and air-blast\nquenching for weldments and complex\nformed parts made from 6061, an alloy\nwhose corrosion resistance is insensitive to\nquenching rate. Straightening requirements\nare negligible and, through careful control\nof racking and coolant flow, the decrease in\nmechanical properties is minimized, as\nshown by the data in Fig 10.\nAnother development for reducing straight-\nening costs is quenching in water-polymer\nsolutions. Quenching of formed sheet-metal\nparts in aqueous solutions of polyalkylene\nglycol or in similar inversely soluble media\nhas significantly reduced the cost of straight-\nening these parts after quenching. The SAE\nheat-treatment specification AMS-2770 rec-\nommends, for several alloys, maximum thick-\nnesses that can be quenched in solutions of\nspecific concentrations while maintaining ac-\nceptable property levels. Typical parameters\nfor quenching wrought products (other than\nforgings) in glycol-water solutions are pre-\nsented in Table 7. Additional information on\npolymer quenchants for aluminum alloys can\nbe found in Ref 5.\nForming and Straightening after Quench-\ning. Immediately after being quenched,\nmost aluminum alloys are nearly as ductile\nas they are in the annealed condition. Con-\nsequently, it is often advantageous to form\nor straighten parts in this temper. More-\nover, at the mill level, controlled mechani-\ncal deformation is the most common meth-\nod of reducing residual quenching stresses.\nBecause precipitation hardening will occur\nat room temperature, forming or straighten-\ning usually follows as soon after quenching\nas possible. In addition, maximum effec-\ntiveness in stress relief is obtained by work-\ning the metal immediately after quenching.\nForming and straightening operations\nvary in degree from minor corrections of\nwarpage to complete forming of complex\nparts from solution-treated flat blanks. Par-\nticular value is gained when enough forming\ncan be done at this stage of processing to\neliminate the distortion caused by quench-\ning. However, production operations must\nbe adjusted so that most of the plastic\ndeformation is accomplished before an ap-\npreciable amount of precipitation hardening\ntakes place.\nAlthough the most severe forming opera-\ntions may have to be arranged to avoid\n\n\n0.5\n\n\n0\n\n\n1.0\n\n\n1.5\n\n\n2.0\n\n\n2.5\n\n\n3.0\n\n\n90\n\n\nHardness, HRB\n\n\n70\n\n\n60 O Control specimen\n\n\nQuenched from side A only\n\n\nA Quenched from side B, interrupted after 3 s\n\n\n50\n\n\n0\n\n\n10\n\n\n20\n\n\n30\n\n\n40\n\n\n50\n\n\n70\n\n\n60\n\n\n80\n\n\nSide A\n\n\nSide B\n\n\nDepth, mm\nDepth, in.\n\n\n0\n\n\n0.5\n\n\n1.5\n\n\n1.0\n\n\n2.0\n\n\n2.5\n\n\n3.0\n\n\nLongitudinal tensile strength, MPa\n\n\n600\n\n\nLongitudinal tensile strength, ksi\n\n\nb\n\n\n80\n\n\n500\n\n\n70\n\n\n60\n\n\n400\n\n\n-50\n\n\n300\n0\n\n\n80\n788\nSide B\n\n\n18\n50\n\n\n10\n\n\n20\n\n\n30\n\n\n40\n\n\n70\n\n\n60\n\n\nSide A\n\n\nDepth, mm\nDepth, in.\n\n\n0\n\n\n0.5\n\n\n1.0\nT\n\n\n1.5\n\n\n2.0\n\n\n2.5\n\n\nB\n3.0\nT\n\n\n550\n\n\nLongitudinal yield strength, MPa\n\n\n\u043e\n\n\n70\n\n\nLongitudinal yield strength, ksi\n\n\n450\n\n\n60\n\n\n50\n50\n\n\n350\n\n\n40\n40\n\n\n250\n\n\n30\n30\n\n\n150\n\n\n10\n\n\n0\n\n\n20\n\n\n30\n\n\n40\n\n\n50\n\n\n70\n\n\n60\n\n\n80\n\n\nSide A\n\n\nSide B\n\n\nDepth, mm\n\n\nThrough-thickness property variations due to quench rate and temperature-rise effects in 7075-T62 plate\n75 mm (3 in.) thick\n\n\nFig 6\n\n\nmetrical a section, the more it will warp\nduring quenching, and the residual stresses\nresulting from straightening of warped parts\n(plus straightening costs) often are less de-\nsirable than the quenching stresses. Holding\nfixtures and die quenching may be helpful,\nbut precautions must be taken to ensure\nthat they do not retard quenching rates\nexcessively. Other factors that must be con-\nsidered are the availability of heat-treating\nfacilities and whether or not the advantages\nof such a manufacturing sequence offset the\ndelay and cost entailed in a double-machin-\ning setup.\nWarpage of thin sections during quench-\ning is also a problem. Even in the same\nload, symmetry of cooling usually varies\nsignificantly among identical parts and the\n\n\nresultant inconsistent warpage usually re-\nquires costly hand straightening. Conse-\nquently, a significant amount of effort has\nbeen devoted to reducing or eliminating\nwarpage by changing racking positions to\nachieve symmetry of cooling.\nFor sheet-metal parts, one manufacturer\nuses a double screen floor in the quenching\nrack to reduce the force of initial contact\nbetween water and parts. Others allow parts\nto \"free fall\" from rack to quench tank.\nSpacing and positioning on the rack are\ncarefully controlled so that parts will enter\nthe water with minimum impact. With this\ntechnique, water turbulences must be\navoided, because it will often cause parts to\nfloat for a few seconds, greatly reducing\ntheir cooling rate.\n\n\n856 / Heat Treating of Nonferrous Alloys\n\n\nAverage cooling rate at 400-290 \u00b0C, \u00b0C/s\n\n\nprecipitation rate is low despite the high\ndegree of supersaturation. At intermediate\ntemperatures, precipitation rate is highest.\nConsequently, times to produce equal\namounts of precipitation follow a C-shape\npattern.\nUsing isothermal quenching techniques,\nFink and Willey pioneered the attempts to\ndescribe the effects of quench rates with the\nuse of C-curves (Ref 6). The C-curves plot\nthe time required at different temperatures\nto precipitate a sufficient amount of solute\nto: reduce strength by a certain amount (Fig\n5); cause a change in the corrosion behavior\nfrom pitting to intergranular (Fig 12); pro-\nduce a given electrical conductivity (Fig\n13); or relate other properties, such as frac-\nture toughness, to isothermal quench con-\nditions. The nose of the C-curves identifies\nthe critical temperature range (the region of\nhighest precipitation rates). Investigators\nuse critical temperature ranges in conjunc-\ntion with properties of samples quenched\ncontinuously from the solution temperature\nto compare relative sensitivities of alloys to\nquenching condition.\nAlthough average quench rates through a\ncritical temperature range can provide rea-\nsonable property predictions if cooling rates\nare fairly uniform, average quench rates can-\nnot provide quantitative predictions when\ncooling rates vary considerably during the\n\n\n103\n\n\n0.1\n\n\n1\n\n\n10\n\n\n100\n\n\n250\n\n\n10\n\n\nComputed maximum\n(assumes instantaneous\ncooling of surface from\n875-210 \u00b0F)\n\n\n75\n\n\n25\n25\n\n\nThickness, mm\n\n\nThickness, in.\n\n\n7.5\n\n\nAir cool\n\n\n2.5\n\n\n0.1\n\n\n212 \u00b0F\n\n\n180 \u00b0F\n\n\n75 \u00b0F\n\n\n150 \u00b0F\n\n\n200 \u00b0F\n\n\n0.75\n\n\nImmersion in water at\nindicated temperature\n\n\n0.25\n\n\n0.01\n\n\n103\n\n\n0.1\n\n\n1\n\n\n104\n\n\n10\n\n\n100\n\n\nAverage cooling rate at 750-550 \u00b0F, \u00b0F/s\n\n\nEffects of thickness and quenching medium on average cooling rates at midplane of aluminum alloy\nsheet and plate quenched from solution temperatures. The dashed line delineates the maximum cooling\nrates theoretically obtainable at the midplane of plate, assuming an infinite heat transfer coefficient (C) and a\ndiffusivity factor of 1400 cm\u00b2/s. Source: Ref 3\n\n\nFig 7\n\n\nciously for parts that are critical in fatigue quench. For such instances, a procedure\n(Fig 11) or stress corrosion.\n\n\nnatural aging, it often is desirable to allow\nsome natural aging to occur and thus avoid\nformation of L\u00fcders lines. This condition of\nnonuniform deformation is most likely to\noccur shortly after quenching and diminish-\nes significantly after a few hours of natural\naging. Complete freedom from L\u00fcders\nlines, however, may require one or two\ndays of natural aging prior to forming. Thus,\nthe forming operation may have to be timed\nso as to obtain the most appropriate trade-\noff of these characteristics for the specific\nparts involved. L\u00fcders lines also can be\nreduced by employing low strain rates or by\nforming at temperatures of 150 to 175 \u00b0C\n(300 to 350 \u00b0F).\nResidual stresses in sheet-metal parts\nformed in the quenched condition are higher\nthan those in parts formed in the annealed\ncondition. Consequently, forming in the\nquenched condition should be selected judi-\n\n\nknown as \"quench-factor analysis\" uses in-\nformation from the entire C-curve to predict\nhow any quench curve affects properties.\nQuench-factor analysis is useful in designing\nsuitable limits for quench delays, or when it is\nnot sufficient just to ensure that the cooling\ncurve misses the nose of the C-curve.\nThe method of quench-factor analysis, as\noutlined by Evancho and Staley (Ref 7), is\nbased on the determination of a quench\nfactor (T), which is the major variable in the\nfollowing equation for precipitation kinetics\nduring continuous cooling:\n\n\nof parts\n\n\nRe-solution heat treatment\nformed after quenching often causes exces-\nsive grain growth in critically strained re-\ngions and thus is not recommended.\nQuench-Factor Analysis\n\n\nDuring the quenching of alloys from a\nsolid-solution temperature condition, the\nrate of precipitation during quenching is\nmaximized in a so-called \"critical\" temper-\nature range, because the diffusion of dis-\nsolved species and the subsequent nucle-\nation of precipitates exhibit opposite\nbehavior as a function of temperature. At\nhigh temperatures, nucleation rates are\nsmall because of the low degree of super-\nsaturation, and so precipitation rates are\nlow despite the high diffusion rates. At low\ntemperatures, diffusion rate is low, and thus\n\n\n(Eq 1)\n\n\n(=1- exp (KT)\n\n\nwhere is the fraction transformed and k is\na constant related to the transformation\nfraction of a given C-curve. The quench\nfactor (T) is defined as:\n\n\n- \u1042\n\n\ndt\n\n\n(Eq 2)\n\n\nT=\n\n\nTable 7 Limits for quenching in glycol-water solutions\nData are for wrought aluminum alloy products other than forgings.\nGlycol\n\n\nwhere is time and C, is critical time as a\nfunction of temperature to transform a spec-\nified fraction (x). The locus of critical times\nfor a given transformation fraction x (or a\npercentage of mechanical properties from\nprecipitation) is the C-curve, and the value\nof k is related to x as follows: k = ln(1 - x),\nor e\u2b51\n= 1 x. Therefore, when T = 1, the\nfraction transformed, 5, equals the fraction\nvalue designated by the C-curve. Equation\n2 is based on the assumption that the reac-\ntion rate is a function only of the amount\ntransformed and temperature.\n\n\nMaximum thickness\n\n\nconcentration,\n\n\nvol%\n\n\nAlloys\n\n\nin.\n0.080\n1.000\n0.071\n0.500\n0.063\n0.375\n0.040\n0.250\n0.080\n\n\nmm\n\n\n12-16\n\n\n2014, 2017, 2117, 2024, 2219\n7075, 7175\n2014, 2017, 2117, 2024, 2219\n7075, 7079, 7175, 7178, 6061\n2014, 2017, 2117, 2024, 2219\n7075, 7079, 7175, 7178, 6061\n2014, 2017, 2117, 2024, 2219\n7075, 7079, 7175, 7178, 6061\n7075, 7079, 7175, 7178, 6061\n\n\n2.03\n\n\n25.4\n\n\n17-22\n\n\n1.80\n\n\n12.7\n\n\n23-28\n\n\n1.60\n9.53\n1.02\n6.35\n2.03\n\n\n29-34\n\n\n35-40\n\n\nHeat Treating of Aluminum Alloys / 857\n\n\n500\n\n\nThe numerical evaluation of the quench\nfactor involves the integration of Eq 2. This\nintegral can be graphically integrated using\nthe method illustrated in Fig 14. Examples\nof the way to use the quench factor (T) in the\nanalysis of quench methods are described\nbelow. Neither the average quenching rate\nthrough a critical temperature range nor\nquench-factor analysis can predict strength\nwhen the temperature increases during\nquenching after it is cooled below some\ncritical temperature. Under this condition,\nstrength in the affected areas can be signif-\nicantly lower than in other areas of the\nmaterial. The most likely way for this phe-\nnomenon to occur is during spray quench-\ning, when the surface cools rapidly by the\nimpinging spray, but reheats by heat flow\nfrom the hotter interior when the spray is\ninterrupted.\nPredicting Strengths of Thick Products.\nEffects of the quenching rate on alloy\nstrengths can be represented on a general-\nized graph of the type shown in Fig 3, and\nthe expected quenching rates of products\nhaving various dimensions can be deter-\nmined from Fig 7. Nevertheless, combining\nthese two kinds of information to predict\nmechanical properties must be done with\ncaution. Inconsistencies were encountered,\nfor example, in correlating properties of\nthick sections quenched in high-cooling-rate\nmedia with properties of thinner sections\nquenched in media affording milder quench-\ning action. One of the reasons for the incon-\nsistencies is believed to be the different\nshapes of the cooling curves. This difficulty\ncan be overcome by using quench-factor\nanalysis. The other reason is that the degree\nof recrystallization and texture of the thick\nand thin sections may be different.\nPredicting Corrosion Behavior. Alloy\n2024-T4, for example, is susceptible to in-\ntergranular corrosion when a critical\namount of solute is precipitated during\nquenching, but will corrode in the less se-\nvere pitting mode when lesser amounts are\nprecipitated. For predicting the effects of\nproposed quenching conditions on the cor-\nrosion characteristics of 2024-T4, the postu-\nlated quench curve is drawn and the quench\nfactor is calculated using the C-curve in Fig\n12. Corrosion characteristics are predicted\nfrom the plot in Fig 15. When the quench\nfactor (T) is less than 1.0, continuously\nquenched 2024-T4 will corrode by pitting.\nThese relationships are applied to studies\nof effects of proposed changes in quench\npractice on design of new quenching sys-\ntems. For example, consider that the goal of\na proposed quenching system for 2024-T4\nsheet products is to minimize warpage while\npreventing susceptibility to intergranular\ncorrosion. Warpage occurs when the stress-\nes imposed by temperature differences\nacross the parts exceed the flow stress. As\nquenching rate decreases, the tendency for\nlarge differences in temperature to occur\n\n\n20 \u00b0C water quench\n\n\n800\n\n\n400\n\n\nTemperature, \u00b0F\n\n\nTemperature, \u00b0C\n\n\n600\n\n\n300\n\n\n200\n\n\n400\n\n\n100\n\n\n200\n\n\nBlack oxide\netched coating\n\n\nAs-rolled\nsurface\n\n\nSanded\nsurface\n\n\n1\n\n\n2\n\n\n3\nTime, s\n\n\n5\n\n\n4\n\n\n6\n\n\n(a)\n\n\n500\n\n\nBoiling water quench\n\n\n800\n\n\n400\n\n\nTemperature. \u00b0F\n\n\nTemperature, \u00b0C\n\n\n600\n\n\n300\n\n\nSurface\n\n\n200\n\n\n400\n\n\nBlack oxide\netched coating\n\n\n100\n\n\n200\n\n\nPowdered oxide\nsprayed coating\n\n\nAs-rolled\nsurface\n\n\nSanded\nsurface\n\n\n0\n\n\n10\n\n\n20\n\n\n30\n\n\n40\n\n\n50\n\n\n60\n\n\n70\n\n\n80\n\n\nTime, s\n\n\n(b)\n\n\nEffect of surface conditions on the midplane cooling of a 13 mm (0.5 in.) thick plate of 7075 from\nquenching in (a) 20 \u00b0C (70 \u00b0F) water and (b) boiling water. Source: Ref 5\n\n\nFig 8\n\n\nCross section of solid cylindrical\nspecimen, in.2\n\n\nCross section of solid cylindrical\nspecimen, in.\n\n\n2\n\n\na\n\n\n10\n\n\n20 30\n\n\n40\n\n\n0\n\n\n20 30\nT\n\n\n10\n\n\n40\n\n\n100\n\n\n100\n\n\n1\n\n\nTension\u2192\u2192\n\n\nTension \u2192\n\n\nTension\n\n\n10\n\n\nQuenched\nin boiling\n\n\nQuenched\n\n\n10\n\n\nTension\n\n\nLongitudinal stress, MPa\n\n\nLongitudinal stress, ksi\nLongitudinal stress, MPa\n\n\nin cold\nwater\n\n\n50\n\n\nLongitudinal stress, ksi\n\n\nwater\n\n\n5\n\n\n0\n\n\nCompression\n\n\n\u2713 Compression\n\n\n- Compression\n\n\n- Compression\n\n\nT\n\n\n5\n\n\n50\n\n\n50\n\n\n10\n\n\n10\n\n\n100\n\n\n100\n\n\n0\n\n\n10\n\n\n20\n\n\n0\n\n\n30\n\n\n10\n\n\n20\n\n\n30\n\n\nCross section of solid cylindrical\nspecimen, 103 mm\u00b2\n\n\nCross section of solid cylindrical\nspecimen, 103 mm\u00b2\n\n\nFig 9 Effect of quenching from 540 \u00b0C (1000 \u00b0F) on residual stresses in solid cylinders of alloy 6151\n\n\n858 / Heat Treating of Nonferrous Alloys\n\n\n400\n\n\nThickness, 0.001 in.\n\n\n150\n\n\n200\n\n\n100\n\n\n250\n\n\nCondition during flattening\n\n\nCurve Bend radius\n\n\n100\n\n\n3\n123\n4\n\n\nNot bent\n3.2 mm\n3.2 mm\n3.2 mm\n1.6 mm\n\n\nNot applicable\n\n\nWater\n\n\n50\n\n\nAnnealed\n\n\nspray\n\n\nAs-quenched + 3 days storage\nAs-quenched + 14 days storage\nAs-quenched + 3 days storage\n\n\n90\n\n\n% of maximum yield strength\n\n\n300\n\n\n5\n\n\nStress, MPa\n\n\nStress, ksi\n\n\n40\n\n\nAir blast\n\n\n70\n\n\n30\n\n\n200\n\n\n50\n60\n\n\nStress ratio, 0.1\n\n\n6061-T6 sheet\n\n\n1\n\n\n20\n\n\n2\n\n\n50\n\n\n3750\n\n\n2500\n\n\n6250\n\n\n1250\n\n\n7500\n\n\n5000\n\n\n3\n\n\n5\n\n\n4\n\n\n100\n\n\nThickness, um\nThickness, 0.001 in.\n\n\n0.1\n\n\n1\n\n\n0.01\n\n\n10\n\n\nMillions of cycles to failure\n\n\n150\n\n\n200\n\n\n250\n\n\n100\n\n\n100\n\n\nFig 11 Fatigue characteristics of 1 mm (0.04 in.) alclad 2024-T4 sheet after 90\u00b0 bending in the annealed\ncondition and subsequent flattening as indicated. Flattening (unbending) was done either in the\nannealed condition (curve 2), or in the solution-treated and quenched condition (curves 3, 4, 5) with indicated\nstorage times at -18 to -12 \u00b0C (0 to 10 \u00b0F).\n\n\n% of maximum tensile strength\n\n\nWater spray\n\n\n90\n\n\n80\n\n\nable corrosion behavior in 2024-T4 sheet\n(quench factor, 0.99) were calculated. Some\nof these curves are plotted in Fig 16. This\nillustration shows that 2024 can be\nquenched at a rate of 470 \u00b0C/s (850 \u00b0F/s) or\nhigher and still develop acceptable corro-\nsion characteristics if the quenching rate is\nlinear from the solution temperature to 150\n\u00b0C (300 \u00b0F). If sheet 3.2 mm (0.125 in.) thick\nis air-blast quenched (rate of heat removal,\n5.68 W/m\u00b2 \u00b0C) to 395 \u00b0C (740 \u00b0F), however,\nthe quenching rate from 395 to 150 \u00b0C must\nbe at least 945 \u00b0C/s (1700 \u00b0F/s) to maintain\nthe acceptable corrosion behavior. It may\nalso be air-blast quenched to 395 \u00b0C (740\n\u00b0F), spray quenched at 3300 \u00b0C/s (6000 \u00b0F/s)\nto 250 \u00b0C (480 \u00b0F), then air-blast quenched\nto 150 \u00b0C (300 \u00b0F).\n\n\nOther curves could be drawn, of course,\nbut the important points are that air-blast\nquenching cannot be continued to more\nthan a few degrees below 395 \u00b0C (740 \u00b0F)\nand cannot be initiated at more than a few\ndegrees above 270 \u00b0C (520 \u00b0F) even if infi-\nnite quenching rates are attained from 395\nto 270 \u00b0C (740 to 520 \u00b0F).\nPredicting yield strength is more complex\nthan predicting corrosion behavior and re-\nquires some knowledge of the relationship\nbetween extent of precipitation and loss in\nability to develop property. Because attain-\nable strength of precipitation-hardening alu-\nminum alloys is a function of the amount of\nsolute remaining in solid solution after\nquenching, relationships between strength\n(\u03c3) attainable after continuous cooling and\n\n\nAir blast\n\n\n70\n\n\n6061-T6 sheet\n\n\n601250\n\n\n2500\n\n\n3750\n\n\n5000\n\n\n6250\n\n\n7500\n\n\nThickness, um\n\n\nEffect of quenching medium on strength of\n6061-T6 sheet. Water-immersion quench\nequals 100%. Control of coolant flow will minimize\ndecrease in mechanical properties.\n\n\nFig 10\n\n\ndecreases but the tendency for intergranular\ncorrosion to occur increases.\nThe C-curve in Fig 12 indicates that\nquenching rate can be decreased near the\nsolution heat-treating temperature and near\nroom temperature without greatly sacrific-\ning corrosion characteristics, but this infor-\nmation does not provide a quantitative an-\nswer. Simple calculations, however, can\nreveal a multitude of hypothetical cooling\ncurves that provide slow quenching during a\nlarge portion of the quench cycle but suffi-\nciently rapid quenching where critical times\nare short so that desirable corrosion char-\nacteristics are obtained.\nAs an example, one-, two-, and three-step\nquench curves that would ensure accept-\n\n\n425\n\n\n700\n\n\nIACS\n\n\n18%\n\n\n17%\n\n\n375\n\n\n650\n\n\nHolding temperature, \u00b0C\n\n\nHolding temperature, K\n\n\n19%\n\n\n325\n\n\n600\n\n\n20%\n\n\n21%\n\n\n275\n\n\n550\n\n\nTemperature, \u00b0C\n\n\n500\n\n\nTemperature, \u00b0F\n\n\n900\n\n\n500\n\n\n225\n\n\n19%\n\n\n400\n\n\n17%\n\n\n20%\n\n\nPredominantly\nintergranular\ncorrosion\n\n\n21%\n\n\n700\n\n\n22%\n\n\n300\n\n\n18%\n\n\n500\n\n\nPredominantly\npitting\n1\n\n\n200\n\n\n175\n\n\n450\n\n\n300\n\n\n104\n\n\n105\n\n\n106\n\n\n103\n\n\n10\n\n\n100\n\n\n100\n0.1\n\n\n103\n\n\n10\n\n\n100\n\n\nHolding time, s\n\n\nCritical time, s\n\n\nFig 13 Change in electrical conductivity of an Al-2.5% Li binary alloy after the following: solution treated at 540\n\u00b0C (1000 \u00b0F) for 12 h, immersed into an adjacent salt or oil bath for the appropriate isothermal holding\ntemperature and time, then quenched into water. Source: Ref 1\n\n\nC-curve indicating type of corrosion attack\non 2024-T4 sheet\n\n\nFig 12\n\n\nHeat Treating of Aluminum Alloys / 859\n\n\nQuench curve\n\n\nC-curve\n\n\n500\n\n\n900\n\n\nAir-blast quench\n\n\n3.2-mm (0.125-in.) sheet\n\n\non both sides\n\n\n800\n\n\n\u30c9\u30c9\u30c9\n\n\n(T\u2081 + T\u2082)/2\n\n\n400\n\n\n470 \u00b0C/s (850 \u00b0F/s)\n\n\nTemperature, \u00b0C\n\n\nTemperature, 'F\n\n\n+ T3)/2\n\n\n\u3051\n\n\n700\n\n\nTemperature\n\n\nTemperature\n\n\n2780 \u00b0C/s (5000 \u00b0F/s)\nAir-blast quench\n\n\n600\n\n\n300\n\n\n3.2-mm (0.125-in.)\n\n\nsheet\n\n\n500\n\n\n200\n\n\n400\n\n\n(TF + TF)/2\n\n\nTEITI\n1++AG\n\n\n945 \u00b0C/s\n\n\n3330 \u00b0C/s\n(6000 F/s)\n\n\n300\n\n\n(1700 F/s)\n\n\n|\n\n\n1+AtF-1\nTF-1 tF\nElapsed time\n\n\n1\n4\n\n\n100\n\n\n0\n\n\n2\n\n\n6\n\n\n\u04212\n\n\nC\u2081\u2081 CF-1\nCritical time\n\n\nTime, s\n\n\nFig 16 Quench curves for 2024-T4 sheet, to elimi-\nnate susceptibility to intergranular corrosion\n\n\n+\n\u0394\u03b9\u03b1 \u0394\u03b9\u03b1\nC\u2082\n\n\nAtF-1\nCF-1\n\n\nx =\n\n\n+\n\n\nThe advantage of using the quench factor\nfor predicting yield strength from cooling\ncurves is apparent. Cooling curves that\nhave long holding times either above or\nbelow the critical temperature range from\n400 to 290 \u00b0C (750 to 550 \u00b0F) cannot be used\nto predict yield strength from average\nquenching rate. In such instances, predic-\ntion of yield strength on the basis of quench\nfactor is particularly advantageous.\nWith the use of finite-element analysis,\nquench factors can also be plotted as a\nfunction of Grossmann quench severity val-\nues (H) or the heat transfer coefficients (C)\nbetween the quenchant and a particular part\n(Fig 20). However, an underlying assump-\ntion of both quench-factor analysis and av-\nerage-cooling-rate estimation is that the\nonly effect of temperature is on the kinetics\nof precipitation. This assumption is not val-\nid, however, when portions of the metal are\nquenched locally but reheated significantly\nbefore quenching is complete.\n\n\nFig 14 Method of determining quench factor, &, using a cooling curve and a C-curve\n\n\n0.20\n\n\nQuench\n\n\no Air-water-air\n\n\n\u2022 Water-air\nA Air-water\n\n\n0.006\n\n\n0.15\n\n\nAverage depth, mm\n\n\nAverage depth, in.\n\n\n\u25b2 Water\n\n\n\u25a1 Air\n\n\nQA\n\n\n0.004\n\n\n0.10\n\n\n0.05\n\n\n0.002\n\n\n1\n0.2\n\n\n\u0567\n0.4\n\n\nT\n\n\n1\n\n\nL\n4.0\n\n\n0\n\n\n1\n2.0\n\n\n0.6 0.8 1.0\n\n\n6.0 8.0 10.0\n\n\n20.0\n\n\n0.1\n\n\nQuench factor, T\n\n\nP-Pitting\n\n\n2 P+SI - Pitting and slight intergranular\nP+1 Pitting and intergranular\nIntergranular\n\n\nType of corrosion\n\n\nP + SI\n\n\nQuench\nAir-water-air\n\n\nP +\n\n\n\u2022 Water-air\n\u25b2 Air-water\n\n\n0800\n\n\n\u25b2 Water\n\n\n\u25a1 Air\n\n\nAge Hardening\n\n\n0.6 0.8 1.0\n2.0\nQuench factor, 1\n\n\n0.2\n\n\n4.0 6.0 8.0 10.0\n\n\n20.0\n\n\n0.1\n\n\n0.4\n\n\nAfter solution treatment and quenching,\nhardening is achieved either at room tem-\nperature (natural aging) or with a precipita-\ntion heat treatment (artificial aging). In\nsome alloys, sufficient precipitation occurs\nin a few days at room temperature to yield\nstable products with properties that are\nadequate for many applications. These al-\nloys sometimes are precipitation heat treat-\ned to provide increased strength and hard-\nness in wrought or cast products. Other\nalloys with slow precipitation reactions at\nroom temperature are always precipitation\nheat treated before being used.\nIn some alloys, notably those of the 2xxx\nseries, cold working of freshly quenched\nmaterial greatly increases its response to\nlater precipitation heat treatment. Mills take\nadvantage of this phenomenon by applying\na controlled amount of rolling (sheet and\nplate) or stretching (extrusion, bar, and\nplate) to produce higher mechanical proper-\nties. However, if the higher properties are\nused in design, reheat treatment must be\navoided.\n\n\nFig 15 Type and depth of attack on 2024-T4 sheet versus quench factor\n\n\nquench factor (T) can be expressed as fol-\nlows:\n\n\nfollowing comparison. Four specimens of\nalloy 7075-T6 quenched by various means\n(see Fig 17) were selected. Yield strengths\nwere predicted both from average quench-\ning rate between 400 and 290 \u00b0C (750 and\n550 \u00b0F) and from quench factor. Quench\nfactor was calculated using the C-curve for\n99.5% maximum yield strength for 7075-T6\n(Fig 18), and yield strength was estimated\nfrom the above equation defining the\nquench factor (7) (see Fig 19).\nA comparison of predicted yield strength\nwith actual yield strength is given in Table\n8. Yield strengths predicted from quench\nfactor agree very well with measured yield\nstrengths for all specimens, the maximum\nerror being 19.3 MPa (2.8 ksi). Yield\nstrengths predicted from average quenching\nrates, however, differ from measured val-\nues by as much as 226 MPa (32.8 ksi).\n\n\n\u03c3x=\u03c3max exp (k\u2081T)\nwhere \u03c3 max\nis the strength attainable with\nan infinite quenching rate and:\n\n\n(Eq 3)\n\n\ndt\n\n\nSa\n\n\n(Eq 4)\n\n\nT =\n\n\nwhere t is time and Cx is the C-curve for\n\u03c3 that is, critical time as a function of\ntemperature to reduce attainable strength to\nx of \u03c3max. The constant k\u2081 is related to the\nnatural logarithm of x. For example, if 7 is\nbased on the C-curve for 99.5% of maxi-\nmum yield strength, then k\u2081 = -0.005013 =\nIn (0.995).\nThe advantage of predicting yield\nstrength from quench factor instead of from\naverage quenching rate is illustrated by the\n\n\n860 / Heat Treating of Nonferrous Alloys\n\n\n500\n\n\n900\n\n\nQuenched in still air to 370 \u00b0C (700 \u00b0F) + cold-water quench\n\n\n500\n\n\n900\n\n\n800\n\n\n400\n\n\n700\n00\n\n\n- 400\n\n\nTemperature, \u00b0C\n\n\nTemperature, \"C\n\n\nTemperature, \"F\n\n\nTemperature, \u00b0F\n\n\nQuenched in boiling water to\n\n\n700\n\n\n315 \u00b0C (600 \u00b0F) + cold-water quench\n\n\n-600\n\n\n300\n\n\n300\n\n\n\" 500\n\n\n500\n\n\n200\n\n\nF6 Quenched in denatured alcohol to 290 \u00b0C\n\n\n200\n\n\n400\n\n\n\u2020(550 \u00b0F) cold-water quench\n\n\n300\n\n\n100\n0.1\n\n\n300\n\n\nCold-water quench\n\n\n10 000\n\n\n100\nCritical time, s\n\n\n1\n\n\n10\n\n\n1000\n\n\n100\n\n\n0\n\n\n5\n\n\n20\n\n\n10\n\n\n30\n\n\n35\n\n\n15\n\n\n25\n\n\nTime, s\n\n\nFig 17 Cooling curves for 7075-T6 sheet\n\n\nFig 18 C-curve for 99.5% maximum yield strength of 7075-T6 sheet\n\n\nNatural Aging. The more highly alloyed\nmembers of the 6xxx wrought series, the\ncopper-containing alloys of the 7xxx group,\nand all of the 2xxx alloys are almost always\nsolution heat treated and quenched. For\nsome of these alloys-particularly the 2xxx\nalloys the precipitation hardening that re-\nsults from natural aging alone produces\nuseful tempers (T3 and T4 types) that are\ncharacterized by high ratios of tensile to\nyield strength and high fracture toughness\nand resistance to fatigue. For the alloys that\nare used in these tempers, the relatively\nhigh supersaturation of atoms and vacan-\ncies retained by rapid quenching causes\nrapid formation of GP zones, and strength\nincreases rapidly, attaining nearly maxi-\nmum stable values in four or five days.\nTensile-property specifications for products\nin T3- and T4-type tempers are based on a\nnominal natural aging time of four days. In\nalloys for which T3- or T4-type tempers are\nstandard, the changes that occur on further\nnatural aging are of relatively minor magni-\ntude, and products of these combinations of\nalloy and temper are regarded as essentially\nstable after about one week.\nIn contrast to the relatively stable condi-\ntion reached in a few days by 2xxx alloys\nthat are used in T3- or T4-type tempers, the\n6xxx alloys and to an even greater degree\nthe 7xxx alloys are considerably less stable\nat room temperature and continue to exhibit\nsignificant changes in mechanical properties\nfor many years. The differences in rate and\n\n\nduration of changes in tensile yield strength\nof representative alloys of the three types\nare illustrated in Fig 21. Because of the\nrelative instability of the 7xxx alloys, the\nnaturally aged temper (after solution heat\ntreatment and quenching) is designated by\nthe suffix letter W. For a specific descrip-\ntion of this condition, the time of natural\naging should be included (example: 7075-W,\n1 month).\nAging characteristics vary from alloy to\nalloy with respect to both time to initial\nchange in mechanical properties and rate of\nchange, but aging effects always are less-\nened by reductions in aging temperature\n(see Fig 21). With some alloys, aging can be\nsuppressed or delayed for several days by\nholding at a temperature of -18 \u00b0C (0 \u00b0F) or\nlower. It is usual practice to complete form-\ning and straightening before aging changes\nmechanical properties appreciably. When\nscheduling makes this impractical, aging\nmay be avoided in some alloys by refriger-\nating prior to forming. It is conventional\npractice to refrigerate alloy 2024-T4 rivets\nto maintain good driving characteristics.\nFull-size wing plates for current-generation\njet aircraft have been solution heat treated\nand quenched at the primary fabricating\nmill, packed in dry ice in specially designed\ninsulated shipping containers and transport-\ned by rail about 2000 miles to the aircraft\nmanufacturer's plant for forming.\n\n\nUnanticipated difficulties may arise as a\nresult of failure to control refrigerator or\npart temperature closely enough. If opening\nof the cold box to insert or remove parts is\ndone too frequently, the cooling capacity of\nthe refrigerator may be exceeded. At times,\nthe rate at which heavy-gage parts can be\ncooled in a still-air cold box has been found\nto be insufficient. This problem has been\nsolved in one plant by immersing parts in a\nsolvent at -40 \u00b0C (-40 \u00b0F) before placing\nthem in the refrigerator.\nThe T3-type tempers are distinguished\nfrom T4-type tempers by significant me-\nchanical-property differences resulting from\ncold work strain hardening associated with\ncertain mechanical operations performed\nafter quenching. Roller or stretcher leveling\nto achieve flatness or straightness introduc-\nes modest strains (on the order of 1 to 4%)\nthat cause changes in mechanical properties\n(primarily, increases in strength). Further\nincreases in strength can be obtained by\ncold rolling, additional stretching, combina-\ntions of these operations, or for products\nsuch as hand forgings, compressive defor-\nmation. The tempers produced by these\noperations followed by natural aging alone\n(no precipitation heat treatment) are classi-\nfied as T3-type tempers, and an additional\ndigit is used to indicate a variation in strain\nhardening that results in significant changes\nin properties. In the most recently intro-\n\n\nTable 8 Yield-strength values for 7075-T6 sheet predicted from cooling curves using\naverage quench rate and quench factor\n\n\nYield strength\npredicted from\naverage quench\nrate\n\n\nAverage quench\nrate from 400 to\n290 \u00b0C (750 to\n550 \u00b0F)\n\n\n% of maximum yield strength\n\n\nYield strength\npredicted from\nquench factor\n\n\nMeasured yield\nstrength\n\n\n100\n\n\nQuench\nfactor,\nT\n\n\n80\n\n\n\u00b0C/s\n\n\nksi\n\n\nQuench\n\n\n\u00b0F/s\n\n\nMPa\n\n\nMPa\n\n\nksi\n\n\nMPa\n\n\nksi\n\n\nCold water\n\n\n72.3\n\n\n0.464\n\n\n73.4\n\n\n498\n\n\n935\n\n\n1680\n\n\n506\n\n\n499\n\n\n72.4\n\n\n100 o/max exp(-0.005013 t)\n\n\n60\n\n\nDenatured alcohol to\n290 \u00b0C (550 \u00b0F),\nthen cold water\n\n\n40\n\n\n8.539\n\n\n476\n\n\n463\n\n\n67.2\n\n\n478\n\n\n50\n\n\n90\n\n\n69.1\n\n\n69.4\n\n\n20\n\n\nBoiling water to\n\n\n315 \u00b0C (600 \u00b0F),\nthen cold water\n\n\n0\n\n\n64.2\n\n\n15.327\n\n\n67.1\n\n\n55\n\n\n458\n\n\n66.4\n\n\n443\n\n\n463\n\n\n30\n\n\n103\n\n\n0.01\n\n\n10\nQuench factor, t\n\n\n1\n\n\n0.1\n\n\n100\n\n\nStill air to 370 \u00b0C\n\n\n(700 \u00b0F), then cold\nwater\n\n\n21.334\n\n\n9\n\n\n468\n\n\n67.9\n\n\n242\n\n\n35.1\n\n\n449\n\n\n65.1\n\n\n5\n\n\nFig 19 Yield strength versus quench factor\n\n\nHeat Treating of Aluminum Alloys / 861\n\n\nSheet thickness, in.\n\n\nPlate thickness, in.\n\n\n0.05\n\n\n0.15\n\n\n0.2\n\n\n0.25\n\n\n0.3\n\n\n0\n\n\n2\n\n\n0\n\n\n0.1\n\n\n0.5\n\n\n1\n\n\n1.5\n\n\n2.5\n\n\n3\n\n\n50\n550\n\n\n50\n\n\nC = 0.216\n\n\nC=0.036 W/cm2.K\n\n\n45\n\n\n45\n\n\nC = 0.072\n\n\nattainable yield strength for alloy 7075-T73\n\n\nattainable yield strength for alloy 7075-T73\n\n\n40\n\n\nQuench factor for 99.5% of\n\n\nQuench factor for 99.5% of\n\n\nC=0.108\n\n\nC = 2.16\nC = 2.88\nC = 4.32\nC=7.20\n\n\nC = 0.144\n\n\nC = 0.216\n\n\n1\n\n\n1\n\n\nC = 0.288\n\n\n10\n\n\n10\n\n\nC = 0.432\n\n\nC = 0.720\nC = 1.44\n6.25\n\n\n5\n\n\n5\n\n\n3.75\nSheet thickness, mm\n\n\n7.5\n\n\n1.25\n\n\n0\n\n\n2.5\n\n\n5\n\n\n25\n\n\n37.5\n\n\n50\n\n\n62.5\n\n\n75\n\n\n12.5\n\n\nPlate thickness, mm\n\n\n(a)\n\n\n(b)\n\n\nPlot of quench factors derived from finite element analysis with given product sizes and film (heat transfer) coefficients (C). Heat transfer coefficients between\nFig 20\nthe quenchant and part are expressed in W/cm\u00b2 K. Source: Ref 4\n\n\nduced 2xxx aircraft alloy, 2324, high\nstrength is achieved by cold rolling plate to\na T39 temper.\nPrecipitation heat treatments generally\nare low-temperature, long-term processes.\nTemperatures range from 115 to 190 \u00b0C (240\nto 375 \u00b0F); times vary from 5 to 48 h.\nChoice of time-temperature cycles for\nprecipitation heat treatment should receive\ncareful consideration. Larger particles of\nprecipitate result from longer times and\nhigher temperatures; however, the larger\nparticles must, of necessity, be fewer in\nnumber with greater distances between\nthem. The objective is to select the cycle\nthat produces optimum precipitate size and\ndistribution pattern. Unfortunately, the cy-\ncle required to maximize one property, such\nas tensile strength, is usually different from\nthat required to maximize others, such as\nyield strength and corrosion resistance.\nConsequently, the cycles used represent\ncompromises that provide the best combi-\nnations of properties.\nProduction of material in T5- through\nT10-type tempers (see the section on tem-\nper designations near the end of this article)\nnecessitates precipitation heat treating at\nelevated temperatures (artificial aging). Al-\nthough the hardening precipitate developed\nby this operation is submicroscopic, struc-\ntures before and after precipitation heat\ntreatment often can be distinguished by\netching metallographic specimens. In alumi-\nnum alloys in the solution heat treated and\nquenched condition, coloration contrast be-\ntween grains of differing orientation is rela-\ntively high, particularly in 2xxx series\n\n\ntures of the precipitation-hardening process\nreduce the probability of obtaining the re-\nquired properties.\nT6 and T7 Tempers. Precipitation heat\ntreatment following solution heat treatment\nand quenching produces T6- and T7-type\ntempers. Alloys in T6-type tempers gener-\nally have the highest strengths practical\nwithout sacrifice of the minimum levels of\nother properties and characteristics found\nby experience to be satisfactory and useful\nfor engineering applications. Alloys in T7\ntempers are overaged, which means that\nsome degree of strength has been sacrificed\nor \"traded off\" to improve one or more\nother characteristics. Strength may be sac-\nrificed to improve dimensional stability,\nparticularly in products intended for service\nat elevated temperatures, or to lower resid-\nual stresses in order to reduce warpage or\ndistortion in machining. T7-type tempers\nfrequently are specified for cast or forged\nengine parts. Precipitation heat-treating\ntemperatures used to produce these tem-\npers generally are higher than those used to\nproduce T6-type tempers in the same al-\nloys.\nTwo important groups of T7-type\ntempers-the T73 and T76 types-have\nbeen developed for the wrought alloys of\nthe 7xxx series, which contain more than\nabout 1.25% copper. These tempers are\nintended to improve resistance to exfolia-\ntion corrosion and stress-corrosion crack-\ning, but as a result of overaging, they also\nincrease fracture toughness and, under\nsome conditions, reduce rates of fatigue-\ncrack propagation. The T73-type temper\n\n\nwrought alloys and 2xx.0 series casting al-\nloys. This contrast is noticeably decreased\nby precipitation heat treatment.\nDifferences in type, volume fraction,\nsize, and distribution of the precipitated\nparticles govern properties as well as the\nchanges observed with time and tempera-\nture, and these are all affected by the initial\nstate of the structure. The initial structure\nmay vary in wrought products from unre-\ncrystallized to recrystallized and may ex-\nhibit only modest strain from quenching or\nadditional strain from cold working after\nsolution heat treatment. These conditions,\nas well as the time and temperature of\nprecipitation heat treatment, affect the final\nstructure and the resulting mechanical prop-\nerties.\nBecause mechanical properties and other\ncharacteristics change continuously with\ntime and with temperature, as shown in Fig\n22(a), (b), and (c) by typical curves for three\nwrought alloys, treatment to produce a\ncombination of properties corresponding to\na specific alloy-temper combination re-\nquires one or more rather specific and co-\nordinated combinations of time and temper-\nature, with both parameters being subject to\npractical limitations. Recommended com-\nmercial treatments often are compromises\nbetween time and cost factors and the prob-\nability of obtaining the intended properties,\nwith consideration of allowances for varia-\nbles such as composition within specified\nrange and temperature variations within the\nfurnace and load. Use of higher tempera-\ntures may reduce treatment time; but if the\ntemperature is too high, characteristic fea-\n\n\n862 / Heat Treating of Nonferrous Alloys\n\n\n2014\n\n\n2024\n\n\nsition. If first-step aging time is too short,\nif first-step aging temperature is too far\nbelow the GP-zone solvus, or if heating\nrates are too high, the GP zones will\ndissolve above 150 \u00b0C (300 \u00b0F), and the\nresultant coarse and widely distributed\nprecipitate will provide lower strength.\nThe T76-type treatments have the same\noperational sequence but employ second-\nstage heating only long enough to develop\na resistance to exfoliation corrosion higher\nthan that provided by the T6-type tempers.\nMaterials in the T73-type temper also have\nhigh resistance to exfoliation corrosion.\nRecommended treatments to produce T5-\nand T6-type tempers, and those of the T7-\ntype employed for dimensional and proper-\nty stabilization, provide adequate tolerance\nfor normal variations encountered with\ngood operating practices. On the other\nhand, the T73, T74 (formerly T736), and\nT76 tempers for alloys 7049, 7050, 7075,\n7175, and 7475 involve changes in strength\nthat occur significantly more rapidly at the\ntemperatures employed in the second stage\nof the T7x precipitation heat-treatment cy-\ncle compared to the changes occurring at\nthe temperatures employed to produce the\nT6 temper.\nAs illustrated in Fig 23, variations in soak\ntime of several hours, and variations in soak\ntemperature of up to 11 \u00b0C (20 \u00b0F) from the\nnominal aging practice of 24 h at 120 \u00b0C (250\n\u00b0F) affect the strength of 7075-T6 by as\nmuch as 28 MPa (4 ksi). In contrast, similar\nvariations in second-step soak time and\ntemperature for 7075-T73-that is, varia-\ntions for 24 h at 165 \u00b0C (325 \u00b0F)-affect\nstrength by up to 150 MPa (22 ksi).\nConsequently, control of both tempera-\nture and time to achieve the mechanical\nproperties and corrosion resistance speci-\nfied for these tempers is more critical than\nthe control required in producing the T6\ntemper. Moreover, rate of heating from the\nfirst to the second aging step must be con-\nsidered, because precipitation occurs dur-\ning this period.\nHeat treaters attempt to adjust these new\nproblems by empirically modifying soak\ntimes to compensate for precipitation dur-\ning heating and for effects of soaking at\ntemperatures above or below the nominal.\nA method has been developed (Ref 8) that\npermits quantitative compensation for the\neffects of precipitation during heating and of\nsoaking either above or below the recom-\nmended temperature. For overaging, these\neffects can be described by the following\nequation:\n\n\n600\n\n\n600\n\n\nT\n\n\n80\n\n\n80\n\n\n500\n\n\n500\n\n\n70\n\n\n70\n\n\nRT\n\n\nTensile strength, MPa\n\n\nRT\n\n\nTensile strength, MPa\n\n\nTensile strength, ksi\n\n\nTensile strength, ksi\n\n\n0 \u00b0C (32\u00b0F)\n\n\n60\n\n\n60\n\n\n0 \u00b0C (32 \u00b0F).\n\n\n400\n\n\n400\n\n\n-18 \u00b0C (0 \u00b0F)\n\n\nT\n\n\n50\n\n\n50\n\n\n--18\u00b0C (0\u00b0F)\n\n\n300\n\n\n300\n\n\nT\n\n\n40\n\n\n40\n\n\n30\n\n\n30\n\n\n200\n\n\n200\n\n\n1 year\n\n\n1 year\n\n\n1\n\n\n1 week 2 months\n\n\n30 min\n\n\n20\n10\n\n\n30 min\n\n\n1 week 2 months.\n\n\n1 day\n\n\n1 day\n\n\n20\n\n\n100\n\n\n100\n\n\n104\n\n\n102\n\n\n102\n\n\n104\n\n\n103\n\n\n103\n\n\n1\n\n\n0.1\n\n\n10\n\n\n10\n\n\n0.1\n\n\n1\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\n500\n\n\n70\n20\n\n\n500\n\n\nT\n\n\n70\n\n\n1\n\n\n60\n\n\n60\n\n\n400\n\n\n400\n\n\nT\n\n\nI\n\n\n50\n50\n\n\nYield strength, MPa\n\n\n50\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\nYield strength, ksi\n\n\n300\n\n\n300\n\n\n40\n\n\n40\n\n\nRT\n\n\nRT\n\n\n0\u00b0C (32\u00b0F)\n\n\n30\n30\n\n\n0\u00b0C (32 F)\n\n\n30\n\n\n200\n\n\n200\n\n\n-18 C (OF)\n\n\n20\n20\n\n\n20\n\n\n-18 \u00b0C (0 \u00b0F)\n\n\n100\n\n\n100\n\n\n10\n\n\n1 year\n\n\n10\n\n\n1 year\n\n\n30 min\n\n\n1 week 2 months\n\n\n1 week 2 months\n\n\n1 day\n\n\n1 day\n\n\n30 min\n\n\n11\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n0.1\n\n\n102\n\n\n104\n\n\n103\n\n\n102\n\n\n103\n\n\n104\n\n\n0.1\n\n\n1\n\n\n10\n\n\n1\n\n\n10\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\n40\n\n\n40\n\n\nElongation, % in 50 mm or 2 in.\n\n\nElongation, % in 50 mm or 2 in.\n\n\n30\n\n\n30\n\n\n-18 \u00b0C (0\u00b0F)\n0\u00b0C (32\u00b0F)\n\n\n-18 \u00b0C (0 \u02daF)\n\n\nRT\n\n\n20\n20\n\n\n20\n\n\n0 \u00b0C (32\u00b0F)\n\n\n10\n10\n\n\n10\n\n\n1 year\n\n\n1 year\n\n\n30 min\n\n\n1 week 2 months\n\n\n1 day\nI\n\n\n30 min\n\n\n1 day\nI\n\n\n1 week 2 months\n\n\n1\n102\n\n\n11\n103\n\n\n0\n\n\n0\n\n\n102\n\n\n104\n\n\n1\n\n\n104\n\n\n0.1\n\n\n10\n\n\n103\n\n\n0.1\n\n\n1\n\n\n10\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\nFig 21 Aging characteristics of aluminum sheet alloys at room temperature, at 0 \u00b0C (32 \u00b0F), and at -18 \u00b0C (0 \u00b0F)\n\n\nhas greatly minimized stress-corrosion\ncracking of large and complex machined\nparts made of these alloys, which occa-\nsionally occurred with T6-type tempers.\nThe precipitation heat treatment used to\nproduce the T73- and T76-type tempers\nconsist either of a two-stage isothermal\nprecipitation heat treatment or of heating\nat a controlled rate to a single treatment\ntemperature. The microstructural/\nelectrochemical relationships that are re-\nquired in order to achieve the desired\ncorrosion-resisting characteristics can be\ndeveloped by using only a single-stage\nprecipitation heat treatment above about\n150 \u00b0C (300 \u00b0F), but higher strength is\nobtained by preceding this with a lower-\ntemperature stage or with a slow-con-\n\n\ntrolled heatup. Extended natural aging can\nprovide the same results, but the times\nrequired at room temperature are imprac-\ntical. Either during the preliminary stage\nor during slow heatup, a fine, high-density\ndispersion of GP zones is nucleated. Either\nthe time and temperature of the first step\nor the rate of heating must be controlled to\nproduce GP zones that will not dissolve\nbut will transform to the n' precipitate\nwhen heated to the aging temperature\nabove 150 \u00b0C (300 \u00b0F). The aging practice\nthat produces the results in the shortest\ntime depends on the GP-zone solvus tem-\nperature. This temperature, in turn, de-\npends on vacancy concentration, a factor\ninfluenced by solution heat-treating tem-\nperature and quench rate, and on compo-\n\n\n(\n\n\nYS = Y exp\n\n\n(Eq 5)\n\n\n+ 0\nFYS\n\n\n-\n\n\nwhere YS is yield strength; Y is a term\nhaving units of strength that is dependent on\nalloy, fabrication, and test direction; to is\ntime at soak temperature; Fys is a temper-\nature-dependent term; and\n\n\nYS\n\n\nHeat Treating of Aluminum Alloys / 863\n\n\n7050\n\n\n7075\n\n\n6061\n\n\n600\n\n\n600\n\n\n600\n\n\n00\n80\n\n\n80\n\n\nT\n\n\n80\n\n\n500\n\n\n500\n\n\n500\n\n\n70\n\n\n70\n\n\n70\n\n\n0\u00b0C (32\u00b0F)\n\n\nTensile strength, MPa\n\n\nRT\n\n\nTensile strength, MPa\n\n\nTensile strength, ksi\n\n\nTensite strength, MPa\n\n\nTensile strength, ksi\n\n\nRT\n\n\nTensile strength, ksi\n\n\n60\n\n\nT\n\n\n60\n\n\n60\n\n\n400\n\n\n400\n\n\n400\n\n\nT\n\n\n-18 \u00b0C (0\u00b0F)\n\n\n50\n\n\nT\n\n\n50\n\n\n50\n\n\n300\n\n\n300\n\n\n300\n\n\nT\n\n\n40\n40\n\n\n40\n\n\n40\n10\n\n\nT\n\n\n--\n\n\nRT\n\n\n0 \u00b0C (32\u00b0F)\n\n\n30\n30\n\n\n30\n\n\n30\n\n\n200\n\n\n200\n\n\n200\n\n\n-18 \u00b0C (0 \u00b0F)\n\n\n1 year\n\n\n-\n\n\n1 year\n\n\n1 year\n\n\n20\n20\n\n\n20\n\n\n2 months\n\n\n1 week\n\u2610 1\n102\n\n\n30 min\n\n\n1 day 1 week 2 months\n\n\n1 day\n1\n\n\n30 min\n\n\n1 week 2 months\n\n\n30 min\n\n\n1 day\n1\n\n\n\u2610\n103\n\n\n104\n24\n20\n\n\n| L\n102\n\n\n11\n103\n\n\n100\n\n\n11\n103\n\n\n100\n\n\n100\n\n\n104\n\n\n104\n\n\n102\n\n\n1\n\n\n10\n\n\n10\n\n\n0.1\n\n\n1\n\n\n0.1\n\n\n1\n\n\n10\n\n\n0.1\n\n\nElapsed time after quenching, h\n\n\nElapsed time quenching, h\n\n\nElapsed time after quenching, h\n\n\n500\n\n\n500\n\n\nT\n\n\n500\n\n\n70\n0\n\n\n70\n\n\n70\n\n\n60\n60\n\n\n660\n\n\nT\n\n\n60\n\n\n400\n\n\n400\n\n\n400\n\n\n50\n50\n\n\nT\n\n\n650\n\n\nT\n\n\nYield strength, ksi\n\n\n50\n\n\nYield strength, MPa\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\nYield strength, ksi\n\n\nYield strength, MPa\n\n\nRT\n\n\n300\n\n\n300\n\n\n300\n\n\n40\n40\n\n\nRT\n\n\nT\n\n\nT\n\n\nT\n\n\n40\n\n\n0 \u00b0C (32\u00b0F)\n\n\n40\n\n\n30\n\n\n200\n\n\n30\n\n\n200\n\n\n30\n\n\n200\n\n\nT\n\n\n20\n20\n\n\n20\n\n\n20\n\n\nRT\n\n\n-18 \u00b0C (0 \u00b0F)\n\n\n100\n\n\n0 \u00b0C (32 F:\n\n\n100\n\n\n100\n\n\n-18 \u00b0C (0 \u00b0F) 1 year 10.\n2 months\n\n\n10\n\n\n1 year\n\n\n10\n\n\n1 year\n\n\n1 day\n\n\n1 week 2 months\n\n\n1 day 1 week\n\n\n30 min\n\n\n30 min\n\n\n\u2610\n102\n\n\n11\n\n\n\u2610\n103\n\n\n1 week 2 months\n\n\n30 min\n\n\n1 day\n1\n\n\n\u2610\u2610\n10\n\n\n11\n102\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n11\n102\n\n\n0\n\n\n104\n\n\n10\n\n\n104\n\n\n0.1\n\n\n1\n\n\n103\n\n\n1\n\n\n0.1\n\n\n104\n\n\n103\n\n\n1\n\n\n10\n\n\n0.1\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\n40\n\n\n40\n\n\n40\n\n\nElongation, % in 50 mm or 2 in.\n\n\nElongation, \" in 50 mm or 2 in.\n\n\nElongation, % in 50 mm or 2 in.\n\n\n30\n\n\n30\n\n\n-18 \u00b0C (0\u00b0F)\n0 \u00b0C (32\u00b0F)\n\n\n30\n\n\nRT\n\n\n-18 \u00b0C (0 \u00b0F)\n\n\nRT\n\n\n20\n\n\n20\n\n\n0 \u00b0C (32\u00b0F)\n\n\nRT\n\n\n10\n\n\n10\n\n\n1 year\n\n\n1 year\n\n\n1 year\n\n\n30 min\n\n\n1 week 2 months.\n\n\n1 day\nT\n\n\n1 week 2 months\n\u2610\n103\n\n\n30 min\n\n\n1 day\n\n\n\u2610 1\n103\n\n\n\u2610\n\n\n1\n102\n\n\n1 week 2 months\n\u2610\n102\n\n\n1 day\n1\n\n\n0\n0.1\n\n\n30 min\n\n\n0\n\n\n104\n\n\n104\n\n\n102\n\n\n10\n\n\n1\n\n\n0\n\n\n1\n\n\n10\n\n\n0.1\n\n\n104\n\n\n103\n\n\n0.1\n\n\n1\n\n\n10\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\nElapsed time after quenching, h\n\n\nFig 21 (continued)\n\n\n18 090\n)\nTK\n\n\ndt\nFys\n\n\nS\n\n\nFys 1.45\u00d710-16\n\n\n(Eq 7b)\n\n\nThe effects of neglecting to compensate\nfor soaking at temperatures other than the\nnominal can be large (Fig 25). For exam-\nple, the calculated difference in strength\nbetween alloy 7050 extrusions soaked 29 h\nat 160 \u00b0C (320 \u00b0F) and at 165 \u00b0C (325 \u00b0F) is\nabout 50 MPa (7 ksi), and the calculated\ndifference in strength between 7050 extru-\nsions soaked 29 h at 155 \u00b0C (315 \u00b0F) and at\n170 \u00b0C (335 \u00b0F) is about 100 MPa (14 ksi).\nNeglecting to compensate for time spent\nheating the work to the soak temperature\nwill increase the variability. Strength loss\nattributed to heatup was 14 MPa (2 ksi).\nThese kinetic relationships also can as-\nsist in selection of equivalent aging times\nfor alternate second-step aging tempera-\ntures. Equations 5 and 7 can be rearranged\n\n\n(Eq 6)\n\n\nexp (\n\n\n0=\n\n\nwhere t is time during heating.\nEquation 5 provides the basis for selec-\ntion of a nominal aging time that will result\nin the desired yield strength and gives the\nfurnace operator a method of compensating\nfor heating rate and for differences between\ndesired and attained soak temperatures.\nSpecifics will be illustrated using data for\nalloy 7050. The value of Fys (in units of\nhours) for 7050 can be calculated by the\nfollowing equation:\n\n\nwhere Tk is temperature in K.\n\n\nIn one experiment, lengths of 7050-W (4\ndays) extrusions were aged at 24 h at 120\n\u00b0C (250 \u00b0F) plus the equivalent of 3 to 42 h\nat 165 \u00b0C (325 \u00b0F). For the second step, a\nlogarithmic heatup was used in which 10 h\nwere required for the load to reach 155 \u00b0C\n(315 \u00b0F), and nominal soak temperature\nwas 165 \u00b0C (325 \u00b0F). Figure 24 indicates\nthat yield strength generally agreed with\nvalues predicted using Eq 5. The deviation\nof the curve for short-transverse strength\nat the short aging times indicates that the\nmethod is inadequate for predicting\nstrength on the underaging side of the\naging curve.\n\n\n32 562\nTF+460\n\n\n(Eq 7a)\n\n\nFys 1.45\u00d710-16 exp (\n\n\n)\n\n\nwhere T is temperature in \u00b0F, or\n\n\n864 / Heat Treating of Nonferrous Alloys\n\n\n50\n\n\n550\n\n\n\u2610\n\n\n500\n\n\n500\n\n\n105 C\n\n\nRT\n\n\n100 C\n(212 F)\n\n\n(225 F) -70\n\n\n70\n\n\n450\n\n\n150 C\n\n\n450\n\n\nNW(320F) (300)\n\n\nTensile strength, MPa\n\n\nTensile strength, MPa\n\n\nTensile strength, ksi\n\n\nTensile strength, ksi\n\n\n60\n\n\n60\n\n\n135 \"C\n(275 F)\n\n\n400\n\n\n400\n\n\n190 C\n(375 \u00b0F)\n\n\n150 \u00b0C\n(300 \"F)\n\n\n50\n\n\nIT\n\n\n350\n\n\n350\n\n\n175 \"C\n(350 (F)\n\n\n300\n\n\n300\n\n\n205 C\n(400 \u00b0F)\n\n\n190 \"C\n(375 \u00b0F)\n205 \"C\n(400 \"F)\n\n\n40\n\n\n260 \u00b0C\n(500 \u02daF)\n\n\n260 \"C\n(500 \u00b0F)\n\n\nT\n\n\n40\n\n\n250\n\n\n250\n\n\n230 \"C\n(450 \"F)\n\n\n1 day 1 week 2 months 1 year\n\n\n30 min\n\n\n00\n200\n\n\n30\n\n\n30 min\n\n\n1 day\n\n\n1 week 2 months 1 year.\n\n\n1|\n104\n\n\n1\n\n\n30\n\n\n200\n\n\n105\n\n\n103\n\n\n100\n\n\n105\n\n\n0\n\n\n0.01\n\n\n0.1\n\n\n1\n\n\n10\n\n\n103\n\n\n104\n\n\n1\n\n\n100\n\n\n0\n\n\n0.01\n\n\n0.1\n\n\n10\n\n\nAging time, h\n\n\nDuration of precipitation heat treatment, h\n\n\n500\n\n\n500\n\n\n175 C\n(350 \u00b0F)\n\n\n70\n10\n\n\n130 \u00b0C (265\u00b0F) 150 \u00b0C\n\n\n70\n\n\n(300 \u02daF)\n\n\n105 \u00b0C\n(225\u00b0F)\n\n\n60\n60\n\n\n60\n\n\n400\n\n\n400\n\n\n+\n135\u00b0C\n(275 \u00b0F)\n\n\nRT\n\n\n100\u00b0C\n(212 \"F)\n\n\nYield strength, MPa\n\n\n50\n\n\n50\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\nYield strength, ksi\n\n\n160 \u00b0C\n(320 \u00b0F)\n\n\n300\n\n\n300\n\n\n40\n40\n\n\n190 \u00b0C\n(375 \"F)\n\n\n150 \u00b0C\n(300 \u00b0F)\n\n\n40\n\n\n175 \u00b0C (350 \u00b0F) - 30\n\n\n30\n\n\n200\n\n\n200\n\n\n205 C\n(400 F)\n\n\n190 \u00b0C (375 \u00b0F)\n-205 \u00b0C (400 \u00b0F)\n\n\n20\n\n\n20\n\n\n260 \u00b0C (500 \u00b0F).\n\n\n100\n00\n\n\n230 \u00b0C (450 \u00b0F).\n\n\n100\n\n\n260 \u02daC\n(500 \"F)\n\n\n10\n\n\n10\n\n\n1 day 1 week 2 months 1 year\n\n\n30 min\n1\n\n\n1 day 1 week 2 months 1 year\n\n\n30 min\n\n\n\u2610\n103\n\n\n\u2610\n\n\n11\n103\n\n\n1\n104\n\n\n0\n\n\n0\n\n\n0\n\n\n0\n\n\n105\n\n\n105\n\n\n104\n\n\n0\n\n\n100\n\n\n0.1\n\n\n1\n\n\n10\n\n\n0\n\n\n0.1\n\n\n100\n\n\n0.01\n\n\n0.01\n\n\n1\n\n\n10\n\n\nDuration of precipitation heat treatment, h\n\n\nAging time, h\n\n\n40\n\n\n20\n\n\nElongation, % in 50 mm or 2 in.\n\n\n100 C\n(212 \u00b0F)\n\n\nRT\n\n\nElongation, % in 50 mm or 2 in.\n\n\n190 \u00b0C\n(375 \u00b0F)\n\n\n30 205 \u00b0C\n\n\n175 \u00b0C\n(350 \u00b0F)\n\n\n150 \"C 135 C\n(300 \u02daF) (275 \u00b0F)\n\n\n15\n\n\n(400 \u00b0F)\n\n\n130 C\n(265\u00b0F)\n\n\n205 C\n(400 F)\n\n\n260 \u00b0C\n(500 \u00b0F)\n\n\n20\n\n\n10\n\n\n150 C\n(300 F)\n\n\n105 C\n(225 FI\n\n\n190 \u00b0C 1\n(375 \"F)\n\n\n160 \u00b0C\n\u2020175 C\n(350 F) (320 \u02daF)\n\n\n5\n\n\n10\n\n\n260 \u00b0C\n(500 \u00b0F)\n\n\n230 C\n(450 \u00b0F)\n\n\n1 day 1 week 2 months 1 year\n\n\n30 min\n1\n\n\n1 week 2 months 1 year\n\u2610 1\n103\n\n\n1 day\n\n\n30 min\n\n\n11\n103\n\n\n1\n\n\n\u2610\n104\n\n\n0\n0\n\n\n0\n\n\n105\n\n\n104\n\n\n105\n\n\n0.01\n\n\n0.1\n\n\n10\n\n\n100\n\n\n100\n\n\n1\n\n\n0.01\n\n\n0.1\n\n\n1\n\n\n10\n\n\n0\n\n\nAging time, h\n\n\nDuration of precipitation heat treatment, h\n\n\nFig 22(a) Aging characteristics of alloy 2014 sheet\n\n\nFig 22(b) Aging characteristics of alloy 2024 sheet (see also Fig 22d)\n\n\nquenching, whereas other alloys show little\nor no added strengthening when treated by\nthis sequence of operations.\nAlloys of the 2xxx series such as 2014,\n2124, and 2219 are particularly responsive\nto cold work between quenching and aging,\nand this characteristic is the basis for the\nhigher-strength T8 tempers. The strength\nimprovement accruing from the combina-\ntion of cold working and precipitation heat\ntreating is a result of nucleation of addition-\n\n\nto yield the following equation:\n\n\n1350 29/exp (1.28)=29/3.6=8 h\nThermomechanical effects on aging occur\nfrom deformation after solution heat treat-\nment. The deformation step may be warm\nor cold and before, after, or during aging.\nThe simplest thermomechanical practices\nare those of the conventional T3, T8, or T9\ntempers. The rate and extent of precipita-\ntion strengthening are distinctly increased\nin some alloys by cold working after\n\n\n32 562 32 562\nT\u2081+460 T\u2082+460\n\n\n(Eq 8)\n\n\nt2=t\u2081 exp\n\n\nwhere 11 is aging time at temperature T\u2081, t\u2082 is\naging time at temperature T\u2082 that will provide\nequivalent yield strength, and T, and T\u2082 are in\n\u00b0F. For example, the time at 175 \u00b0C (350 \u00b0F)\nequivalent to aging alloy 7050 for 29 h at 165\n\u00b0C (325 \u00b0F) is calculated as follows:\n\n\nHeat Treating of Aluminum Alloys / 865\n\n\ntype of approach results in strengths sim-\nilar to those obtained with T8 processing\nbut with the better toughness and fatigue\ncharacteristics of T3 products. Alloys\n2024, 2124, and 2219 in T8-type tempers\nare particularly well suited for supersonic\nand military aircraft; alloy 2219 in such\ntempers, and alloy 2014-T65, were the\nprincipal materials for the fuel and oxidiz-\ner tanks (which also served as the primary\nstructure) of the Saturn V space vehicles.\nRe-solution heat treatment of mill prod-\nucts supplied in these tempers can result in\ngrain growth and in substantially lower\nstrength than is normal for the original\ntemper. Such reheat treatment is not rec-\nommended.\nAlloys of the 7xxx series do not respond\nfavorably to the sequence of operations\nused to produce T8-type tempers, and no\nsuch tempers are standard for these alloys.\nThe strains associated with stretching or\ncompressing of 7xxx alloys have relatively\nlittle effect on the mechanical properties of\nmaterial precipitation heat treated to T6-\ntype tempers. On the other hand, these\noperations have measurable detrimental\neffects on final strength when T73-, T736,\nor T76-type tempers are produced, partic-\nularly in the direction opposite the direc-\ntion of cold work. Accordingly, specifica-\ntion properties are somewhat lower for\nthe stress-relieved versions of these tem-\npers. Decreasing the overaging time to\ncompensate for the loss in strength is not\nadvisable, because this would impair de-\nvelopment of the desired corrosion char-\nacteristics.\nTemperature control and uniformity pre-\nsent essentially the same problems in pre-\ncipitation heat treating as they do in solu-\ntion heat treating.\nGood temperature control and uniformity\nthroughout the furnace and load are re-\nquired for all precipitation heat treating.\nRecommended temperatures are generally\nthose that are least critical and that can be\nused with practical time cycles. Except for\n7xxx alloys in T7x tempers, these tempera-\ntures generally allow some latitude and\nshould have a high probability of meeting\nproperty specification requirements. Fur-\nnace radiation effects seldom are trouble-\nsome except in those few furnaces that are\nused for both solution and precipitation heat\ntreating. Generally, such situations should\nbe avoided, because the high heat capacity\nneeded for the higher temperatures may be\ndifficult to control at normal aging temper-\n\n\n350\n\n\n325\n\n\n(250 \u00b0F)\n\n\n445\n\n\n300\n\n\n150 \u00b0C\n(300 \u00b0F)\n\n\nTensile strength, MPa\n\n\nTensile strength, ksi\n\n\n40\n\n\n275\n\n\n170 \u00b0C\n(340 \u00b0F)\n\n\n250\n\n\n35\n35\n\n\n205 \u00b0C\n(400 \u00b0F)\n\n\n225\n\n\n230 \u00b0C\n(450 \u00b0F)\n\n\n30\n\n\n200\n\n\n260 \u00b0C\n(500 \u00b0F)\n\n\n1 day\n\n\n30 min\n\n\n1 week 2 months 1 year\n\u2610 1\n103\n\n\nJ\n\n\n175\n\n\n104\n\n\n105\n\n\n1\n\n\n0\n\n\n0.01\n\n\n0.1\n\n\n10\n\n\n100\n\n\nDuration of precipitation heat treatment, h\n\n\n300\n\n\n120 \u00b0C\n(250 \u00b0F)\n\n\n40\n\n\n150 \u00b0C\n(300 \u00b0F)\n\n\n250\n\n\n35\n\n\n170 \u00b0C\n(340 \u00b0C)\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\n30\n\n\n200\n\n\n205 \u00b0C\n(400 \u00b0F)\n\n\n25\n\n\n1\n\n\n150\n\n\n230 \u00b0C\n(450 \u00b0F)\n\n\n20\n20\n\n\n15\n\n\n100\n\n\n260 \u00b0C\n(500 \u00b0F)\n\n\n30 min\n\n\n1 week 2 months 1 year\n\n\n1 day\n\n\n10\n\n\n50\n\n\n103\n\n\n104\n\n\n105\n\n\n0.01\n\n\n0\n\n\n1\n\n\n100\n\n\n0.1\n\n\n10\n\n\nDuration of precipitation heat treatment, h\n\n\n40\n\n\nElongation, % in 50 mm or 2 in.\n\n\n170 \u00b0C\n(340 \u00b0F)\n\n\n_ 150 \u00b0C\n\n\n230 \u00b0C 205 \u00b0C\n(450 \u00b0F) (400 \u00b0F)\n\n\n30\n\n\n(300 \u00b0F)\n\n\n260 \u00b0C\n(500 \u00b0F)\n\n\n201\n\n\n120\u00b0C\n(250 \u00b0F)\n\n\n10-260 \u00b0C\n\n\n(500 \u00b0F)\n\n\n30 min\n\n\n1 day 1 week 2 months 1 year\n\n\n11\n104\n\n\n0\n\n\n10\n1\n103\nDuration of precipitation heat treatment, h\n\n\n105\n\n\n0\n\n\n0.01\n\n\n100\n\n\n0.1\n\n\nFig 22(c) Aging characteristics of alloy 6061 sheet\n\n\nal precipitate particles by the increased\nstrain. In some alloys of the 2xxx series,\nstrain introduced by cold working after so-\nlution heat treatment and quenching also\ninduces nucleation of a finer precipitate\ndispersion that increases strength. Depend-\ning on the aging temper, however, tough-\nness may be adversely affected, as illustrat-\ned in Fig 26 for 2024 sheet.\nStrengthening from thermomechanical\nprocessing is the basis for the higher-\n\n\nstrength T8-type tempers of alloys 2011,\n2024, 2124, 2219, and 2419, which are\nproduced by applying controlled amounts\nof cold rolling, stretching, or combinations\nof these operations. Normally, cold work\nis introduced by stretching; however, oth-\ner methods such as cold rolling can be\nused. Recently, 2324-T39 was developed.\nThe T39 temper is obtained by cold rolling\napproximately 10% after quenching fol-\nlowed by stretching to stress relieve. This\n\n\natures.\n\n\nSoak time in precipitation heat treating is\nnot difficult to control; the specified times\ncarry rather broad tolerances. Heavier loads\nwith parts racked closer together, and even\nnested, are not abnormal. The principal haz-\nard is undersoaking due to gross excesses in\nloading practices. Some regions of the load\nmay reach soak temperature long after soak\n\n\n866 / Heat Treating of Nonferrous Alloys\n\n\n2024-T3\n(cold worked 1 to 2%)\n\n\n2024-T36\n(cold worked 5 to 6%)\n\n\n2024-T4\n(not cold worked)\n\n\n600\n\n\n600\n\n\n600\n\n\n190 \u00b0C (375 \u00b0F) 175 \u00b0C (350 \u00b0F)-80\n\n\n80\n\n\n80\n\n\nTensile strength, MPa\n\n\nTensile strength, MPa\n\n\n190 \u00b0C (375 \u00b0F)\n\n\nTensile strength, ksi\n\n\nTensile strength, MPa\n\n\nTensile strength, ksi\n\n\nTensile strength, ksi\n\n\n190 \u00b0C (375 \u00b0F).\n\n\n000\n500\n\n\n500\n\n\n500\n\n\nF205 \u00b0C (400 \u00b0F)\n\n\n70\n\n\n70\n\n\n70\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n60\n60\n\n\n60\n60\n\n\n60\n\n\n400\n\n\n400\n\n\n400\n\n\n50\n50\n\n\n50\n50\n\n\n220 \u00b0C (425 \u00b0F) - 50\n\n\nT\n\n\n300\n\n\n300\n\n\n300\n\n\n0\n\n\n4\n\n\n8\n\n\n12\n\n\n16\n\n\n0\n\n\n4\n\n\n8\n\n\n12\n\n\n0\n\n\n12\n\n\n16\n\n\n4\n\n\n8\n\n\n16\n\n\nAging time, h\n\n\nAging time, h\n\n\nAging time, h\n\n\n600\n\n\n600\n\n\n600\n\n\n80\n80\n\n\n175 \u00b0C (350 \u00b0F) - 80\n\n\n80\n\n\n190 \u00b0C (375 \u00b0F)\n\n\n500\n\n\n500\n\n\n500\n\n\nYield strength, MPa\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\n70\n\n\nYield strength, ksi\n\n\n70\n\n\nYield strength, MPa\n\n\n70\n\n\nYield strength, ksi\n\n\n190 \u00b0C (375 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n60\n\n\n60\n\n\n60\n\n\n190 \u00b0C (375 \u00b0F)-\n\n\n400\n\n\n400\n\n\n400\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n50\n50\n\n\nT\n\n\n220 \u00b0C (425 \u00b0F)\n\n\nT\n\n\n50\n\n\n300\n\n\n300\n\n\n300\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n40\n\n\n- 40\n\n\n40\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n30\n\n\n30\n\n\n30\n\n\n200\n\n\n200\n\n\n200\n\n\n0\n\n\n4\n\n\n8\n\n\n12\n\n\n0\n\n\n8\n\n\n16\n\n\n4\n\n\n12\n\n\n4\n\n\n8\n\n\n12\n\n\n16\n\n\n0\n\n\n16\n\n\nAging time, hi\n\n\nAging time, h\n\n\nAging time, h\n\n\n30\n\n\n30\n\n\n30\n\n\nElongation, % in\n\n\nElongation, % in\n\n\nElongation, % in\n50 mm or 2 in.\n\n\n50 mm or 2 in.\n\n\n50 mm or 2 in.\n\n\n190 \u00b0C (375 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n190 \u00b0C (375 \u00b0F)\n\n\n190 \u00b0C (375 \u00b0F)\n\n\n175 \u00b0C (350 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n220 \u00b0C (425 \u00b0F)\n\n\n0\n\n\n8\n\n\n4\n\n\n12\n\n\n16\n\n\n0\n\n\n4\n\n\n8\n\n\n12\n\n\n4\n\n\n8\n\n\n12\n\n\n16\n\n\n16\n\n\n0\n\n\nAging time, h\n\n\nAging time, h\n\n\nAging time, h\n\n\nFig 22(d) Effects of cold work after quenching and before aging on tensile properties of alloy 2024 sheet\n\n\nto the T73 and T76 tempers. As discussed\nabove, soak time is not as critical for peak-\naged (T6 and T8) tempers.\n\n\nHardening of Cast Alloys\n\n\ntime has been called. Placement of load ther-\nmocouples is critical, and limiting the size and\nspacing of a load may be necessary for aging\n\n\nIn general, the principles and procedures\nfor heat treating wrought and cast alloys are\nsimilar. The major differences between so-\nlution-treating conditions for castings and\nthose for wrought products are found in\nsoak times and quenching media. Solution\nof the relatively large microconstituents\npresent in castings requires longer soaking\n\n\n225\n\n\n360\n\n\n385\n\n\n400\n\n\n200\n\n\n410\n\n\nMPa\n\n\nAging temperature, \u00b0C\n\n\nAging temperature, \u00b0F\n\n\n440\n\n\n470\n\n\n350\n\n\n175\n\n\n700\n\n\n500\n\n\nT73 aging\n\n\n100\n\n\nYield strength, MPa\n\n\n525\n\n\nYield strength, ksi\n\n\nLongitudinal\n\n\n60\n\n\n550\n\n\n90\n\n\n150\n\n\n600\n\n\n300\n\n\n64\n\n\n70\n88\n80\n\n\nShort transverse\n\n\n68\n\n\n500\n\n\n125\n\n\nPredicted from Eq 3\n\n\n72-\n\n\n250\n\n\nksi\n\n\nT6 aging\n\n\n\u043e\n\n\n\u2022 Actual\n\n\n76\n\n\n60\n\n\n400\n\n\n]\n\n\n80\n\n\n4 6 8 10\n\n\n100\n\n\n20 30 40 60\n\n\n2\n\n\nEquivalent aging time\nat 165 \u00b0C (325 \u00b0F), h\n\n\n0.01\n\n\n0.1\n\n\n10\n\n\n100\n\n\n1000\n\n\n1\n\n\nAging time, h\n\n\nActual versus predicted yield strengths for\nalloy 7050 extrusions\n\n\nFig 24\n\n\nFig 23 Iso-yield-strength curves for alloy 7075\n\n\nHeat Treating of Aluminum Alloys / 867\n\n\nTransverse yield strength, ksi\n\n\nnot by the heat treater. This effect of casting\nmethods on property development is shown\nin Fig 27. Because of the finer cast structure\nand higher supersaturation of the more rap-\nidly solidified permanent mold castings,\ntheir tensile properties are superior to those\nof sand castings of the same composition\nsimilarly heat treated.\nTempers. Cast products of heat-treatable\naluminum alloys have the highest combina-\ntions of strength, ductility, and toughness\nwhen produced in T6-type tempers. Devel-\noping T6-type tempers in cast products re-\nquires the same sequence of operations\nemployed in developing tempers of the\nsame type in wrought products-solution\nheat treating, quenching, and precipitation\nheat treating. Premium-quality casting spec-\nifications such as MIL-A-21180 can require\ndifferent strengths and ductility levels in the\nsame casting.\nAmong precipitation treatments unique to\ncastings are those resulting in the T5 and T7\ntempers. The T5 temper is produced merely\nby applying a precipitation treatment to the\nas-cast casting, without previous solution\ntreatment. A moderate increase in strength\nis achieved without warpage and subse-\nquent straightening. High hardness and di-\nmensional and strength stability at elevated\ntemperatures account for the almost univer-\nsal use of materials in T5 tempers for pis-\ntons and other engine parts. Some applica-\ntions demand combinations of strength,\ntoughness, and dimensional stability that\ncannot be met by heat treating to T5-, T6-,\nor T8-type tempers. For these applications,\nT7-type tempers are developed by solution\nheat treating, quenching in a medium that\nprovides a moderate cooling rate, and then\nprecipitation heat treating at a temperature\nhigher than those used to develop T5-, T6-,\nand T8-type tempers. Heat treating to T7-\ntype tempers results in lower strength than\nthat of material in T6- or T8-type tempers,\ndevelops high ductility and toughness, and\ncarries precipitation far enough to minimize\nfurther precipitation during service.\n\n\nAging temperature, \u00b0F\n\n\n40 45 50 55 60 65 70\n\n\n310\n\n\n320\n\n\n330\n\n\n340\n\n\n1.8\n\n\n100\n\n\n5000\n\n\n10\n10\n\n\nFe Si\nCu Mg Mn\n4.2 1.4 0.6 0.34 0.14\n\n\nEffect on yield strength, MPa\n\n\nT4\n\n\n50\n50\n\n\nEffect on yield strength, ksi\n\n\nTear strength/Yield strength\n\n\n4500\n\n\n1.6\n\n\nEstimated K, MPav mm\n\n\n5\n\n\nT3\n\n\nStretched 5%\n\n\nT7\n\n\n4000\n\n\nC\n\n\n0\n\n\n1.4\n\n\nG\n\n\n3500\n\n\n-50\n\n\n1.2\n\n\nT6\n\n\nCompared for standard\n\n\n3000\n\n\nNot stretched\n\n\nconditions of 29 h\n\n\n-100\n\n\n-15\n\n\nat 165 \u00b0C\n\n\n1.0\n\n\n2500\n\n\n-20\n\n\n-150\n\n\nT7\n\n\nT8)\n\n\n160 165 170 175\n\n\n150\n\n\n155\n\n\n2000\n\n\n0.8\n\n\nAging temperature, \u00b0C\nEffect of aging temperature on yield strength\nFig 25 of alloy 7050-1736\n\n\n300 350 400 450 500\nTransverse yield strength, MPa\n\n\n250\n\n\nEffect of stretching and aging on the tough-\nness and yield strength of 2024 sheet\n\n\nFig 26\n\n\nperiods than those used for wrought prod-\nucts (Table 3). When heat treatment of\ncastings must be repeated, solution times\nbecome similar to those for wrought prod-\nucts, because the gross solution and homog-\nenization has been accomplished and is\nirreversible under normal conditions. Re-\nduction of stresses and distortion from\nquenching are also important, because cast-\nings generally are complex shapes with vari-\nations in section thickness.\nDifferent casting processes and foundry\npractices also result in microstructural dif-\nferences with relevance to heat-treatment\npractice, because the coarser microstruc-\ntures associated with slow solidification\nrates require a longer solution heat treat-\nment exposure. Therefore, the time re-\nquired at temperature to achieve solution is\nprogressively shorter for investment, sand,\nand permanent mold castings. Foundry\npractice (chills, gating, type of mold) also\nplays an important role in the response of a\ncasting, or a portion of a casting, to heat\ntreatment. For example, thin-wall sand\ncastings produced with extensive use of\nchills can often display finer microstruc-\ntures than heavy-section permanent mold\nparts produced in such a way that process\nadvantages are not exploited.\nFor these reasons, solution heat-treat-\nment practices can be optimized for any\nspecific part to achieve solution with the\nshortest reasonable cycle once production\npractice is finalized, even though most\nfoundries and heat treaters will standardize\na practice with a large margin of safety.\nThere also exists a fundamental difference\nbetween unmodified and modified alloys in\nwhich the size and shape of silicon crystals\nare modified with additions of elements\nsuch as calcium, sodium, strontium, or an-\ntimony. Modified alloys undergo rapid\nspheroidization while complete sphe-\nroidization is not achieved in unmodified\nalloys even after very long times. The prac-\ntical implication is that shorter solution heat\ntreatment could be employed in fully mod-\n\n\nified castings. The microsegregation of sili-\ncon and magnesium is not severe in the\naluminum-silicon-magnesium casting al-\nloys, and hence it takes only a short time to\nhomogenize the alloy and to place the\nMg2Si into solution.\nQuenchants. Quenching of aluminum\ncastings is often done in boiling water or a\nmilder medium to reduce quenching stress-\nes in complex shapes. A commercially im-\nportant variety is a mixture of polyalkylene\nglycol and water, which has no detrimental\neffect on properties for thicknesses under\napproximately 3.2 mm (0.125 in.). Quen-\nchant additions can be made for the follow-\ning purposes:\n\n\n\u2022 To promote stable vapor film boiling by\nthe deposition of compounds on the sur-\nface of parts as they are submerged in the\nquench solution\n\u2022 To suppress variations in heat flux by\nincreasing vapor film boiling stability\nthrough chemically decreased quench so-\nlution surface tension\n\u2022 To moderate quench rate for a given water\ntemperature\n\n\nThe key to the compromise between\ngoals involving property development and\nthe physical consequences of quenching is\nuniformity of heat extraction, which is in\nturn a complex function of the operable heat\nextraction mechanism. Nucleate, vapor\nfilm, and convective boiling occur with dra-\nmatically different heat extraction rates at\ndifferent intervals. Differences in section\nthickness, load density, positioning, racking\nmethods, surface condition, and casting ge-\nometry also influence the results.\nProperty Development, Yield strength is\nlargely controlled by the limiting hardening-\nelement level, and tensile strength (in a\ngeneral sense) is related to the ductility at a\ngiven yield strength. Ductility, however, is\ncontrolled for a given yield strength by\nsoundness and microstructural fineness,\nand is thus determined in the foundry and\n\n\nStress Relief\n\n\nImmediately after being quenched, most\naluminum alloys are nearly as ductile as\nthey are in the annealed condition. Conse-\nquently, it is often advantageous to stress\nrelieve parts by working the metal immedi-\nately after quenching. Numerous attempts\nalso have been made to develop a thermal\ntreatment that will remove, or appreciably\nreduce, quenching stresses. Normal precip-\nitation heat-treating temperatures are gener-\nally too low to provide appreciable stress\nrelief. Exposure to higher temperatures (at\nwhich stresses are relieved more effective-\nly) results in lower properties. However,\nsuch treatments are sometimes utilized\nwhen even moderate reduction of residual\nstress levels is important enough so that\n\n\n868 / Heat Treating of Nonferrous Alloys\n\n\n7079\nSpar forging\n\n\nPermanent mold castings\n\n\nSand castings\n\n\n350\n\n\n350\n\n\nPlanes of\nsaw cuts\n\n\n50\n\n\n50\n\n\nA\n\n\n300\n\n\n300\n\n\nTensile strength, MPa\n\n\n150 \u00b0C (300 \u00b0F)\n\n\nTensile strength, MPa\n\n\n-160 mm\n\n\n150 \u00b0C (300\u00b0F)\n\n\nTensile strength, ksi\n\n\nTensile strength, ksi\n\n\n40\n\n\n40\n\n\n\u00b0F)-\n\n\n205 \u00b0C (400)\n\n\n250\n\n\n250\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n30\n\n\n200\n\n\n200\n\n\n260 \u00b0C (500 \u00b0F)\n\n\nB\n\n\nB\n\n\n260 \u00b0C (500 \u00b0F)\n\n\n150\n\n\n150\n\n\nA\n\n\n20\n\n\n20\n\n\nDirection of\ncompressive deformation\n\n\n100\n\n\n100\n\n\n10\n\n\n20\n\n\n30\n\n\n20\n\n\n30\n\n\n0\n\n\n0\n\n\n10\n\n\n7500\n\n\nAging time, h\n\n\nAging time, h\n\n\nT6 parallel\n\n\nDeflection, 0.001 in.\n\n\nSaw cut A-A\n\n\n250\n\n\nDeflection, \u03bcm\n\n\n200\n\n\n250\n\n\n4500\n\n\n250\n\n\n150 \u00b0C (300 \u00b0F)\n\n\nT6 normal\n\n\n150\n\n\n100\n\n\n30\n\n\nYield strength, MPa\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\nYield strength, ksi\n\n\n-150 \u00b0C (300 \u00b0F)-\n\n\n200\n\n\n200\n\n\n1500\n\n\n50\n\n\n0\n\n\n150\n150\n\n\nT652 normal\n\n\n-205 \u00b0C (400 \u00b0F):\n\n\n150\n\n\nT652 parallel\n\n\n-1500\n\n\n-50\n\n\n20\n\n\n20\n\n\n205 \u00b0C (400 \u00b0F)\n\n\nT\n\n\nLength, 1.8 m\n\n\n260 \u00b0C (500 \u00b0F)\n\n\n260 \u00b0C (500 \u00b0F)\n\n\n100\n\n\n100\n\n\n2500\n\n\nDeflection, 0.001 in.\n\n\nSaw cut B-B\n\n\nT6 parallel\n\n\n10\n\n\n10\n\n\n50\n50\n\n\n80\n\n\nDeflection, um\n\n\n50\n\n\nT6 normal\n\n\n10\n\n\n20\n\n\n30\n\n\n0\n\n\n0\n\n\n10\n\n\n20\n\n\n30\n\n\n60\n\n\n1500\n\n\nAging time, h\n\n\nAging time, h\n\n\n40\n\n\nT652 normal\nT652 parallel\n\n\n20\n\n\n500\n\n\n1\n\n\n2\n\n\nin 50 mm or 2 in.\n\n\nElongation, %\nin 50 mm or 2 in.\n\n\n260 \u00b0C (500 \u00b0F)\n\n\nElongation, %\n\n\n0\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n260 \u00b0C (500 \u00b0F)\n\n\n-20\n\n\n-500\n\n\n-150 \u00b0C (300 \u00b0FH\n\n\n1\n\n\nLength, -1.8 m\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n150 \u00b0C (300 \u00b0F).\n\n\nEffect of 3% permanent deformation in com-\nFig 28\npression (T652 treatment) on distribution of\nstress in a large forging. Parallel and normal refer to\nwarpage directions with respect to the plane of the\n\n\n10\n\n\n20\n\n\n10\n\n\n30\n\n\n20\n\n\n30\n\n\nAging time, h\n\n\nAging time, h\n\n\nFig 27 Comparison of the precipitation-hardening characteristics of 356.0-T4 sand and permanent mold\n\n\nsaw cut.\n\n\ncastings\n\n\nsome sacrifice in mechanical properties can\nbe accepted. The T7 temper for castings is a\ntypical example of this kind of treatment.\nMechanical Stress Relief. Deformation\nconsists of stretching (bar, extrusions, and\nplate) or compressing (forgings) the product\nsufficiently to achieve a small but controlled\namount (1 to 3%) of plastic deformation. If\nthe benefits of mechanical stress relieving\nare needed, the user should refrain from\nreheat treating.\nFigure 28 illustrates the beneficial effect\nof 3% permanent deformation in compres-\nsion on a large forging.\nThese methods are most readily adapt-\nable to mill and forge shop products and\nrequire equipment of greater capacity than\nthat found in most manufacturing plants.\nApplication of these methods to die forgings\nand extrusions usually requires construc-\ntion of special dies and jaws. Stretching\ngenerally is limited to material of uniform\ncross section; however, it has been applied\nsuccessfully to stepped extrusions and to a\n3 by 14 m (10 by 47 ft) aircraft wing skin\nroll-tapered to a thickness range of 7.1 to\n3.2 mm (0.280 to 0.125 in.).\nSpecific combinations of the supplemen-\ntal digits are used to denote the tempers\nproduced when mechanical deformation is\nused primarily to relieve residual stresses\n\n\ninduced during the quenching operation.\nFor products stress relieved by stretching,\nthe digits 51 follow the basic Tx designation\n(T451, for example). For products stress\nrelieved by compressive deformation, the\nsupplementary digits are 52.\nAn additional digit is added to designa-\ntions for extrusions: an added zero specifies\nthat the product has not been straightened\nafter final stretching; an added one indicates\nthat straightening may have been performed\nafter final stretching.\nEffect of Precipitation Heat Treating on\nResidual Stress. The stresses developed dur-\ning quenching from solution heat treatment\nare reduced during subsequent precipitation\nheat treatment. The degree of relaxation of\nstresses is highly dependent upon the time\nand temperature of the precipitation treat-\nment and the alloy composition. In general,\nthe precipitation treatments used to obtain\nthe T6 tempers provide only modest reduc-\ntion in stresses, ranging from about 10 to\n35%. To achieve a substantial lowering of\nquenching stresses by thermal stress relax-\nation, higher-temperature treatments of the\nT7 type are required. These treatments are\nused when the lower strengths resulting\nfrom overaging are acceptable.\nOther thermal stress-relief treatments,\nknown as subzero treatment and cold stabi-\n\n\nlization, involve cycling of parts above and\nbelow room temperature. The temperatures\nchosen are those that can be readily ob-\ntained with boiling water and mixtures of\ndry ice and alcohol-namely, 100 and -73\n\u00b0C (212 and -100 \u00b0F)-and the number of\ncycles ranges from one to five. The maxi-\nmum reduction in residual stress that can be\neffected by these techniques is about 25%.\nThe maximum effect can be obtained only if\nthe subzero step is performed first, and\nimmediately after quenching from the solu-\ntion-treating temperature while yield\nstrength is low. No benefit is gained from\nmore than one cycle.\nA 25% reduction in residual stress is\nsometimes sufficient to permit fabrication\nof a part that could not be made without this\nreduction. However, if a general reduction\nis needed, as much as 83% relief of residual\nstress is possible by increasing the severity\nof the uphill quench\u2014that is, more closely\napproximating the reverse of the cooling-\nrate differential during the original quench.\nThis may be accomplished by a patented\nprocess that involves extending the subzero\nstep to 195 \u00b0C (-320 \u00b0F) and then very\nrapidly uphill quenching in a blast of live\nsteam (Fig 29). The rate of reheating is\nextremely critical, and therefore, to ensure\nproper application of the steam blast, a\n\n\nHeat Treating of Aluminum Alloys / 869\n\n\nTable 9 Reheating schedules for wrought aluminum alloys\nThe schedules given in this table normally will not decrease strength more than 5%.\n\n\nPlate\nthickness\n50 mm\n\n\nReheating time at a temperature of:\n\n\n150 \u00b0C\n(300 \u00b0F)\n\n\n165 \u00b0C\n(325 \u00b0F)\n\n\n175 \u00b0C\n(350 \u00b0F)\n\n\n220 \u00b0C\n(425 \u00b0F)\n\n\n190 \u00b0C\n(375 \u00b0F)\n\n\n205 \u00b0C\n(400 \u00b0F)\n\n\n230 \u00b0C\n(450 \u00b0F)\n\n\nAlloy and temper\n\n\nRange of\nresidual\nstress\nMPa ksi\n\n\n(a)\n8-10 h\n(a)\n\n\n2014-T4\n2014-T6\n\n\n(a)\n2-50 h\n(a)\n20-40 h\n\n\n(a)\n2-4 h\n(a)\n2-4 h\n\n\n(a)\n1/2-1 h\n(a)\n1 h\n\n\n(a)\n5-15 min\n(a)\n12 h\n\n\n(a)\n(b)\n(a)\n15 min\n\n\n(a)\n(b)\n(a)\n5 min\n\n\nA B C D\n\n\n2024-T3, 2024-T4\n\n\n2024-T81, 2024-T86\n6061-T6, 6062-T6,\n\n\nA 28\nB 86\nC 130\nD 165\n\n\n4.0\n12.4\n19.0\n24.0\n\n\n50-100 h\n1-2 h\n\n\n5 min\n(a)\n\n\n6063-T6\n\n\n100-200 h\n10-12 h\n\n\n1-2 h\n1/2-1 h\n\n\n15 min\n(b)\n\n\n8-10 h\n1-2 h\n\n\n1/2 h\n5-10 min\n\n\n7075-T6, 7178-T6\n\n\n(a) Reheating not recommended. (b) Bring to temperature\n\n\nTreatment\n\n\nA: Cooled to -195 \u00b0C, then uphill\nquenched in a steam blast\nB: Cooled to -75 \u00b0 C, then uphill\nquenched in a steam blast\n\n\ntemper without first carefully testing the\neffects of such reheating. In one such test,\n2024-T4 sheet was found to be very suscep-\ntible to intergranular corrosion when sub-\njected to a 15-min drying operation at 150 \u00b0C\n(300 \u00b0F) during the first 8 h after quenching;\nno susceptibility was evident when the\nsame drying operation was performed more\nthan 16 h after quenching. In another test,\n7075-W (0.2 to 600 h) bar and plate were\nreheated for hot forming at 175 \u00b0C (350 \u00b0F)\nfor 20 min. Strengths after aging to the T6\ntemper were 10 to 15% lower than those for\nstandard 7075-T6. In contrast, similar re-\nheating of T6 material for up to 1 h at 175 \u00b0C\n(350 \u00b0F) produced no detrimental effect.\nIf reheating is performed on material in\nthe W or T4 condition, its effect can be\nestimated from families of precipitation\nheat-treating curves such as those present-\ned in Fig 22. Such curves can also be used\nfor reheating of precipitation heat-treated\nmaterial at the precipitation heat-treating\ntemperature. For reheating at other temper-\natures, other data may be needed (Fig 31).\nThe heat-treating and reheating curves may\nbe used as the bases for limitations on\nreheating (Table 9).\n\n\nFor both heat-treatable and non-heat-treat-\nable aluminum alloys, reduction or elimina-\ntion of the strengthening effects of cold work-\ning is accomplished by heating at a\ntemperature from about 260 to about 440 \u00b0C\n(500 to 825 \u00b0F). The rate of softening is\nstrongly temperature-dependent; the time re-\nquired to soften a given material by a given\namount can vary from hours at low tempera-\ntures to seconds at high temperatures.\nIf the purpose of annealing is merely to\nremove the effects of strain hardening, heat-\ning to about 345 \u00b0C (650 \u00b0F) will usually\nsuffice. If it is necessary to remove the\nhardening effects of a heat treatment or of\ncooling from hot-working temperatures, a\ntreatment designed to produce a coarse,\nwidely spaced precipitate is employed. This\nusually consists of soaking at 415 to 440 \u00b0C\n(775 to 825 \u00b0F) followed by slow cooling (28\n\u00b0C/h, or 50 \u00b0F/h, max) to about 260 \u00b0C (500\n\u00b0F). The high diffusion rates that exist dur-\ning soaking and slow cooling permit maxi-\nmum coalescence of precipitate particles\nand result in minimum hardness.\nAs a result of this treatment, only partial\nprecipitation occurs in 7xxx alloys, and a\nsecond treatment (soaking at 230 \u00b1 6 \u00b0C, or\n450 \u00b1 10 \u00b0F, for 2 h) is required. When the\nneed arises for small additional improve-\nments in formability, cooling at 28 \u00b0C/h (50\n\u00b0F/h) should be extended to 230 \u00b0C (450 \u00b0F),\nand the material should be soaked at 230 \u00b0C\nfor 6 h. The effects of eliminating or pro-\nlonging the 230 \u00b0C second step on the duc-\ntility of 7075-0 sheet are compared with the\nstandard treatment in Table 10.\nIn annealing, it is important to ensure that\nthe proper temperature is reached in all\nportions of the load; therefore, it is common\nto specify a soaking period of at least 1 h.\nThe maximum annealing temperature is\nmoderately critical; it is advisable not to\nexceed 415 \u00b0C (775 \u00b0F), because of oxidation\nand grain growth. The heating rate can be\ncritical, especially for alloy 3003, which\nusually requires rapid heating for preven-\ntion of grain growth. Relatively slow cool-\ning, in still air or in the furnace, is recom-\nmended for all alloys to minimize distortion.\nTypical annealing conditions used for some\nalloys in common use are listed in Table 11.\n\n\nC: Cooled to -75 or -195 \u00b0 C, then\nuphill quenched in boiling water\n\n\nD: Standard specimen, quenched and\naged to T6 temper in conventional\nmanner with no further treatment\n\n\nEffectiveness of various uphill quenching\ntreatments in reducing residual quenching\nstresses in 2014 plate. Note: uphill quenching treat-\nments (single-cycle only) were applied from 1/2 to 11/2 h\nafter quenching from the recommended solution-\ntreating temperature. All specimens were aged to the\nT6 temper after uphill quenching.\n\n\nFig 29\n\n\nspecial fixture usually is required for each\npart.\nThis process will not solve all problems of\nwarpage in machining. It may reduce\nwarpage internally but increase warpage of\nthe extreme outer layers, although in the\nopposite direction (Fig 30). Also, the effect\nof the altered residual-stress pattern on per-\nformance must be evaluated carefully for\neach part. This is particularly important for\nparts subjected to cyclic loading or exposed\nto corrosive environments such as marine\natmospheres, especially if the process is\nintroduced after the start of production and\noriginal performance tests are not repeated.\nFurther disadvantages are the cost and haz-\nard involved in handling liquid nitrogen and\nlive steam.\n\n\nAnnealing\n\n\nAnnealing treatments employed for alu-\nminum alloys are of several types that differ\nin objective. Annealing times and tempera-\ntures depend on alloy type as well as on\ninitial structure and temper.\nFull Annealing. The softest, most ductile,\nand most workable condition of both non-\nheat-treatable and heat-treatable wrought\nalloys is produced by full annealing to the\ntemper designated \u201cO.\u201d Strain-hardened\nproducts in this temper normally become\nrecrystallized, but hot-worked products\nmay remain unrecrystallized. In the case of\nheat-treatable alloys, the solutes are suffi-\nciently thoroughly precipitated to prevent\nnatural age hardening. A higher maximum\ntemperature than that used for stress-relief\nannealing, controlled cooling to a lower\ntemperature, and additional holding time at\nthe lower temperature generally are em-\nployed.\n\n\nEffects of Reheating\n\n\nThe precipitation characteristics of alumi-\nnum alloys must be considered frequently\nduring evaluation of the effects of reheating\non mechanical properties and corrosion re-\nsistance. Such evaluations are necessary for\ndetermining standard practices for manu-\nfacturing operations, such as hot forming\nand straightening, adhesive bonding, and\npaint and dry-film lubricant curing, and for\nevaluating the effects of both short-term\nand long-term exposure in elevated temper-\natures in service.\nThe stage of precipitation that exists in an\nalloy at the time of reheating plays a signif-\nicant role in the effects of reheating. Con-\nsequently, it is extremely dangerous to re-\nheat material in a solution heat-treated\n\n\n870 / Heat Treating of Nonferrous Alloys\n\n\n2600\n\n\nTensile strength, MPa\n\n\n+100\n\n\n475\n\n\nTensile strength, ksi\n\n\n25-by-25-mm (1-by-1-in.) bar\n\n\n65\n99\n\n\n450\n\n\n+1300\n\n\n425\n\n\nO 120 \u00b0C\n\u2022 135 \u00b0C\nA 150 \u00b0C\n\n\n60\n\n\n400\n\n\nReheating\n\n\n0\n\n\n0\n\n\ntemperature\n\n\n375\n\n\n55\n\n\nDeflection, 0.001 in.\n\n\nDeflection, m\n\n\n425\n\n\n-1300\n\n\nYield strength, MPa\n\n\n60\n\n\nYield strength, ksi\n\n\n400\n\n\n55\n\n\n375\n\n\n-100\n\n\n-2600\n\n\n350\n\n\n50\n\n\nO 120 \u00b0C\n\n\n-3900\n\n\n325\n\n\n135 \u00b0C\n\n\nA 150 \u00b0C\n\n\n7075\n\n\n1\n\n\n300\n\n\no Control specimen\n\n\n50 mm (2 in), %\n\n\nQuenched from 75 \u00b0C -200\n(-100 \u00b0F) to steam\n\n\nElongation in\n\n\n-5200\n\n\n10\n\n\nA Quenched from liquid\nnitrogen to steam\n\n\n0\n\n\n-6500\n\n\n103\n\n\n10\n\n\n100\n\n\n1\n\n\n2\n\n\n3\n\n\n4\n\n\nDuration of reheating, days\n\n\nTine number\n\n\n50\n+50\n\n\nEffects of reheating on tensile properties of\nalclad 2024-T81 sheet\n\n\nFig 31\n\n\n+1300\n\n\n1 2 3 4 5 6\n\n\n0\n\n\n0\n\n\ntions, this ductility often is slightly lower\nthan that of material that has not been\nsubjected to prior heat treatment\u2014that is,\nmaterial annealed at the producing source.\nTherefore, when maximum ductility is re-\nquired, annealing of a previously heat-treat-\ned product is sometimes unsuccessful.\nPartial Annealing. Annealing of cold-\nworked non-heat-treatable wrought alloys to\nobtain intermediate mechanical properties\n(H2-type tempers) is referred to as partial\nannealing or recovery annealing. Tempera-\ntures used are below those that produce ex-\ntensive recrystallization, and incomplete soft-\nening is accomplished by substructural\nchanges in dislocation density and rearrange-\nment into cellular patterns (polygonization).\nBendability and formability of an alloy an-\nnealed to an H2-type temper generally are\nsignificantly higher than those of the same\nalloy in which an equal strength level is de-\nveloped by a final cold-working operation\n(H1-type temper). Treatments to produce H2-\ntype tempers require close control of temper-\nature to achieve uniform and consistent me-\nchanical properties.\nFigure 32 shows changes in yield strength\nas functions of temperature and time for\nsheet of two non-heat-treatable alloys (1100\nand 5052) initially in the highly cold-worked\ncondition (H18 temper). From these curves,\nit is apparent that, by selection of appropri-\nate combinations of time and temperature,\nmechanical properties intermediate to those\nof cold-worked and fully annealed material\ncan be obtained. It is also evident that yield\nstrength depends much more strongly on\ntemperature than on time of heating.\nStress-Relief Annealing. For cold-worked\nwrought alloys, annealing merely to remove\nthe effects of strain hardening is referred to\nas stress-relief annealing. Such treatments\n\n\n50-by-50-mm\n\n\n(2-by-2-in.) bar\n\n\nDeflection, 0.001 in.\n\n\nDeflection, um\n\n\n-50\n50\n\n\n-1300\n\n\n-100\n\n\n-2600\n\n\n7075\n\n\n-150\n\n\n-3900\n\n\n\u3002 Control specimen\n\n\n\u2022 Quenched from -75 \u00b0C (-100 \u00b0F) to steam\nA Quenched from liquid nitrogen to steam\n\n\n-200\n\n\n-5200\n\n\n3\n\n\n2\n\n\n6\n\n\n4\n\n\n5\n\n\nTine number\n\n\n2600\n\n\n100\n\n\n75-by-75-mm (3-by-3-in.) bar\n\n\n500\n+50\n\n\n+1300\n\n\nDeflection, 0.001 in.\n\n\nDeflection, um\n\n\n0\n\n\n-50\n\n\n-1300\n\n\n7075\n\n\n-100\n\n\n-2600\n\n\no Control specimen\n\n\nQuenched from -75 \u00b0C (-100 \u00b0F) to steam\n\n\nA Quenched from liquid nitrogen to steam\n\n\n-3900\n\n\nA-150\n\n\n1\n\n\n2\n\n\n3\n\n\n4\n\n\n5\n\n\n6\n\n\n7\n\n\n8\n\n\nTine number\n\n\nEffect of uphill quenching on deflection of tines. Six-tine specimen was machined from 50 by 50 mm\nFig 30\n(2 by 2 in.) bar. Similar specimens machined from 25 by 25 mm (1 by 1 in.) and 75 by 75 mm (3 by 3 in.)\nbars had four and eight tines, respectively.\n\n\nfew minutes. For these extremely rapid\noperations, maximum temperature may ex-\nceed 440 \u00b0C (825 \u00b0F).\nAlthough material annealed from the pre-\ncipitation-hardened condition usually has\nsufficient ductility for most forming opera-\n\n\nProducts that can be heated and cooled\nvery rapidly, such as wire, are annealed by\ncontinuous processes that require a total\nheating and cooling time of only a few\nseconds. Continuous annealing of coiled\nsheet is accomplished in a total time of a\n\n\nHeat Treating of Aluminum Alloys / 871\n\n\nTable 10 Effects of annealing treatments on ductility of 7075-0 sheet\n\n\nly controlled. Even allowing the load to cool\nin the furnace may result in an excessively\nhigh rate. Similarly, lowering the furnace-\ncontrol instrument by 28 \u00b0C (50 \u00b0F) each\nhour may produce stepped cooling, which is\nnot satisfactory for severe forming opera-\ntions. For maximum softening, a continu-\nous cooling rate of not more than 28 \u00b0C/h (50\n\u00b0F/h) is recommended.\nAnnealing of castings for 2 to 4 h at\ntemperatures from 315 to 345 \u00b0C (600 to 650\n\u00b0F) provides the most complete relief of\nresidual stresses and precipitation of the\nphases formed by the excess solute retained\nin solid solution in the as-cast condition.\nSuch annealing treatments provide maxi-\nmum dimensional stability for service at\nelevated temperatures. The annealed tem-\nper is designated \u201cO.\" (This temper was\ndesignated \"T2\" prior to 1975.)\n\n\nElongation in bending(c), %\nin 50 mm (2 in.) for\nthickness of:\n\n\nBend angle(b), degrees, for\nthickness of:\n\n\nfor thickness of:\nElongation in tension(a), % in 50 mm (2 in.)\n\n\nAnnealing\ntreatment\n\n\n0.5 mm\n(0.020 in.)\n\n\n1.6 mm\n(0.064 in.)\n\n\n1.6 mm\n(0.064 in.)\n\n\n2.6 mm\n(0.102 in.)\n\n\n2.6 mm\n(0.102 in.)\n\n\n1.6 mm\n(0.064 in.)\n\n\n2.6 mm\n(0.102 in.)\n\n\nTreatment 1(d)\nTreatment 2(e)\nTreatment 3(f)\n\n\n12\n14\n16\n\n\n12\n14\n16\n\n\n12\n14\n\n\n82\n91\n92.5\n\n\n73\n76\n84\n\n\n50\n57\n60\n\n\n48\n58\n56\n\n\n(a) Uniform elongation of gridded tension specimens. (b) Bend angle at first fracture. (c) Elongation in bend test for 1.3 mm (0.05 in.)\ngage spanning fracture. (d) Soak 2 h at 415 14 \u00b0C (775 \u00b1 25 \u00b0F); furnace cool to 260 \u00b0C (500 \u00b0F) at 30 \u00b0C/h (50 \u00b0F/h); air cool. (c) Soak\n2 h at 425 \u00b0C (800 \u00b0F), air cool; soak 2 h at 230 \u00b0C (450 \u00b0F), air cool. (f) Soak 1 h at 425 \u00b0C (800 \u00b0F); furnace cool to 230 \u00b0C (450 \u00b0F) at\n30 \u00b0C/h (50 \u00b0F/h); soak 6 h at 230 \u00b0C (450 \u00b0F), air cool\n\n\nemploy temperatures up to about 345 \u00b0C\n(650 \u00b0F), or up to 400 \u00b1 8 \u00b0C (750 \u00b1 15 \u00b0F)\nfor 3003 alloy, and cooling to room temper-\nature. No appreciable holding time is re-\nquired. Such treatment may result in simple\nrecovery, partial recrystallization, or full\nrecrystallization. Age hardening may follow\nstress-relief annealing of heat-treatable al-\nloys, however, because a concentration of\nsoluble alloying elements sufficient to cause\nnatural aging remains in solid solution after\nsuch treatments.\nA special form of stress-relief temper is\nused for heat-treatable alloy products that\nsubsequently will be inspected ultrasonical-\nly. The product is heated to its normal\nsolution heat-treating temperature, then\ncooled in still air to room temperature. This\ntemper is referred to as the O1 temper.\nControlled-Atmosphere Annealing and\nStabilizing. Aluminum alloys that contain\n\n\neven very small amounts of magnesium will\nform a surface magnesium oxide unless the\natmosphere in the annealing furnace is free\nof moisture and oxygen. Examples include\nalloy 3004, which is used for cooking uten-\nsils, and alloys of the 5xxx series.\nAnother problem that control of the an-\nnealing atmosphere helps to overcome or\navoid is oil staining by oil-base roll lubri-\ncants that do not burn off at lower annealing\ntemperatures. If the oxygen content of the\nfurnace atmosphere is kept very low during\nsuch annealing, the oil will not oxidize and\nstain the work.\nTemperature control for full and partial\nannealing is somewhat more critical than for\nstress-relief annealing; the temperatures\nand times specified are selected to produce\nrecrystallization and, in the case of heat-\ntreatable alloys, a precipitate of maximum\nsize; for this the cooling rate must be close-\n\n\nGrain Growth\n\n\nMany of the aluminum alloys in common\nuse are subject to grain growth during solu-\ntion treatment or annealing. This phenome-\nnon can occur during or after recrystallization\nof material that has been subjected to a small\ncritical amount of prior cold work. It is usu-\nally manifested by surface roughening during\nsubsequent fabrication operations and fre-\nquently results in rejections for appearance or\nfunctional reasons. Less frequently, some de-\nterioration of mechanical properties is en-\ncountered, and this is undesirable regardless\nof surface-roughening effects.\nDegree of susceptibility to grain growth\nvaries with alloy, structure, and chemical-\ncomposition variation, and from one prod-\nuct form to another. The critical range of\ncold work is ordinarily about 5 to 15%.\nUsually, temperatures of 400 \u00b0C (750 \u00b0F)\nand above must be reached before grain\ngrowth occurs, but some growth has been\nencountered at temperatures as low as 345\n\u00b0C (650 \u00b0F). Grain growth that occurs during\ninitial recrystallization is more a function of\ncomposition, structure, and degree of cold\nwork than of temperature per se; tempera-\ntures in excess of 455 \u00b0C (850 \u00b0F) in common\nalloys can lead to secondary-recrystalliza-\ntion grain-growth problems. The common\nsymptom indicating moderately large-grain\nmaterial is roughening or \"orange peel\" on\nthe external surfaces of bends. Severe\ngrowth of grains to fingernail size and larger\nsometimes is evident in parts made from\nannealed (O temper) material by stretch\nforming and then thermal treating or similar\noperations. This type of grain growth often\nis detected during subsequent anodizing,\netching, and chemical milling operations.\nCracking during welding or brazing is\nanother characteristic which may indicate\nthat severe grain growth has occurred. In\nsuch instances, cracks propagate along\ngrain boundaries that provide little obstruc-\ntion to their progress.\n\n\nTypical full annealing treatments for some common wrought aluminum alloys\nThese treatments, which anneal the material to the O temper, are typical for various sizes and methods of\nmanufacture and may not exactly describe optimum treatments for specific items.\n\n\nTable 11\n\n\nApproximate\ntime at\ntemperature, h\n\n\nApproximate\ntime at\ntemperature, h\n\n\nMetal temperature\n\n\nMetal temperature\n\n\n\u00b0C\n\n\nAlloy\n\n\n\u00b0F\n\n\nAlloy\n\n\n\u00b0C\n\n\n\u00b0F\n\n\n1060\n1100\n1350\n2014\n2017\n2024\n2036\n2117\n2124\n2219\n3003\n3004\n3105\n5005\n5050\n5052\n5056\n5083\n5086\n5154\n5182\n5254\n5454\n5456\n\n\n345\n345\n345\n\n\n650\n650\n650\n\n\n5457\n5652\n6005\n6009\n6010\n6053\n6061\n6063\n6066\n7001\n7005\n7049\n7050\n7075\n7079\n7178\n7475\n\n\n(a)\n(a)\n(a)\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n(a)\n(a)\n(a)\n(a)\n(a)\n(a)\n(a)\n(a)\n\n\n345\n345\n\n\n650\n650\n\n\n(a)\n(a)\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n2-3\n\n\n415(b)\n\n\n775(b)\n\n\n415(b)\n415(b)\n415(b)\n385(b)\n415(b)\n415(b)\n415(b)\n\n\n775(b)\n775(b)\n775(b)\n725(b)\n775(b)\n775(b)\n775(b)\n\n\n415(b) 775(b)\n415(b) 775(b)\n415(b) 775(b)\n415(b) 775(b)\n415(b) 775(b)\n415(b) 775(b)\n\n\n775(c)\n650(d)\n775(c)\n775(c)\n775(c)\n775(c)\n\n\n415(c)\n345(d)\n415(c)\n415(c)\n415(c)\n415(c)\n\n\n775\n650\n650\n650\n650\n650\n650\n650\n650\n650\n650\n650\n650\n650\n\n\n415\n345\n345\n345\n345\n345\n345\n345\n345\n345\n345\n345\n345\n345\n\n\n415(c) 775(c)\n\n\n775(c)\n\n\n415(c)\n\n\nBrazing sheet\n\n\n(a)\n(a)\n(a)\n(a)\n\n\nNo. 11 and 12\nNo. 21 and 22\nNo. 23 and 24\n\n\n345\n345\n345\n\n\n650\n650\n650\n\n\n(a)\n(a)\n(a)\n\n\n(a) Time in the furnace need not be longer than necessary to bring all parts of the load to appealing temperature. Cooling rate is\nunimportant. (b) These treatments are intended to remove the effects of solution treatment and include cooling at a rate of about 30 \u00b0C/h\n(50 \u00b0F/h) from the annealing temperature to 260 \u00b0C (500 \u00b0F). Rate of subsequent cooling is unimportant. Treatment at 345 \u00b0C (650 \u00b0F),\nfollowed by uncontrolled cooling, may be used to remove the effects of cold work or to partly remove the effects of heat treatment. (c)\nThese treatments are intended to remove the effects of solution treatment and include cooling at an uncontrolled rate to 205 \u00b0C (400 \u00b0F)\nor less, followed by reheating to 230 \u00b0C (450 \u00b0F) for 4 h. Treatment at 345 \u00b0C (650 \u00b0F), followed by uncontrolled cooling, may be used\nto remove the effects of cold work or to partly remove the effects of heat treatment. (d) Cooling rate to 205 \u00b0C (400 \u00b0F) or below is less\nthan or equal to 30 \u00b0C/h (50 \u00b0F/h).\n\n\n872 / Heat Treating of Nonferrous Alloys\n\n\nreducing the size of furnace loads or by\nchanging from an air furnace to a salt bath.\nIn one application, severe grain growth was\nfound during bending of alloy 1100 rectan-\ngular tubing. The roughening of the inside\nsurfaces of the parts, which occurred during\nforming of the large-grain material, im-\ntheir functioning as radar\nwaveguides. Investigation disclosed that, to\nminimize handling marks, the material was\nprocured in the strain-hardened (H14) tem-\nper and was stress-relief annealed at 345 \u00b0C\n(650 \u00b0F) immediately prior to forming. Grain\ngrowth occurred during annealing as a re-\nsult of the moderate amount of cold work\nintroduced at the mill. The problem was\neliminated by changing the stress-relieving\noperation to a 5-min heating period in an air\nfurnace operating at 540 \u00b0C (1000 \u00b0F). The\nexplanation advanced for the success of this\ntreatment was that, due to the rapid heating\nrate, the temperature of the material was\nraised through the recrystallization range\nfor the less severely cold-worked grains\nbefore the critically cold-worked grains had\ntime to grow appreciably.\n\n\n200\n\n\n\u201c100 B\n\n\n175 \u00b0C (350 \u00b0F)-\n\n\n150\n\n\nYield strength, ksi\n\n\nYield strength, MPa\n\n\n20\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n100\n\n\npaired\n\n\n230 \u00b0C (450 \u00b0F)\n\n\n10\n\n\n50\n\n\n260 \u00b0C (500 \u00b0F)\n\n\n288 and 316 \u00b0C (550 and 600 \u00b0F)\n\n\n0\n\n\n1\n2\n\n\n1\n3\n\n\n0\n\n\n0.5\n\n\n1\n\n\n1.5\n\n\n2.5\nTime, h\n\n\n3.5\n\n\n4\n\n\n4.5\n\n\n5\n\n\n350\n\n\n\u25ac\u25ac\u25ac\u25ac\u25ac 5052-+18 7 50\n\n\n300\n\n\n40\n40\n\n\n175 \u00b0C (350 \u00b0F)\n\n\n205 \u00b0C (400 \u00b0F)\n\n\n250\n\n\n230 \u00b0C (450 \u00b0F)\n\n\nYield strength, MPa\n\n\nYield strength, ksi\n\n\n30\n\n\n200\n\n\n260 \u00b0C (500 \u00b0F)\n\n\n150\n\n\n20\n\n\nHeating Equipment and\nAccessories\n\n\n290 \u00b0C (550 \u00b0F)\n\n\n100\n\n\n\u2610 315 \u00b0C (600 \u00b0F)\n\n\nThe general methods for heat treating\naluminum alloys include the use of molten\nsalt baths, air-chamber furnaces, and induc-\ntion heaters. The choice of heating equip-\nment depends largely on the alloy and the\nconfiguration of the parts to be processed.\nThe type of heat treatment can also influ-\nence the choice of heating equipment. For\nexample, both molten salt baths and air-\nchamber furnaces are suitable for solution\ntreating of aluminum alloys, while induction\nheating requires additional analysis to de-\nfine the proper temperature range for solu-\ntion treatment. Induction methods can pro-\nvide high heating rates, which affect\ntransformation behavior (see, for example,\nthe section \"Nonequilibrium Melting\" in\nthis article).\nMolten salt baths and air-chamber furnac-\nes both have advantages and disadvantages\nin solution heat treatments, as discussed\nbelow. Oil and gas-fired furnaces, in de-\nsigns that allow the products of combustion\nto come in contact with the work, are\nusually unsatisfactory because they pro-\nmote high-temperature oxidation.\nSalt baths heat the work faster (see Table\n2) than air furnaces, provided that the\namount of work introduced at any one time\nis controlled to prevent the temperature\nfrom falling below the desired range. If the\ntemperature is permitted to fall below the\nminimum limit, much of the advantage of\nthe salt bath is lost, because of the necessity\nfor reheating the large mass of salt.\nSalt baths are also more readily adapted\nto the introduction, at any time, of small\namounts of work requiring different soaking\n\n\n10\n\n\n50\n\n\n0\n0\n\n\n0.5\n\n\n1\n\n\n1.5\n\n\n2\n\n\n2.5\n\n\n3\n\n\n3.5\n\n\n4.5\n\n\n5\n\n\n4\n\n\nTime, h\n\n\nFig 32 Representative isothermal annealing curves for alloys 1100-H18 and 5052-H18\n\n\nIn other similar investigations, no detri-\nmental effects have been discovered, and in\nmany cases such parts have served satisfac-\ntorily in critical applications.\nWhen a grain-growth problem is discov-\nered, it is too late to change the condition of\nthe parts in question, but several possible\nmethods are available for preventing recur-\nrence of the difficulty. The simplest of these\nis relieving the causative stress by interject\ning a stress-relief anneal into the manufac-\nturing sequence immediately prior to the\nsolution-treating or full-annealing cycle in\nwhich the grain growth occurred. This ap-\nproach is usually successful and practical.\nAnother possibility is to adjust the amount\nof stress present in the part immediately\nprior to the critical heat treatment so that\nthe stress level is outside the critical range.\nThis may be done by adding a cold-working\noperation before forming, such as pre-\nstretching of blanks, or by forming in mul-\ntiple stages with a stress-relief anneal before\neach stage.\nA third method that is sometimes suc-\ncessful consists of increasing the heating\nrate during the critical heat treatment by\n\n\nIf the surface roughening is objectionable\nfrom either an appearance or a functional\naspect, the desirability of surface-smooth-\ning operations, such as sanding or buffing,\nmust be evaluated. If reductions in mechan-\nical properties are suspected, these must be\nestablished by test and evaluated in relation\nto the anticipated service.\nIn one application, a part that had been\nmade by stretch forming O-temper 2 mm\n(0.080 in.) sheet and heat treating exhibited\nsignificantly lower tensile and yield\nstrengths in portions where severe grain\ngrowth had occurred than in portions hav-\ning normal grain size:\n\n\nYield strength\n\n\nTensile strength\n\n\nGrain\nstructure\n\n\nTest\n\n\nksi\n\n\nMPa\n\n\nksi\n\n\nMPa\n\n\nTransverse\n\n\nCoarse\nCoarse\n\n\n35.8\n35.0\n37.8\n\n\n1\n2\n3\n\n\n265\n263\n311\n\n\n38.5\n38.2\n45.1\n\n\n247\n241\n261\n\n\nFine\n\n\nLongitudinal\n\n\n259\n269\n305\n\n\n243\n245\n270\n\n\nCoarse\nCoarse\n\n\n37.6\n39.0\n44.2\n\n\n35.3\n35.6\n39.1\n\n\n1\n2\n3\n\n\nFine\n\n\nHeat Treating of Aluminum Alloys / 873\n\n\nProduction rate, tons/h\n\n\nTable 12 Frequency selection for induction heating with a longitudinal-flux coil and a\ntransverse-flux coil\n\n\n1\n\n\n7\n\n\n0\n\n\n2\n\n\n3\n\n\n5\n\n\n6\n\n\n4\n\n\n1800\n\n\nMinimum part thickness, mm (in.), for a frequency of:\n200 Hz\n\n\nSolution heat treat at 590 \u00b0C\n\n\n1 kHz\n\n\nMaterial\n\n\n1500\n\n\n60 Hz\n\n\n3 kHz\n\n\n10 kHz\n\n\nFull anneal at 425 \u00b0C\nPartial anneal at 315 \u00b0C\n\n\nPower source, kW\n\n\nSolenoid (longitudinal-flux) coil\nSteel below Curie temperature\nSteel above Curie temperature\nBrass\n\n\n1200\n\n\n2.3 (0.09)\n25 (1.0)\n7 (0.28)\n5 (0.2)\n\n\n>38 (1.5)\n>175 (7.0)\n>50 (2.0)\n>38 (1.5)\n\n\n1 (0.04)\n13 (0.5)\n4 (0.16)\n3 (0.12)\n\n\n13 (0.5)\n100 (4.0)\n28 (1.1)\n22 (0.85)\n\n\n5 (0.2)\n43 (1.7)\n13 (0.5)\n\n\n900\n\n\nAluminum\n\n\n9 (0.375)\n\n\n600\n\n\nTransverse-flux coil\n\n\n0.25 (0.01)\n0.5 (0.02)\n2.5 (0.1)\n\n\nAluminum\n\n\n1.3 (0.05)\n\n\n0.04 (0.0016)\n0.08 (0.0032)\n0.5 (0.020)\n\n\n0.08 (0.003)\n0.15 (0.006)\n0.9 (0.035)\n\n\n300\n\n\n>5 (0.2)\n>10 (0.4)\n>50 (2.0)\n\n\n2.5 (0.1)\n13 (0.5)\n\n\nBrass\nSteel\n\n\n0\n\n\n15 20\nLine speed, m/min\n\n\n30\n\n\n5\n\n\n10\n\n\n25\n\n\n0\n\n\nperiods. (Economical utilization of air fur-\nnaces usually dictates accumulation of a\nlarge load of parts of similar thickness be-\nfore charging.) Also, the buoyant effect of\nthe salt reduces distortion during heating,\nand the large reservoir of heat facilitates\ntemperature control and uniformity.\nSalt bath operation entails special house-\nkeeping requirements. Dragout is costly and\nunsightly. Because residual salt on parts\nmay result in corrosion, all salt must be\ncompletely removed, including that from\ncrevices and blind holes. In addition, salt\nresidue from the quench water must be kept\nto a minimum by a constant water overflow\nor by providing a fresh-water rinse for all\nparts after quenching. When these provi-\nsions are impractical, corrosion can be in-\nhibited by adding 14 g (\u00bd oz) of sodium or\npotassium dichromate to each 45 kg (100 lb)\nof the molten salt.\nPrecautions. Molten salt baths are poten-\ntially hazardous and require special precau-\ntions. Operators must be protected from\nsplashing and dripping of the hot salt. Be-\ncause heated nitrates are powerful oxidizing\nagents, they must never by allowed to come\nin contact with combustibles and reducing\nagents, such as magnesium and cyanides.\nMost authorities advise against inserting\naluminum alloys containing more than a few\npercent of magnesium into molten nitrate.\nTo avoid exposure of personnel to nitrous\nfumes produced during decomposition of\nnitrates, good ventilation is essential.\nWhen molten nitrates are being used, the\npossibilities of explosions resulting from\nboth physical and chemical reactions must\nbe avoided. The former result from rapid\nexpansion of gases entrapped beneath the\nsurface of the bath. Hence, parts entering\nthe bath must be clean and dry; they must\nalso be free of pockets or cavities that\ncontain air or other gases. Chemical-reac-\ntion explosions result from rapid break-\ndown of the nitrates due to overheating or\nreaction with the pot material. Stainless\nsteel pots (preferably of type 321 or 347) are\nmore resistant to scaling than those made of\ncarbon steel or cast iron and therefore pre-\nsent a lower probability of local overheat-\ning. Sludge or sediment accumulations in\n\n\nPower requirement for transverse-flux induc-\nFig 33\ntion heating of aluminum strip 1 mm (0.04 in.)\nthick and 1270 mm (50 in.) wide. Source: Ref 9\n\n\nbottom-heated pots can also lead to local\noverheating. Overheat controls are essen-\ntial to ensure against temperatures exceed-\ning 595 \u00b0C (1100 \u00b0F).\nIt is vitally important that water be kept\naway from a nitrate tank. In controlling a\nnitrate fire, do not use water or any fire\nextinguisher containing water. The best ex-\ntinguisher is dry sand, a supply of which\nshould be kept near the tank.\nExtra sacks of salt should be stored in a\ndry place, distant from the tank. If the fresh\nsalt being added to the bath is even slightly\ndamp, it should be added very slowly or\nwhen the bath is frozen.\nAir furnaces are used more widely than salt\nbaths because they permit greater flexibility\nin operating temperature. When production\nschedules and the variety of alloys requiring\nheat treatment necessitate frequent changes\nin temperature, the time and cost of adjusting\nthe temperature of a large mass of salt makes\nthe use of an air furnace almost mandatory.\nHowever, waiting periods are often required\nto allow the walls of air furnaces to stabilize at\nthe new temperature before parts are intro-\nduced. Otherwise, parts may radiate heat to\ncolder walls or absorb radiant heat from hot-\nter walls, and the temperature indicated by\nthe control instrument will not reflect actual\nmetal temperature in the usual manner. Air\nfurnaces are also more economical when the\nproduct mix includes a few rather large parts;\nholding the temperature of a large volume of\nsalt in readiness for an occasional large part is\nfar more expensive than heating an equal\nvolume of air.\nInduction heating with either solenoid\n(longitudinal-flux) coils or transverse-flux\ncoils provides an efficient method for in-line\nheating of flat-rolled products (particularly\nstrip). Solenoid coils create a longitudinal\nflux, which can produce efficient heating for\nthicker and/or lower resistivity materials.\nSolenoid coils can also be used efficiently in\nthe heating of thinner magnetic material\n(see, for example, steel below the Curie\ntemperature in Table 12).\nFor several nonferrous materials (alumi-\nnum, copper, brass), however, efficiency\nand power factors with solenoid coils are\nsignificantly lower than for ferrous materi-\n\n\nals. Therefore, transverse-flux coils are ide-\nally suited for heating nonferrous materials,\nbecause transverse-flux lines do not exhibit\nthe degree of current cancellation associat-\ned with longitudinal flux lines. This aspect\nof transverse-flux coils improves efficiency\nand also permits the use of lower frequen-\ncies (Table 12). This reduces the capital\nequipment costs, and where it shifts from\nrequiring RF frequencies, the power source\nconversion efficiency is also significantly\nimproved. Aluminum, brass, copper, and\naustenitic stainless steel strip lines are ide-\nally suited for transverse-flux heating. Each\nof these materials often requires in-line pro-\ncesses like partial or full annealing or solu-\ntion heat treating to provide necessary me-\nchanical properties for subsequent finishing\noperations.\nTransverse-flux induction heating offers\nseveral benefits for in-line strip heating and\nhas been known for many years. However,\nit requires specially designed iron-cored\nlaminated inductor coils and tighter control\nof the power, strip handling, and process\nparameters. Frequency selection is dictated\nby the resistivity and thickness of the ma-\nterial, while power requirements depend on\nthe production rates, the specific heat, and\nthe processing temperatures for a given\nmaterial. Figure 33 shows typical power\nsource requirements for transverse-flux\nheating of aluminum strip.\n\n\nFurnace Temperature Control\n\n\nThe importance of close temperature con-\ntrol in solution treating has been noted in\nthe previous section on solution treating.\nEach control zone of each furnace should\ncontain at least two thermocouples. One\nthermocouple, with its instrument, should\nact as a controller, regulating the heat input;\nthe other should act independently as a\nsafety cutoff, requiring manual reset if its\nset temperature (usually the maximum of\n\n\n874 / Heat Treating of Nonferrous Alloys\n\n\nfactors should be applied after each probe\ncheck, but if the correction required ex-\nceeds \u00b13 \u00b0C (\u00b15 \u00b0F), the source of the\ndeviation should be corrected. MIL-H-6088\nrecommends that this check be made week-\nly, but many operators make the check as\nfrequently as once each shift.\nTemperature-Uniformity Surveys. In con-\ntrolling the temperature of parts that are\nbeing heat treated it must first be deter-\nmined that the temperature indicated by the\nfurnace instruments truly represents the\ntemperature of the nearby air or salt. Sec-\nond, the uniformity of temperature within\nthe working zone must be shown to be\nwithin a range of 11 \u00b0C, or 20 \u00b0F (6 \u00b0C, or 10\n\u00b0F, for precipitation heat treatment of alloy\n2024). This is accomplished by measuring\nthe temperature at several test locations,\nusing calibrated test thermocouples and a\ncalibrated test potentiometer, and reading\nfurnace instruments nearly simultaneously.\nMIL-H-6088 recommends monthly surveys\nwith one test location per 1.1 m\u00b3, or 40 ft\u00b3\n(0.7 m\u00b3, or 25 ft\u00b3, for air furnaces on initial\nsurvey), but with a minimum of nine test\nlocations distributed as shown in Fig 34.\nDespite the large size of some furnaces,\nrather surprising temperature uniformities\nhave been reported. In one instance the\ninitial survey of an air furnace measuring\n12.5 by 1.2 by 3.0 m (41 by 4 by 10 ft)\nshowed maximum temperature variations of\n+1.7, \u22121.1 \u00b0C (+3, \u22122 \u00b0F). When a parti-\ntion 0.3 m (1 ft) thick was lowered, convert-\ning the furnace to two chambers 6.1 by 1.2\nby 3.0 m (20 by 4 by 10 ft) each, the spread\nwas 1.1, 0.6 \u00b0C (+2, -1 \u00b0F) in one\nsection and +0.6, \u22121.1 \u00b0C (+1, \u22122 \u00b0F) in\nthe other.\nFor each furnace load, one thermocouple\n(the \"cold\" couple) should be placed in the\ncoldest area of the furnace and another (the\n\"hot\" couple) in the hottest area. In addi-\ntion to these two thermocouples, a load\nthermocouple should be installed. The load\ncouple should be of approximately the same\ngage as the sheet or other product being\nheat treated. If heavy plate, forgings, or\ncastings are being heat treated, a similar\ndiscarded item should be used at the con-\ntrolling load couple. The thermocouple\nshould be placed in a drilled hole and\npacked to hold it firmly in place during the\nheat-treating cycle. In some instances, the\nitems being heat treated can be used as the\nload couples. The thermocouples can be\nplaced in holes drilled in areas that will be\nremoved in making the finished article.\nIt is important that items of different\nthicknesses-1 mm (0.040 in.) sheet and 25\nmm (1 in.) plate, for example-not be heat\ntreated in the same furnace load.\nIn salt baths, uniformity surveys usually\nare made by holding a probe thermocouple\nin each location until thermal equilibrium is\nreached; in air furnaces, a mock heat-treat-\ning cycle is required. First, the air furnace is\n\n\nthe specified range) is exceeded during the\nsolution-treating cycle.\nSafety cutoffs are mandatory for salt\nbaths to guard against explosions and often\nhave paid for themselves in air furnaces by\nsaving a load of parts or even the furnace\nitself. It is important, however, that they be\ntested periodically (by deliberately over-\nshooting the empty furnace) to guard\nagainst \"frozen\" corroded contacts result-\ning from prolonged periods of idleness.\nAt least one of the instruments for each\nzone should be of the recording type, and\nboth instruments should have restricted\nscales for instance, 400 to 600 \u00b0C (750 to\n1110 \u00b0F), rather than 0 to 600 \u00b0C (32 to 1110\n\u00b0F). This is required for maximum accuracy\nbecause manufacturers' guarantees are\nspecified in terms of percent of scale.\nIn the placement of instruments, expo-\nsure to extremes in ambient temperature,\nhumidity, vibration, dust, and corrosive\nfumes should be avoided. Ambient temper-\natures between 5 and 50 \u00b0C (40 and 120 \u00b0F)\nare satisfactory, but temperature changes of\n6 \u00b0C/h (10 \u00b0F/h) or more should be avoided.\nIt is also essential that instruments and\nthermocouple circuits be shielded from\nelectromagnetic fields commonly associat-\ned with the leads of high-amperage furnace\nheating elements.\nTemperature-sensing elements must be\ncapable of responding more rapidly to tem-\nperature changes than the materials being\nprocessed. Therefore, thermocouple wire\ndiameter should not exceed 1\u00bd times the\nthickness of the minimum-gage material to\nbe heat treated, and should in no case\nexceed 14 gage. Thermocouples for salt\nbaths should be enclosed in suitable protec-\ntion tubes. Air-furnace thermocouples\nshould be installed in open-end protection\ntubes, with the thermocouple junction ex-\ntending sufficiently beyond the tube to pre-\nvent any loss in sensitivity.\nTemperature-sensing elements should be\nlocated in the furnace work chamber, not in\nducts and plenums, and should be as close\nas possible to the working zone. Specifica-\ntion MIL-H-6088C restricts distance be-\ntween the sensing element and the working\nzone to a maximum of 100 mm (4 in.). The\nsafety-cutoff thermocouple should be locat-\ned to reflect the highest temperature in the\nworking zone. The control thermocouple\nshould be located in a position where it will\nread a temperature approximately halfway\nbetween the hottest and coldest tempera-\ntures.\nProbe Checks. After the temperature-\nmeasurement equipment is properly in-\nstalled, it must be checked frequently for\naccuracy. This is accomplished by inserting\na calibrated probe thermocouple into the\nfurnace adjacent to each furnace thermo-\ncouple and comparing its reading on a cali-\nbrated test potentiometer with that indicat-\ned by the furnace instrument. Correction\n\n\n6\n\n\n5\n\n\n5-194\n\n\n=\n\n\n8\n\n\nRectangular furnace\n\n\nWINE\n6 5 4\n\n\n\u798f\n\n\n8\n\n\nCylindrical salt bath\n\n\nCylindrical air furnace\nLocation of thermocouples for surveying\nFig 34\ntemperature uniformity in the working\nzones of air furnaces and salt baths\n\n\nstabilized at the test temperature. Then a\nrack containing the test thermocouples is\ninserted into the furnace. By using multiple\nswitches or a multipoint recording instru-\nment, all test thermocouples and furnace\ninstruments are read every 5 min. As the\ntemperature approaches the test range, it is\nadvisable to increase the frequency of read-\nings to detect possible overshooting. After\nthermal equilibrium is reached, readings\nshould be continued until the recurrent tem-\nperature pattern is established.\nSurveys of salt baths generally are con-\nsidered acceptable whether they are made\nwhile the bath is empty or filled with work.\nIt is controversial whether surveys of air\nfurnaces should be made with or without a\nload. Undoubtedly, recovery overshoots\nare most likely to occur with a very light\nload and would not be detected if a heavier\nload were used. Certainly, if all loads are\nessentially alike, surveys should be made\nwith typical loads. With widely varying\nloads, the optimum approach is to make\nseveral surveys initially, including one with\nan empty furnace, and then to make suc-\nceeding surveys with an empty furnace to\nensure against changes in furnace charac-\nteristics. If any changes are made in the\nfurnace that might affect temperature distri-\nbution, such as repair of vanes or louvers,\nseveral surveys should be repeated.\nAnother aspect of the problem of temper-\nature control in air furnaces is the necessity\nof ensuring that the temperature of the parts\nis the same as that of the surrounding air.\nFurnace components whose temperature\ndiffers from the air temperature must be\nsuitably shielded to prevent radiation to or\n\n\nHeat Treating of Aluminum Alloys / 875\n\n\nas possible to the actual temperature. To\nachieve this, it is necessary to apply correc-\ntion factors obtained during calibration to\nthe next lower echelon of accuracy. Even\nthen, if all errors inherent in the chain are in\nthe same direction, a considerable differ-\nence will exist between the measured and\nactual temperatures. Therefore, it is advis-\nable to operate as close to the mean of the\ndesired range as possible.\n\n\nfrom the parts being heat treated. In a\nfurnace used for solution heat treating of\nrivets, unshielded heating elements have\nbeen known to produce part temperatures\nas much as 20 \u00b0C (35 \u00b0F) higher than the\ncontrol temperature, resulting in eutectic\nmelting and cracking. In two other instanc-\nes, reradiation through inadequate shielding\nproduced a radiation effect of as much as 11\n\u00b0C (20 \u00b0F). One of these problems was\nsolved by painting the shield with reflective\naluminum paint and the other by adding a 13\nmm (1/2 in.) thick layer of asbestos to the 1.6\nmm (1/16 in.) stainless steel shield.\nFurnace-wall temperatures that differ ap-\npreciably from the temperature of the parts\nalso must be avoided. Consequently, when\nthe operating temperature of an air furnace\nis changed, waiting periods are required\nafter the furnace instrument indicates sta-\nbility, to allow the furnace walls to stabilize\nat the new temperature. The magnitude of\nthis limitation is directly proportional to the\nefficiency of the furnace as an insulated\nchamber, but possibilities of such radiation\nshould be recognized even in thin-wall fur-\n\n\nAverage cooling rate at center\nof cylinders, \u00b0F/s\n\n\n10\n\n\n100\n\n\n500\n\n\nLongitudinal stress range, MPa\n\n\nLongitudinal stress range, ksi\n\n\n1660\n\n\n400\n\n\n300\n\n\n40\n\n\n200\n\n\n20\n20\n\n\n\u00b0\n\n\n100\n\n\nPoo\n\n\n0\n\n\n0\n\n\nDimensional Changes during\nHeat Treatment\n\n\n10\n\n\n100\n\n\nAverage cooling rate at center\nof cylinders, \u00b0C/s\n\n\nIn addition to the completely reversible\nchanges in dimensions that are simple func-\ntions of temperature change and are caused\nby thermal expansion and contraction, di-\nmensional changes of more permanent char-\nacter are encountered during heat treat-\nment. These changes are of several types,\nsome of mechanical origin and others\ncaused by changes in metallurgical struc-\nture. Changes of mechanical origin include\nthose arising from stresses developed by\ngravitational or other applied forces, from\nthermally induced stresses or from relax-\nation of residual stresses. Dimensional\nchanges also accompany recrystallization,\nsolution, and precipitation of alloying ele-\nments.\nSolution Heat Treatment. Distortion as a\nresult of creep during solution heat treat-\nment should be avoided by proper loading\nof parts in baskets, racks, or fixtures, or by\nprovision of adequate support for long piec-\nes of plate, rod, bar, and extrusions heat\ntreated in horizontal roller hearth furnaces.\nSheet is provided with air-pressure support\nin continuous heat-treating furnaces to\navoid scratching, gouging, and distortion. If\nparts are to be solution heat treated in\nfixtures or racks made of materials (such as\nsteel) with coefficients of thermal expansion\nlower than that of the aluminum being treat-\ned, allowance should be made for this dif-\nferential expansion to ensure that expansion\nof the aluminum is not restricted. Straight-\nening immediately after solution heat treat-\ning may be preferable to fixturing.\nSolution of phases formed by major alloy-\ning elements causes volumetric expansion\nor contraction, depending on the alloy sys-\ntem, and this may have to be taken into\naccount in heat treatment of long pieces.\nFor example, solution heat treatment and\nquenching of alloy 2219 causes lengthwise\ncontraction of about 2 mm/m (0.002 in./in.).\nSolution heat treatment and quenching of\nalloys of the 7xxx series is accompanied by\nlengthwise expansion-about 0.6 mm/m\n(0.0006 in./in.) for alloy 7075 rod or plate.\nQuenching. The most troublesome\nchanges in dimensions and shape are those\nthat occur during quenching or that result\nfrom stresses induced by quenching. Due to\nits nonuniform cooling, quenching may pro-\nduce warpage or distortion, particularly in\n\n\nEffect of quenching rate on longitudinal\nFig 35\nstress ranges in alloy 2014-T4 cylinders\nquenched in various media. Cylinders were 75 mm (3\nin.) in diameter by 230 mm (9 in.) long. Cooling rate\nwas measured from 400 to 290 \u00b0C (750 to 555 \u00b0F). Stress\nrange is maximum tensile stress plus maximum com-\npressive stress.\n\n\nthin material and in thin sections of parts\nthat contain variations in thickness. For\nthick-section products or parts, changes in\nexternal shape may be small because of\nrigidity, but the interior-to-surface temper-\nature gradients that form with rapid cooling\ncreate residual stresses; these stresses nor-\nmally are compressive at the surfaces and\ntensile in the interior.\nAs previously discussed, warpage or dis-\ntortion of thin-section material can be re-\nduced by using a quenching medium that\nprovides slower cooling; however, cooling\nmust be sufficient to produce the required\nproperties. Slower quenching can also re-\nduce the magnitude of residual stresses in\nthicker parts or pieces, as shown in Fig 9 for\ncylindrical specimens of alloy 6151\nquenched in cold or boiling water. Stress\nrange (maximum tensile stress plus maxi-\nmum compressive stress) for a cylinder with\na radius of 89 mm (3.5 in.) is about 205 MPa\n(30 ksi) when the cylinder is quenched in\ncold water but less than 70 MPa (10 ksi)\nwhen it is quenched in boiling water. The\neffects of average cooling rate through the\ntemperature range from 400 to 290 \u00b0C (750\nto 550 \u00b0F) on longitudinal stress ranges\ndeveloped in alloy 2014 cylinders 75 mm (3\nin.) in diameter are shown in Fig 35.\nHigh stresses induced by rapid quenching\ngenerally are reduced only modestly by the\nprecipitation heat treatments used to pro-\nduce T6- or T8-type tempers. Consequent-\nly, for the alloys that require rapid cooling\nto develop the properties of these tempers,\nthose incorporating mechanical stress relief\n(Tx51, Tx52) usually are specified when\nsubstantial metal must be removed to pro-\nduce final shapes. Other T8-type tempers,\nsuch as T86 and T87, also have low residual\nstress as a result of the stretching required\nto produce them.\nHeat Treatments for Precipitation and Sta-\nbilization. The most significant dimensional\nchanges associated with precipitation heat\n\n\nnaces.\n\n\nRadiation effects are potentially danger-\nous because they often cannot be detected\nby ordinary thermocouples. Specially pre-\npared radiation panels with thermocouples\nattached are used, and their readings are\ncompared with adjacent free thermocou-\nples. These panels normally are made of\nmaterial of the same gage as the thinnest\nparts to be heat treated and should have a\nsingle surface area of about 650 cm\u00b2 (100\nin.2). A thermocouple is attached to the\ncenter of the panel by welding or peening.\nIn order to detect the maximum effect,\npanel surfaces should be darkened so that\ntheir emissivity is at least as high as that of\nany material to be processed. During the\ntest, the panel surfaces should be parallel to\nthe suspected source or recipient of radia-\ntion. As an example of the number of panels\nrequired, several aerospace companies\nspecify one panel for every 1.5 linear meters\n(5 linear feet) of furnace wall.\nInstrument Calibration. All instruments\nand thermocouples must be accurately cal-\nibrated, and it is essential that the calibra-\ntions be traceable directly to the National\nBureau of Standards. The chain of trace-\nability should consist of not more than four\nlinks for sensing elements and three links\nfor measuring elements. To illustrate, if the\narticle calibrated by the National Bureau of\nStandards is called a primary standard, then\nthe chain of traceability of measuring ele-\nments should consist of primary standard,\ntest potentiometer, and furnace instrument.\nSimilarly, the chain for sensing elements\nshould consist of primary standard, second-\nary standard, test thermocouple, and fur-\nnace thermocouple. Every effort should be\nmade to ensure that the temperature indi-\ncated by the furnace instruments is as close\n\n\n876 / Heat Treating of Nonferrous Alloys\n\n\ntreatments and stabilizing heat treatments\narise from concurrent dilution of the solid\nsolution (which changes lattice parameter)\nand formation of precipitate. Changes in\ndensity and specific volume resulting from\nthese changes in metallurgical structure are\nthe reverse of those caused by solution of\nthe alloy phases. However, because the\nstrongest tempers are those in which the\nprecipitate is present in nonequilibrium\ntransition forms, the amount of change dur-\ning precipitation heat treatment does not\ntotally compensate for the previous (and\nopposite) change that occurred during solu-\ntion heat treatment. Most of the heat-treat-\nable alloys expand (grow) during precipita-\ntion heat treatment. Exceptions are alloys\nof the 7xxx wrought series and the 7xx.0\ncasting series, which exhibit contraction.\nIn alloys of the 2xxx series, the amount of\ngrowth decreases with increasing magne-\nsium content. Thus, growth of about 1.5\nmm/m (0.0015 in./in.) can be expected dur-\ning precipitation heat treatment of alloy\n2219-T87, about 0.5 mm/m (0.005 in./in.) for\ntreatment of alloy 2014-T6 and less than 0.1\nmm/m (0.0001 in./in.) for treatment of alloy\n2024-T851. Alloys 7050 and 7075, on the\nother hand, contract about 0.3 mm/m\n(0.0003 in./in.) on precipitation heat treating\nfrom the W temper to the T6 temper and\nabout 0.7 mm/m (0.0007 in./in.) on treating\nfrom the W temper to the T73 temper.\nStabilizing T7-type treatments cause great-\ner amounts of growth than the T5-, T6-, or\nT8-type treatments for the same alloys. This\nincreased growth is associated either with\nformation of increased amounts of transition\nprecipitates or with transformation of transi-\ntion precipitates to equilibrium phases.\n\n\nThe T3- and T4-type tempers are the least\nstable dimensionally because of possible\nprecipitation in service. Alloys 2024 and its\nvariants have the smallest dimensional\nchange in aging; the total change from the\nquenched to the average state is of the order\nof 0.06 mm/m (0.00006 in./in.), less than the\nchange due to a temperature variation of 3\n\u00b0C (5 \u00b0F). These alloys therefore can be used\nin the T3- and T4-type tempers, except for\nprecision equipment. For all other alloys,\nT6- or T8-type tempers should be used,\nbecause in these tempers all the alloys have\ngood dimensional stability.\n\n\nYield strength, ksi\n\n\n60\nT\n\n\n65\n1\n\n\n70\nT\n\n\n75\n\n\n50\n\n\n7075-T6\nAlclad sheet\n\n\n40\n\n\n180 specimens\nfrom a single sheet\n\n\n30\n\n\n% of tests\n\n\n4290 routine\nmill tests\n\n\nde 20\n\n\n10\n\n\nStability of Precision Equipment. Proper\nmaintenance of high-precision devices,\nsuch as gyros, accelerometers, and optical\nsystems, requires use of materials in which\ndimensional changes from metallurgical in-\nstability are limited from 10 \u00b5m/m (10 \u03bcin./ Fig 36 Comparison of distribution of yield strength\nin.). Several laboratory investigations and\nconsiderable practical experience have\nshown that wrought alloys 2024 and 6061\nand casting alloy 356.0 are well suited and\ngenerally preferred for such applications.\nDimensional changes were no greater than\n10 \u03bcm/m when alloys 2024-T851 and -T62,\n6061-T651 and -T62, and 356.0-T51, -T6,\nand -T7 were tested for more than a year at\nroom temperature and for several months at\n70 \u00b0C (160 \u00b0F), and then the same alloys\nwere tested with repeated thermal cycling\nbetween 20 and -70 \u00b0C (68 and -94 \u00b0F).\n\n\n525\n\n\n450\nYield strength, MPa\n\n\n475\n\n\n500\n\n\n400\n\n\n425\n\n\nin heat-treated 7075-T6 clad sheet product\nwith distribution in a single sheet. A is 95% probability\nthat not more than 1% of all material will fall below this\nvalue; B is 95% probability that not more than 10% of\nall material will fall below this value. (A and B refer\nonly to curve representing 4290 routine mill tests.)\n\n\nerties allow the use of tensile properties\nalone as acceptance criteria. The minimum\nguaranteed strength is ordinarily that value\nabove which it has been statistically pre-\ndicted with 95% probability that 99% or\nmore of the material will pass. The inherent\nvariability within lots and among specimens\nfrom a given piece is shown in Fig 36.\nTesting provides a check for evidence of\nconformance; process capability and pro-\ncess control are the foundations for guaran-\nteed values.\nPublished minimum guaranteed values\nare applicable only to specimens cut from a\nspecific location in the product, with their\naxes oriented at a specific angle to the\ndirection of working as defined in the appli-\ncable procurement specification. In thick\nplate, for example, the guaranteed values\napply to specimens taken from a plane\nmidway between the center and the surface,\nand their axes parallel to the width dimen-\nsion (long transverse). Different properties\nshould be expected in specimens taken from\nother locations, or in specimens whose axes\nwere parallel to thickness dimension (short\ntransverse). However, the specified \"refer-\nee\" locations and orientations do provide a\nuseful basis for lot-to-lot comparisons, and\nconstitute a valuable adjunct to other pro-\ncess-control measures.\nTensile tests can be used to evaluate the\neffects of changes in the process, provided\nspecimens are carefully selected. A varia-\ntion in process that produces above-mini-\nmum properties on test specimens, howev-\ner, is not necessarily satisfactory. Its\nacceptability can be judged only by compar-\ning the resulting properties with those de-\nveloped by the standard process on similar-\nly located specimens. Finally, variations in\nheat-treating procedure are likely to affect\n\n\nBecause stresses applied or induced by\nacceleration in such devices generally are\nnot high, strength levels lower than those of\nthe highest-strength tempers frequently are\nsatisfactory. To increase precision of ma-\nchining to intended dimensions, as well as\nto promote maximum stability, it is com-\nmon practice to apply additional thermal\ntreatments for stress relief and precipitation\nof 1 to 2 h at temperatures of 175 to 205 \u00b0C\n(350 to 400 \u00b0F) after rough machining. These\nadditional treatments sometimes are repeat-\ned at successive stages of processing, and\neven after final machining. In addition, it\nhas been claimed that one or two cyclic\ntreatments consisting of cooling to -100 \u00b0C\n(-150 \u00b0F), holding for 2 h, heating to 232 to\n240 \u00b0C (450 to 465 \u00b0F) and again holding for\n2 h can improve dimensional stability of\n356-T6 castings.\n\n\nDimensional Stability in Service\n\n\nDimensional stability of heat-treated\nparts in service depends on alloy, temper,\nand service conditions. Of the latter, ex-\ncluding mechanical conditions such as ap-\nplied loads, the most important is service\ntemperature range relative to the range in\nwhich precipitation occurs. Residual stress-\nes constitute another source of dimensional\nchanges. Stress relief minimizes changes\ndue to residual stresses, and most mill prod-\nucts usually are supplied in tempers that\ninclude stress relief. Potential dimensional\nchange as a result of further precipitation in\nparts that operate at elevated temperatures\nis minimized for wrought products by use of\nT7-type stabilizing treatments and for cast-\nings by use of T5-type treatments. Howev-\ner, components of high-precision equip-\nment, such as instruments for aerospace\nguidance systems and optical and telescopic\ndevices, may require special supplementary\ntreatments during manufacture to further\nreduce stresses or subsequent precipitation.\n(These treatments are discussed below, un-\nder \"Stability of Precision Equipment.\")\n\n\nQuality Assurance\n\n\nQuality-assurance criteria that heat-treat-\ned materials must meet always include min-\nimum tensile properties and, for certain\nalloys and tempers, adequate fracture\ntoughness and resistance to detrimental\nforms of corrosion (such as intergranular or\nexfoliation attack) or to stress-corrosion\ncracking. All processing steps through heat\ntreatment must be carefully controlled to\nensure high and reliable performance.\nTensile Tests. In general, the relatively\nconstant relationships among various prop-\n\n\nHeat Treating of Aluminum Alloys / 877\n\n\n600\n\n\nTable 13 Typical acceptable hardness values for wrought aluminum alloys\n\n\n85\n\n\nAcceptable hardness does not guarantee acceptable properties; acceptance should be based on acceptable\nhardness plus written evidence of compliance with specified heat-treating procedures. Hardness values\nhigher than the listed maximums are acceptable provided that the material is positively identified as the\ncorrect alloy.\n\n\n80\n80\n\n\n550\n\n\n75\n\n\nHardness\n\n\nTensile strength, MPa\n\n\n500\n\n\nTensile strength, ksi\n\n\nT\n\n\n70\n\n\nProduct form(a)\n\n\nAlloy and temper\n\n\nHRIST\n\n\nHRB\n\n\nHRE\n\n\nHRH\n\n\n87-95\n103-110\n104-110\n100-109\n97-106\n\n\n2014-T3, T4, -T42\n2014-T6, -T62, -T65\n\n\nAll\n\n\n65-70\n80-90\n81-90\n\n\n450\n\n\n65\n\n\nSheet(b)\nAll others\n\n\n2014-T61\n2024-T3\n\n\nAll\n\n\n60\n\n\nNot clad(c)\n\n\n69-83\n\n\n111-118\n\n\n82.5-87.5\n\n\n400\n\n\nClad, \u22641.60 mm (0.063\n\n\nT\n\n\n55\n\n\n109-116\n\n\n52-71\n\n\nin.)\n\n\n91-100\n\n\n80-84.5\n\n\nClad, 1.60 mm (0.063\n\n\n350\n\n\nin.)\n\n\n93-102\n100-110\n97-106\n\n\n109-116\n\n\n52-71\n76-90\n69-83\n\n\n50\n\n\n2024-T36\n\n\nAll\n\n\n85-90\n82.5-87.5\n\n\n2024-T4, -T42(d)\n\n\nNot clad\n\n\n111-118\n\n\n45\n\n\nClad, 1.60 mm (0.063\nin.)\nClad, >1.60 mm (0.063\n\n\n300\n\n\n80-84.5\n\n\n91-100\n\n\n90\n\n\n100\n\n\n52-71\n\n\n109-116\n\n\n70\nHardness, HRB\n\n\n80\n\n\n50\n\n\n60\n\n\n93-102\n99-106\n99-106\n99-106\n105-110\n79-87\n60-75\n70-81\n\n\nin.)\n\n\n52-71\n74.5-83.5\n74.5-83.5\n\n\n109-116\n\n\nTensile strength versus hardness for various\naluminum alloys and tempers\n\n\nFig 37\n\n\n2024-T6, -T62\n\n\nAll\n\n\n84-88\n84-88\n\n\n2024-T81\n\n\nNot clad\n\n\nClad\nAll\nAll\n\n\n87.5-90\n74.5-78.5\n64-75\n67-78\n\n\n83-90\n\n\n2024-T86\n6053-T6\n\n\n\u2022 Use a specimen that has at least 19 cm\u00b2 (3\nin.2) of surface area\n\n\n88-100\n82-103\n\n\n6061-T4(d)\n\n\nSheet\n\n\nExtrusions; bar\n\n\nRemove any cladding by filing or etching\nClean the specimen by immersing it for 1\nmin in a solution containing 5% concen-\ntrated nitric acid and 0.5% hydrofluoric\nacid at a temperature of 95 \u00b0C (200 \u00b0F);\nrinse in distilled water. Immerse for 1 min\nin concentrated nitric acid at room tem-\nperature; rinse in distilled water\n\n\n\u2022\n\n\nNot clad, 0.41 mm\n\n\n\u2022\n\n\n6061-T6\n\n\n75-84\n\n\n(0.016 in.)\n\n\nNot clad, \u22650.51 mm\n\n\n(0.020 in.)\n\n\n78-84\n\n\n85-97\n84-96\n55-70\n70-85\n91-102\n106-114\n\n\n47-72\n\n\nClad\nAll\nAll\nAll\n\n\n6063-T5\n6063-T6\n6151-T6\n\n\n89-97\n\n\n62.5-70\n\n\n7075-T6, -T65\n\n\nNot clad(e)\n\n\n85-94\n\n\n87.5-92\n\n\nClad:\n\n\n\u2022 Immerse the specimen for 6 h in a freshly\nprepared solution containing 57 g of sodi-\num chloride and 10 mL of 30% hydrogen\nperoxide per liter of water at a tempera-\nture of 30 \u00b1 5 \u00b0C (86 \u00b1 9 \u00b0F). More than\none specimen may be corroded in the\nsame container provided that at least 4.6\nmL of solution is used for each square\ncentimeter (30 mL/in.\u00b2) of specimen sur-\nface and that the specimens are electrical-\nly insulated from each other\n\u2022 After the immersion period, wash the\nspecimen with a soft-bristle brush to re-\nmove any loose corrosion product. Cut a\ncross-sectional specimen at least 19 mm\n(3/4 in.) long through the most severely\ncorroded area; mount and metallographi-\ncally polish this specimen\nExamine the cross-sectional specimen mi-\ncroscopically at magnifications of 100\u00d7\nand 500\u00d7 both before and after etching\nwith Keller's reagent\n\u2022 Describe the results of the microscopic\nexamination in terms of the five degrees\nof severity of intergranular attack illus-\ntrated in Fig 38\n\n\n86-90\n\n\n\u22640.91 mm (0.036 in.)\n\n\n102-110\n\n\n>0.91 1.27 mm\n\n\n(>0.036 0.050 in.)\n\n\n78-90\n\n\n104-110\n\n\n>1.27 1.57 mm\n\n\n104-110\n\n\n(>0.050 \u2264 0.062 in.)\n\n\n76-90\n\n\n>1.57 1.78 mm\n\n\n(>0.062\u2264 0.070 in.)\n\n\n76-90\n73-90\n81-93\n85 min\n\n\n102-110\n102-110\n104-114\n105 min\n\n\n>1.78 mm (0.070 in.)\n\n\n7079-T6, -T65\n7178-T6\n\n\nAll(e)\n\n\n87.5-92\n88 min\n\n\nNot clad(f)\n\n\nClad:\n\n\n\u22640.91 mm (0.036 in.)\n\n\n86 min\n\n\n102 min\n\n\n>0.91 1.57 mm\n\n\n(>0.036\u2264 0.062 in.)\n>1.57 mm (0.062 in.)\n\n\n85 min\n88 min\n\n\n(a) Minimum hardness values shown for clad products are valid for thicknesses up to and including 2.31 mm (0.091 in.); for heavier-gage\nmaterial, cladding should be locally removed for hardness testing or test should be performed on edge of sheet. (b) 126 to 158 HB (10\nmm ball, 500 kg load). (c) 100 to 130 HB (10 mm ball, 500 kg load). (d) Alloys 2024-T4, 2024-T42 and 6061-T4 should not be rejected for\nlow hardness until they have remained at room temperature for at least three days following solution treatment. (e) 136 to 164 HB (10\nmm ball, 500 kg load). (f) 136 HB min (10 mm ball, 500 kg load)\n\n\n\u2022\n\n\nthe relationships among tensile properties\nand other mechanical properties. In appli-\ncations where other properties are more\nimportant than tensile properties, the other\nproperties should be checked also.\nHardness tests are less valuable for ac-\nceptance and rejection of heat-treated alu-\nminum alloys than they are for steel. Nev-\nertheless, hardness tests have some utility\nfor process control. Typical hardness val-\nues for various alloys and tempers are given\nin Table 13. Figure 37 shows the general\nrelationship between longitudinal tensile\nstrength and hardness for aluminum alloys.\nIntergranular-Corrosion Test. The extent\nof precipitation during elevated-temperature\naging of alloys 2014, 2219, and 2024 markedly\n\n\ninfluences the type of corrosion attack and\nthe corrosion resistance. With thin-section\nproducts quenched at rates sufficiently rapid\nto prevent precipitation in the grain bound-\naries during the quench, short periods of\nprecipitation heat treating produce localized\ngrain boundary precipitates adjacent to the\ndepleted areas, producing susceptibility to\nintergranular corrosion. Additional heating,\nhowever, induces extensive general precipi-\ntation within the grains, lowering the corro-\nsion potential differences between the grains\nand the boundary areas, thus removing the\ncause of the selective corrosion.\nThe most common test for susceptibility\nto intergranular corrosion is carried out as\nfollows:\n\n\nElectrical Conductivity. For control of the\ncorrosion and stress-corrosion characteris-\ntics of certain tempers, notably the T73 and\nT76 types, the materials must meet combi-\nnation criteria of yield strength plus electri-\ncal conductivity. Although these criteria are\nbased on indirect measurements of proper-\nties, their validity for ensuring the intended\ncorrosion and stress-corrosion resistance\n\n\n878 / Heat Treating of Nonferrous Alloys\n\n\nmay differ from those employed to produce\nthe same temper in another alloy.\nDesignations for the common heat-treat-\ned tempers, and descriptions of the se-\nquences of operations used to produce\nthose tempers, are given in the following\nparagraphs. (For the entire aluminum alloy\ntemper designation system, including desig-\nnations for non-heat-treatable alloys, see\nVolume 2 of this Metals Handbook series.)\nBasic temper designations for heat-treated\nconditions include the codes O, W, and T.\nOther basic temper designations are F (as\nfabricated) and H (strain hardened).\nO, annealed. Applies to wrought prod-\nucts that are annealed to obtain lowest\nstrength temper and to cast products that\nare annealed to improve ductility and di-\nmensional stability. The O may be followed\nby a digit other than zero.\nW, solution heat treated. An unstable\ntemper applicable to any alloy that naturally\nages (spontaneously ages at room tempera-\nture) after solution heat treatment. This\ndesignation is specific only when the period\nof natural aging is indicated-for example,\nWh. (See also the discussion of the Tx51,\nTx52, and Tx54 tempers, in the section\nbelow on subdivision of the T temper.)\nT, heat treated to produce stable tempers\nother than O. Applies to products that are\nthermally treated, with or without supple-\nmentary strain hardening, to produce stable\ntempers. The T is always followed by one or\nmore digits, as discussed below.\nMajor Subdivisions of T Temper. In T-type\ndesignations, the T is followed by a number\nfrom 1 to 10; each number denotes a specif-\nic sequence of basic treatments, as de-\nscribed below.\nTI, cooled from an elevated-temperature\nshaping process and naturally aged to a\nsubstantially stable condition. Applies to\nproducts that are not cold worked after an\nelevated-temperature shaping process such\nas casting or extrusion, and for which me-\nchanical properties have been stabilized by\nroom-temperature aging. If the products are\nflattened or straightened after cooling from\nthe shaping process, the effects of the cold\nwork imparted by flattening or straightening\nare not recognized in specified property\nlimits.\nT2, cooled from an elevated-temperature\nshaping process, cold worked, and natural-\nly aged to a substantially stable condition.\nApplies to products that are cold worked\nspecifically to improve strength after cool-\ning from a hot-working process such as\nrolling or extrusion, and for which mechan-\nical properties have been stabilized by\nroom-temperature aging. The effects of cold\nwork, including any cold work imparted by\nflattening or straightening, are recognized in\nspecified property limits.\nT3, solution heat treated, cold worked,\nand naturally aged to a substantially stable\ncondition. Applies to products that are cold\n\n\nCondition\n\n\nProduct form\n\n\nAlloy\n\n\nSheet and plate\nSheet and plate\n\n\n** EE\nT8\nT8\nT7\nT7\n\n\n2048\n2124\n2419\n\n\nT3, T8\n\n\nSheet, plate, extrusions, and\nforgings\n\n\n(a)\n\n\n(b)\n\n\n7049\n7050\n\n\nPlate, forgings, and extrusions\nSheet, plate, forgings, and\nextrusions\n\n\nSheet and plate\nSheet, plate, forgings, and\nextrusions\nSheet and plate\n\n\n7150\n7175\n\n\nT6\n\n\nT6, T7\n\n\n(c)\n\n\n(d)\n\n\n7475\n\n\nT6, T7\n\n\nThe fracture toughness of these alloys and\ntempers range in measured K\u2081c values from\nabout 20 MPa\u221a\u221a\u221am (18 ksi\u221ain.) upward. Con-\ntrolled-toughness alloys are often derivatives\nof conventional alloys. For example, 7475\nalloy is a derivative of 7075 with maximum\ncompositional limits on some elements that\nwere found to decrease toughness.\nIn products of the newer controlled-\ntoughness high-strength alloys 2090, 2091,\n2124, 2224, 2324, 7050, 7149, 7150, 7175,\n7475, and 8090, which provide guaranteed\nlevels of fracture toughness, minimum val-\nues of the applicable indices, K\u2081c or Kc,\nare established by accumulation of statis-\ntical data from production lots as a basis\nfor guaranteed minimum values. If the\nminimum specified fracture toughness val-\nue is not attained, the material is not\nacceptable. Some specifications allow use\nof less-expensive screening tests (such as\nthe notch tensile or chevron-notched short\nbar) as a basis for release of high-tough-\nness alloy products. In these instances,\ncorrelations between K\u2081 and the screening\ntest result is used to establish the appro-\npriate notch-yield ratio as a lot-release\ncriterion.\n\n\n(e)\n\n\nFive degrees of severity of intergranular at-\nFig 38\ntack. Severity of intergranular attack (sche-\nmatic), as observed microscopically in transverse sec-\ntions after test for susceptibility to intergranular\ncorrosion. Top of each area shown in surface exposed\nto corrosive solution\n\n\nhas been firmly established by extensive\ncorrelation and testing.\nLow tensile strengths may be accompa-\nnied by high levels of electrical conductiv-\nity, so electrical conductivity is sometimes\nused as a quality-assurance diagnostic\ntool. However, because the correlation\nbetween strength and electrical conductiv-\nity is strongly a function of chemical com-\nposition and fabricating practice, use of\nelectrical conductivity is not recommend-\ned except for rough screening. This\nscreening must be followed by hardness\ntesting, and then by tensile testing if the\nhardness tests indicate that the heat treat-\nment was suspect.\nFracture Toughness Indices. Fracture\ntoughness is rarely, if ever, a design con-\nsideration in the 1000, 3000, 4000, 5000,\nand 6000 series alloys. The fracture tough-\nness of these alloys is sufficiently high that\nthicknesses beyond those commonly pro-\nduced would be required to obtain a valid\ntest.\nFracture toughness is a meaningful de-\nsign-related parameter for some conven-\ntional high-strength alloys and all the con-\ntrolled-toughness, high-strength alloys.\nConventional aerospace alloys for which\nfracture toughness minimums may be use-\nful in design include 2014, 2024, 2219,\n7075, and 7079. These alloys have tough-\nness levels that are inferior to those of\ntheir controlled-toughness counterparts.\nConsequently, these products are not used\nin fracture-critical applications, although\nfracture toughness can be a meaningful\ndesign parameter. Fracture toughness is\nnot guaranteed in conventional high\nstrength alloys.\nFracture toughness quality control and\nmaterial procurement minimums are appro-\npriate for controlled-toughness, high-\nstrength alloys. The alloys and tempers\ncurrently identified as controlled-tough-\nness, high-strength products include:\n\n\nTemper Designations for\nHeat-Treatable Aluminum Alloys\n\n\nThe temper designations used in the Unit-\ned States for heat-treatable aluminum alloys\nare part of the system that has been adopted\nas an American National Standard (ANSI\nH35.1). Used for all wrought and cast prod-\nuct forms except ingot, the system is based\non the sequences of mechanical or thermal\ntreatments, or both, used to produce the\nvarious tempers. The temper designation\nfollows the alloy designation and is separat-\ned from it by a hyphen. Basic temper des-\nignations consist of individual capital let-\nters. Major subdivisions of basic tempers,\nwhere required, are indicated by one or\nmore digits following the letter. These digits\ndesignate specific sequences of treatments\nthat produce specific combinations of char-\nacteristics in the product. Variations in\ntreatment conditions within major subdivi-\nsions are identified by additional digits. The\nconditions during heat treatment (such as\ntime, temperature, and quenching rate)\nused to produce a given temper in one alloy\n\n\nHeat Treating of Aluminum Alloys / 879\n\n\nTemper designations T42 and T62 have\nbeen assigned to wrought products heat\ntreated from the O or the F temper to\ndemonstrate response from the heat treat-\nment described below. Temper designations\nT42 and T62 also may be applied to wrought\nproducts heat treated from any temper by\nthe user when such heat treatment results in\nthe mechanical properties applicable to\nthese tempers.\n\n\nworked specifically to improve strength af-\nter solution heat treatment, and for which\nmechanical properties have been stabilized\nby room-temperature aging. The effects of\ncold work, including any cold work impart-\ned by flattening or straightening, are recog-\nnized in specified property limits.\nT4, solution heat treated and naturally\naged to a substantially stable condition.\nApplies to products that are not cold\nworked after solution heat treatment, and\nfor which mechanical properties have been\nstabilized by room-temperature aging. If the\nproducts are flattened or straightened, the\neffects of the cold work imparted by flatten-\ning or straightening are not recognized in\nspecified property limits.\nT5, cooled from an elevated-temperature\nshaping process and artificially aged. Ap-\nplies to products that are not cold worked\nafter an elevated-temperature shaping pro-\ncess such as casting or extrusion, and for\nwhich mechanical properties or dimensional\nstability, or both, have been substantially\nimproved by precipitation heat treatment. If\nthe products are flattened or straightened\nafter cooling from the shaping process, the\neffects of the cold work imparted by flatten-\ning or straightening are not recognized in\nspecified property limits.\nT6, solution heat treated and artificially\naged. Applies to products that are not cold\nworked after solution heat treatment, and\nfor which mechanical properties or dimen-\nsional stability, or both, have been substan-\ntially improved by precipitation heat treat-\nment. If the products are flattened or\nstraightened, the effects of the cold work\nimparted by flattening or straightening are\nnot recognized in specified property limits.\nT7, solution heat treated and stabilized.\nApplies to products that have been precip-\nitation heat treated to the extent that they\nare overaged. Stabilization heat treatment\ncarries the mechanical properties beyond\nthe point of maximum strength to provide\nsome special characteristic, such as en-\nhanced resistance to stress-corrosion crack-\ning or to exfoliation corrosion.\nT8, solution heat treated, cold worked,\nand artificially aged. Applies to products\nthat are cold worked specifically to improve\nstrength after solution heat treatment, and\nfor which mechanical properties or dimen-\nsional stability, or both, have been substan-\ntially improved by precipitation heat treat-\nment. The effects of cold work, including\nany cold work imparted by flattening or\nstraightening, are recognized in specified\nproperty limits.\nT9, solution heat treated, artificially\naged, and cold worked. Applies to products\nthat are cold worked specifically to improve\n\n\nstrength after they have been precipitation\nheat treated.\nT10, cooled from an elevated-tempera-\nture shaping process, cold worked, and\nartificially aged. Applies to products that\nare cold worked specifically to improve\nstrength after cooling from a hot-working\nprocess such as rolling or extrusion, and for\nwhich mechanical properties or dimensional\nstability, or both, have been substantially\nimproved by precipitation heat treatment.\nThe effects of cold work, including any cold\nwork imparted by flattening or straighten-\ning, are recognized in specified property\nlimits.\nOther Subdivisions of T Temper Codes for\nStress-Relieved Products. When it is desir-\nable to identify a variation of one of the ten\nmajor T tempers described above, addition-\nal digits, the first (x) of which cannot be\nzero, may be added to the designation.\nThe following specific sets of additional\ndigits have been assigned to stress-relieved\nwrought products.\nTx51, stress relieved by stretching. Ap-\nplies to the following products when\nstretched to the indicated amounts after\nsolution heat treatment or after cooling\nfrom an elevated-temperature shaping pro-\n\n\n\u2022 T42. Solution heat treated from the O or\nthe F temper to demonstrate response to\nheat treatment and naturally aged to a\nsubstantially stable condition\n\u2022 T62. Solution heat treated from the O or\nthe F temper to demonstrate response to\nheat treatment and artificially aged\n\n\nSubdivision of the O Temper. In temper\ndesignations for annealed products, a digit\nfollowing the O indicates special character-\nistics. For example, 01 denotes that a prod-\nuct has been heat treated according to a\ntime/temperature schedule approximately\nthe same as that used for solution heat\ntreatment, and then air cooled to room\ntemperature, to accentuate ultrasonic re-\nsponse and provide dimensional stability;\nthis designation applies to products that are\nto be machined prior to solution heat treat-\nment by the user.\n\n\ncess:\n\n\nPermanent set, %\n\n\nProduct form\n\n\n112-3\n\n\nPlate\n\n\nREFERENCES\n\n\nRod, bar, shapes, extruded\ntube\nDrawn tube\n\n\n1-3\n1/2-3\n\n\n1. S. Hirano et al., Quench Sensitivity in\nAl-Li Based Alloys, Proceedings of Con-\nference on Aluminum-Lithium Alloys\n(Vol 1), Materials and Component Engi-\nneering Publications, 1989, p 335-344\n2. T. Sheppard, Mater. Sci. Technol., Vol\n4, July 1988, p 636\n3. J.E. Hatch, in Aluminum Properties and\nPhysical Metallurgy, American Society\nfor Metals, 1984, p 165-166\n4. C.E. Bates, Selecting Quenchants to\nMaximize Tensile Properties and Mini-\nmize Distortion in Aluminum Parts, J.\nHeat Treat., Vol 5 (No. 1), 1987, p 27-40\n5. T. Croucher, Critical Parameters for\nEvaluating Polymer Quenching of Alu-\nminum, Heat Treat., Vol 19 (No. 12),\nDec 1987, p 21-25\n6. W.L. Fink and L.A. Willey, Quenching\nof 75S Aluminum Alloy, Trans. AIME,\nVol 175, 1948, p 414-427\n7. J.W. Evancho and J.T. Staley, Kinetics\nof Precipitation in Aluminum Alloys dur-\ning Continuous Cooling, Metall. Trans.\nA, Vol 5A, Jan 1974, p 43-47\n8. J.T. Staley, Industrial Heating XLIV,\nOct 1977, p 6-9\n9. G.F. Bobart, J. Heat Treat., Vol 6 (No.\n1), 1988, p 47-52\n\n\nTx51 applies directly to plate and to rolled\nor cold finished rod and bar. These products\nreceive no further straightening after\nstretching. Tx51 also applies to extruded\nrod, bar, shapes, and tubing, and to drawn\ntubing, when designated as follows:\n\n\nTx510. Products that receive no further\nstraightening after stretching\n\n\n\u2022\n\n\n\u26ab Tx511. Products that may receive minor\nstraightening after stretching to comply\nwith standard tolerances\n\u26ab Tx52. Stress relieved by compressing. Ap-\nplies to products that are stress relieved\nby compressing after solution heat treat-\nment, or after cooling from a hot-working\nprocess to produce a permanent set of 1\nto 5%\n\u2022 Tx54. Stress relieved by combining\nstretching and compressing. Applies to\ndie forgings that are stress relieved by\nrestriking cold in the finish die. (These\nsame digits and 51, 52, and 54-may be\nadded to the designation W to indicate\nunstable solution heat-treated and stress-\nrelieved tempers)\n"}, "expected_output": {"claims": [{"unit": "\u00b0C", "value": 107, "evidence": ["(h) Two-stage treatment comprised of 6 to 8 h at 107 \u00b0C (225 \u00b0F)", "(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water,\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) Stress relieved by stretching to produce a specified amount of permanent set after solution treatment and\nprior to precipitation heat treatment. (e) No solution heat treatment; 72 h at room temperature following press quench, followed by two-stage precipitation heat treatment comprised of 8 h at 107 \u00b0C (225 \u00b0F)\nplus 16 h at 149 \u00b0C (300 \u00b0F). (f) Aging practice varies with product, size, nature of equipment, loading procedures, and furnace-control capabilities. The optimum practice for a specific item can be ascertained\nonly by actual trial treatment of the item under specific conditions. Typical procedures involve a two-stage treatment comprised of 3 to 30 h at 121 \u00b0C (250 \u00b0F) followed by 15 to 18 h at 163 \u00b0C (325 \u00b0F) for\nextrusions. An alternative two-stage treatment of 8 h at 99 \u00b0C (210 \u00b0F) followed by 24 to 28 h at 163 \u00b0C (325 \u00b0F) also may be used. (g) Aging of aluminum alloys 7050, 7075, 7175, and 7475 from any temper\nto the T73 or T76 temper series requires closer-than-normal controls on aging variables such as time, temperature, heatup rate, and so forth, for any given item. In addition, when material in a T6-type temper\nis reaged to a T73- or T76-type temper, the specific condition of the T6 material (such as property levels and other effects of processing variables) is extremely important and will affect the capability of the\nreaged material to conform to the requirements specified for the applicable T73- or T76-type temper. (h) Two-stage treatment comprised of 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by: 24 to 30 h at 163 \u00b0C (325 \u00b0F)\nfor sheet and plate: 8 to 10 h at 177 \u00b0C (350 \u00b0F) for rolled or cold finished rod and bar; 6 to 8 h at 177 \u00b0C (350 \u00b0F) for extrusions and tube; 8 to 10 h at 177 \u00b0C (350 \u00b0F) for forgings in the T73 temper; and 6\nto 8 h at 177 \u00b0C (350 \u00b0F) for forgings in the T7352 temper. (i) These heat treatments also apply to alclad sheet and plate of these alloys. (j) An alternative two-stage treatment comprised of 4 h at 96 \u00b0C (205 \u00b0F)\nfollowed by 8 h at 157 \u00b0C (315 \u00b0F) also may be used. (k) For sheet, plate, tube, and extrusions, an alternative two-stage treatment comprised for 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by 14 to 18 h at 168 \u00b0C\n(335 \u00b0F) may be used, provided that a heatup rate of approximately 14 \u00b0C/h (25 \u00b0F/h) is employed. For rolled or cold finished rod and bar, the alternative treatment is 10 h at 177 \u00b0C (350 \u00b0F). (I) An alternative\nthree-stage treatment comprised of 5 h at 99 \u00b0C (210 \u00b0F), 4 h at 121 \u00b0C (250 \u00b0F), and then 4 h at 149 \u00b0C (300 \u00b0F) may also be used. (m) Quenched in water at 60 to 80 \u00b0C (140 to 180 \u00b0F). (n) Stress relieved\nby 1 to 5% cold reduction after solution treatment and prior to precipitation heat treatments. (o) 7175-T74 and -T7452 heat treatments are directed to specific results, may vary from supplier to supplier and\nare either proprietary or patented. (p) Must be preceded by soak at 466 to 477 \u00b0C (870 to 890 \u00b0F). See U.S. Patent 3,791,880.\n", "Two-stage treatment comprised of 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by:"]}, {"unit": "\u00b0C", "value": 163, "evidence": ["24 to 30 h at 163 \u00b0C (325 \u00b0F) for sheet and plate", "(a) Material should be quenched from the solution-treating temperature as rapidly as possible and with minimum delay after removal from the furnace. When material is quenched by total immersion in water,\nunless otherwise indicated, the water should be at room temperature, and should be suitably cooled so that it remains below 38 \u00b0C (100 \u00b0F) during the quenching cycle. Use of high-velocity, high-volume\njets of cold water also is effective for some materials. (b) The nominal temperatures listed should be attained as rapidly as possible and maintained within \u00b16 \u00b0C (\u00b110 \u00b0F) of nominal during the time at\ntemperature. (c) Approximate time at temperature. The specific time will depend on the time required for the load to reach temperature. The times shown are based on rapid heating, with soak time measured\nfrom the time the load reaches a temperature within 6 \u00b0C (10 \u00b0F) of the applicable temperature. (d) Stress relieved by stretching to produce a specified amount of permanent set after solution treatment and\nprior to precipitation heat treatment. (e) No solution heat treatment; 72 h at room temperature following press quench, followed by two-stage precipitation heat treatment comprised of 8 h at 107 \u00b0C (225 \u00b0F)\nplus 16 h at 149 \u00b0C (300 \u00b0F). (f) Aging practice varies with product, size, nature of equipment, loading procedures, and furnace-control capabilities. The optimum practice for a specific item can be ascertained\nonly by actual trial treatment of the item under specific conditions. Typical procedures involve a two-stage treatment comprised of 3 to 30 h at 121 \u00b0C (250 \u00b0F) followed by 15 to 18 h at 163 \u00b0C (325 \u00b0F) for\nextrusions. An alternative two-stage treatment of 8 h at 99 \u00b0C (210 \u00b0F) followed by 24 to 28 h at 163 \u00b0C (325 \u00b0F) also may be used. (g) Aging of aluminum alloys 7050, 7075, 7175, and 7475 from any temper\nto the T73 or T76 temper series requires closer-than-normal controls on aging variables such as time, temperature, heatup rate, and so forth, for any given item. In addition, when material in a T6-type temper\nis reaged to a T73- or T76-type temper, the specific condition of the T6 material (such as property levels and other effects of processing variables) is extremely important and will affect the capability of the\nreaged material to conform to the requirements specified for the applicable T73- or T76-type temper. (h) Two-stage treatment comprised of 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by: 24 to 30 h at 163 \u00b0C (325 \u00b0F)\nfor sheet and plate: 8 to 10 h at 177 \u00b0C (350 \u00b0F) for rolled or cold finished rod and bar; 6 to 8 h at 177 \u00b0C (350 \u00b0F) for extrusions and tube; 8 to 10 h at 177 \u00b0C (350 \u00b0F) for forgings in the T73 temper; and 6\nto 8 h at 177 \u00b0C (350 \u00b0F) for forgings in the T7352 temper. (i) These heat treatments also apply to alclad sheet and plate of these alloys. (j) An alternative two-stage treatment comprised of 4 h at 96 \u00b0C (205 \u00b0F)\nfollowed by 8 h at 157 \u00b0C (315 \u00b0F) also may be used. (k) For sheet, plate, tube, and extrusions, an alternative two-stage treatment comprised for 6 to 8 h at 107 \u00b0C (225 \u00b0F) followed by 14 to 18 h at 168 \u00b0C\n(335 \u00b0F) may be used, provided that a heatup rate of approximately 14 \u00b0C/h (25 \u00b0F/h) is employed. For rolled or cold finished rod and bar, the alternative treatment is 10 h at 177 \u00b0C (350 \u00b0F). (I) An alternative\nthree-stage treatment comprised of 5 h at 99 \u00b0C (210 \u00b0F), 4 h at 121 \u00b0C (250 \u00b0F), and then 4 h at 149 \u00b0C (300 \u00b0F) may also be used. (m) Quenched in water at 60 to 80 \u00b0C (140 to 180 \u00b0F). (n) Stress relieved\nby 1 to 5% cold reduction after solution treatment and prior to precipitation heat treatments. (o) 7175-T74 and -T7452 heat treatments are directed to specific results, may vary from supplier to supplier and\nare either proprietary or patented. (p) Must be preceded by soak at 466 to 477 \u00b0C (870 to 890 \u00b0F). See U.S. Patent 3,791,880.\n"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_e1300834056cb7e6"}}
{"id": "996c9b812b13f84f391a1773", "input": {"query": "What is the recycled content percentage in the SSAB cold-rolled steel sheets? Please provide numerical percentages.", "source_url": "https://www.ssab.com/-/media/files/en/support/certificates/epd-ssab-cold-rolled-steel-sheets-and-slit-coils.pdf?m=20250616123413", "document_text": "ECO PLATFORM\n\n\n'EPD\u24c7\n\n\nEPD\n\n\nVERIFIED\n\n\nEnvironmental Product Declaration\n\n\nCold rolled steel\nsheets and slit coils\n\n\nEPD of multiple products, based on the average results of the product group\n\n\nSSAB\n\n\nIn accordance with ISO 14025:2006 and EN 15804:2012+A2:2019/AC:2021\n\n\nProgramme: The International EPD\u00ae System, www.environdec.com\nProgramme operator: EPD International AB\n\n\nEPD owner: SSAB Europe Oy\n\n\nEPD registration number: EPD-IES-0023966\nPublication date: 2025-06-16\n\n\nValid until: 2030-06-04\n\n\nContents\n\n\n1. General information\n\n\n3\n\n\n1.1 SSAB's vision - A stronger, lighter and more sustainable world\n1.2 Company information\n\n\n4\n\n\n2. Product information\n\n\n4\n\n\n2.1 Product technical information and applications\n\n\n4\n\n\n2.2 Product description\n\n\n4\n\n\n2.3 Labeling and packaging\n\n\n\u0441\u043b\n\n\n3. Production and transportation\n\n\n\u0441\u043b\n5\n\n\n3.1 Production sites\n\n\n5\n\n\n3.2 Transportation\n\n\n\u0441\u043b\n5\n\n\n10\n\n\n4. LCA\n\n\n160\n\n\n4.1 LCA information\n\n\n4.2 Product content declaration\n\n\n8\n\n\n4.3 Environmental performance indicators results\n\n\n8\n\n\n5. References\n\n\n11\n\n\n2\n\n\n1. General information\n\n\nPROGRAM INFORMATION\n\n\nThe International EPD\u24c7 System\n\n\nProgram:\n\n\nEPD International AB\n\n\nAddress:\n\n\nBox 210 60\n\n\nSE-100 31 Stockholm\n\n\nSweden\n\n\nWebsite:\n\n\nwww.environdec.com\n\n\ninfo@environdec.com\n\n\nEmail:\n\n\nAccountabilities for PCR, LCA and independent, third-party verification\n\n\nProduct Category Rules (PCR)\n\n\nCore product category rules: CEN standard EN 15804 serves as the core PCR.\n\n\nProduct category rules: PCR 2019:14 Construction products. Version 1.3.4. Date 2024-04-30.\n\n\nProduct group classification: UN CPC 412.\n\n\nPCR review was conducted by: The Technical Committee of the International EPD System. See www.environdec.com for a list of members.\nReview chair: Claudia A. Pe\u00f1a, University of Concepci\u00f3n, Chile. The review panel may be contacted via the Secretariat www.environdec.com/contact.\nLife Cycle Assessment (LCA)\n\n\nLCA accountability: Lisa Hallberg, IVL Swedish Environmental Research Institute.\n\n\nThird-party verification\n\n\nIndependent third-party verification of the declaration and data, according to ISO 14025:2006, via:\n\n\n> EPD verification by individual verifier\n\n\nThird-party verifier: David Althoff Palm, Dalemarken AB.\n\n\nApproved by: The International EPD\u00ae System\n\n\nProcedure for follow-up of data during EPD validity involves third party verifier:\n\n\n> Yes \u2610 No\n\n\n[Procedure for follow-up the validity of the EPD is at minimum required once a year with the aim of confirming whether the information in the EPD\nremains valid or if the EPD needs to be updated during its validity period. The follow-up can be organized entirely by the EPD owner or together with\nthe original verifier via an agreement between the two parties. In both approaches, the EPD owner is responsible for the procedure being carried\nout. If a change that requires an update is identified, the EPD shall be re-verified by a verifier]\n\n\nThe EPD owner has the sole ownership, liability, and responsibility for the EPD.\n\n\nEPDS within the same product category but registered in\ndifferent EPD programmes, or not compliant with\nEN 15804, may not be comparable. For two EPDs to\nbe comparable, they must be based on the same PCR\n(including the same version number) or be based on\nfully-aligned PCRs or versions of PCRs; cover products\nwith identical functions, technical performances and use\n(e.g. identical declared/functional units); have equivalent\n\n\nsystem boundaries and descriptions of data; apply\nequivalent data quality requirements, methods of data\ncollection, and allocation methods; apply identical cut-\noff rules and impact assessment methods (including\nthe same version of characterization factors); have\nequivalent content declarations; and be valid at the\ntime of comparison. For further information about\ncomparability, see EN 15804 and ISO 14025.\n\n\n3\n\n\n1.1 SSAB'S VISION - A STRONGER,\nLIGHTER AND MORE SUSTAINABLE\nWORLD\n\n\nDescription of the organizations:\n\n\nSSAB Europe is responsible for strip, heavy plate, and\ntubular products in Europe as well as for the global\nbusiness in the Automotive customer segment. SSAB\nEurope is also responsible for color coated products.\n\n\n\u2022\n\n\nSSAB is a global steel company with a leading position in\nhigh-strength steels and related services. The company\nis a frontrunner in the green transformation of the steel\nindustry and aims to largely eliminate carbon dioxide\nemissions from its operations and together with suppliers\nand customers create a fossil-free value chain.\nSSAB's production sites are in Sweden, Finland and\nthe USA and have an annual crude steel production\ncapacity of 8.8 million tonnes. SSAB Europe is responsible\nfor sales of strip, heavy plate, and tubular products\nin Europe as well as for the global business in the\nAutomotive customer segment. SSAB Special Steels\nhas global responsibility for sales of SSAB's quenched\nand tempered (Q&T) steels and advanced high-strength\nsteels (AHSS). SSAB Americas is the largest heavy plate\nproducer in North America and has a strong position\nbased on cost efficiency and quality.\n\n\nName and location of production sites:\n\u26ab SSAB EMEA AB (Lule\u00e5, Sweden):\nSvart\u00f6v\u00e4gen 20, 974 37 Lule\u00e5 (Sweden).\n\n\nSSAB EMEA AB (Oxel\u00f6sund, Sweden):\nAspleden 1, 613 80 Oxel\u00f6sund (Sweden).\n\n\n\u2022\n\n\nSSAB EMEA AB (Borl\u00e4nge, Sweden):\n\n\n\u2022\n\n\nKontorsviksv\u00e4gen 1, 781 84 Borl\u00e4nge (Sweden).\n\n\n\u2022 SSAB Europe Oy (Raahe, Finland):\n\n\n\u2022\n\n\nRautaruukintie 155, 92100 Raahe (Finland).\n\n\nSSAB Europe Oy (H\u00e4meenlinna, Finland):\nHarvialantie 420, 13300 H\u00e4meenlinna (Finland).\n\n\nCertifications:\n\n\nCertificates applicable to SSAB sites are ISO 14001 and\nISO 9001.\n\n\n1.2 COMPANY INFORMATION\n\n\nContact:\n\n\nEPD owner:\n\n\nEPDssab@ssab.com.\n\n\nSSAB Europe Oy, Kaisa Ahvonen, Harvialantie 420,\n13300 H\u00e4meenlinna, Finland.\n\n\n2. Product information\n\n\nFor more detailed information about technical product\nproperties and the product portfolio, please visit\nwww.ssab.com.\n\n\n2.1 PRODUCT TECHNICAL INFORMATION\nAND APPLICATIONS\n\n\nSSAB specializes in materials for demanding applications\nwhere high strength and formability are needed for weight\nsavings and increased durability. Cold rolled steel sheets\nand slit coils are used in many industries and applications,\nincluding automotive industry, construction, light\nengineering, domestic and electrical appliances, heating\nand air conditioning equipments, storage systems, and\ntubular products.\n\n\n2.2 PRODUCT DESCRIPTION\n\n\nThe scope of this EPD is SSAB cold rolled blast furnace-\nbased steel sheets and slit coils.\n\n\nThe steel is an alloy of mainly iron and carbon and may\ncontain other alloying metals and trace elements. These\nalloying elements improve the chemical and physical\nproperties of steel, such as strength, ductility, durability,\nand corrosion resistance.\n\n\nSSAB offers a comprehensive selection of steel\nproducts that includes cold-forming steels, advanced\nhigh-strength steels, complex phase steels, martensitic\nsteels, ultrahigh-strength steels, weather-resistant steels\nand hardenable boron steels.\nCold rolled steel sheets and slit coils are produced in\na thickness range of 0.40-3.0 mm.\n\n\nThe exact composition of the steel manufactured by\nSSAB depends on product requirements, either based\non national and/or international standards, such as\nEN 10130, EN 10268, EN 10338, VDA 239-100, on\ncustomer-specific and/or other OEM standards.\nSSAB's unique products also have their own specific\nrequirements.\n\n\nThe products are often customized to meet national\nand/or international standards as well as customer-\nspecific or other Original Equipment Manufacturer (OEM)\nstandards. Besides standardized steel grades, SSAB's\ncold rolled product portfolio also includes products unique\nto SSAB and which in some cases may be patented.\n\n\nContent declaration and average chemical composition\nare presented in section 4.2. More detailed information on\nthe different steel compositions is available from national\nand international standards, and on www.ssab.com.\n\n\n4\n\n\n2.3 LABELING AND PACKAGING\n\n\nSSAB products are labeled to be easily identifiable and\ntraceable. The packaging and protection type of SSAB\nsteel products is specified when ordering.\n\n\nfor cut-to-lengths packaging. The bundles are fastened\nwith strapping bands.\nDepending on orders, coils can be delivered fastened\nwith or without a pallet, protected with cardboard or\nlaminated plastic, and plastic or metallic end rings,\nmetallic corner protection and strapping bands.\n\n\nSteel bands or strappings, wood props, paper or plastic\nfilm, corner protection and other accessories supporting\npackaging are used as appropriate, depending on the\nprotection needed. Paper and plastic film are usually used\n\n\n3. Production and transportation\n\n\n3.1 PRODUCTION SITES\n\n\nannealed and tempered, and cut-to-length or slitted at\nSSAB Borl\u00e4nge (Sweden).\n\n\nBlast furnace-based steel slabs used for cold rolled steels\nare typically manufactured at SSAB Lule\u00e5 and at SSAB\nOxel\u00f6sund in Sweden. Slabs are produced using iron ore\npellets and, as an energy source and reducing agent, coke\nfrom coal, and injection carbon. These raw materials are\ncharged into a blast furnace to produce molten hot metal,\npig iron.\nSteel scrap and alloying elements are then added to the\nhot metal along with slag forming burnt lime, and oxygen\nis blown through the mixture to convert it into liquid steel\nin the basic oxygen furnace (BOF). The liquid crude steel\nis then cast into slabs on a continuous casting line.\n\n\nPickling, cold rolling, batch annealing and tempering,\nand cut-to-length or slitting can also be done at SSAB\nH\u00e4meenlinna (Finland) using SSAB hot-rolled coils\nproduced in Raahe (Finland).\nCo-products, such as slag, mill scale and iron oxide,\ngenerated in SSAB's steel production processes are\nrecycled as industrial raw materials or materials to\nreplace virgin resources. A high percentage of the\nbaghouse dust originating in various processes is\nrecycled to reduce waste and improve efficiency.\n\n\nBlast furnace-based steel slabs are hot rolled, pickled,\ncold rolled, continuous annealed and tempered or batch\n\n\nFIGURE 1. SSAB main production sites and process steps for blast furnace-based cold rolled steel sheets and slit coils.\nSSAB Cold Rolled Steel sheets and slit coils - main production processes\n\n\n| Slabs\n\n\nBOF Steel Making\n\n\nContinuous Casting\n\n\nRaw materials\n\n\nBlast Furnace\n\n\nSSAB LULEA-\n\n\n\u54c1\n\n\nSSAB OXEL\u00d6SUND\n\n\nScrap Lime Alloys\n\n\nIron ore\n\n\nLime\n\n\nCoke\n\n\n|Train\n\n\n\u5712\n\n\nHot rolling\n\n\nCold rolling\n\n\nCut-to-length or slitting\n\n\nPickling\n\n\nAnnealing and tempering\n\n\nSSAB BORL\u00c4NGE\n\n\nSSAB\n\n\nCold rolled steel\nsheets and slit\n\n\ncoils\n\n\nNote: pickling, cold rolling, batch annealing and tempering, and cut-to-length or slitting can also be done at SSAB H\u00e4meenlinna (Finland)\nusing SSAB hot-rolled coils produced in Raahe (Finland).\n\n\n3.2 TRANSPORTATION\n\n\nBorl\u00e4nge. Hot rolled coils from SSAB Raahe are\ntransported by rail to SSAB H\u00e4meenlinna.\n\n\nBlast furnace-based steel slabs from SSAB Lule\u00e5 and\nfrom SSAB Oxel\u00f6sund are transported by rail to SSAB\n\n\n5\n\n\n4. LCA\n\n\nTime representativeness:\n\n\n4.1 LCA INFORMATION\n\n\n2023 for the steel slabs production at SSAB Oxel\u00f6sund,\n2023 for the steel slabs production at SSAB Lule\u00e5,\n2021 for the steel slabs production at SSAB Raahe,\n2022 for the steel processing at SSAB Borl\u00e4nge,\n2021 for the steel processing at SSAB H\u00e4meenlinna.\n\n\nDeclared unit:\n1 kg of product\n\n\nReference service life:\n\n\nNot applicable\n\n\nDatabase(s) and LCA software used:\n\n\nDescription of system boundaries:\n\n\nThe LCA was modelled using the LCA software LCA for\nExperts and corresponding database (version 2024.1)\nprovided by Sphera.\n\n\nThe system boundaries are cradle-to-gate with modules\nC1-C4 and module D.\n\n\nSystem diagram:\n\n\nPre-consumer:\nwith burden (A1)\n\n\nEoL of Core waste (A3)\n\n\nSSAB Lule\u00e5,\nOxel\u00f6sund,\nRaahe:\nBlast furnace\nbased steel\nproduction\n(A3)\n\n\nPost-consumer:\nburden-free (A1)\n\n\nInternal scrap:\nburden-free (A1)\n\n\nProduction of\niron ore, energy and\nancilliary materials\nor chemicals (A1)\n\n\nEoL of Core waste (A3)\n\n\nDownstreams\nat SSAB's sites\nin Europe\n(Sweden and\nFinland) (A3)\n\n\nMaterial\n\n\nto module C\n\n\nWaste processing,\nsteel (C3)\n\n\nCredit,\n\n\nsteel (D)\n\n\nThe product\nis assumed to\nend up in steel\nrecycling\n\n\nDeconstruction\n\n\nTransport\n(C2)\n\n\n(C1)\n\n\nDisposal,\nsteel (C4)\n\n\nModule A1: Production of raw materials and production\nof fuels\n\n\nModule C1: Deconstruction of the product\n\n\n\u2022\n\n\n\u2022\n\n\n\u2022 Module C2: Transport to waste processing and disposal\nModule C3: Waste processing of the product, to be sent\nto steel recycling\n\n\nModule A2: Transportation of raw materials to SSAB's\nmanufacturing site (including transportation of steel\nbetween SSAB sites)\n\n\n\u2022\n\n\nModule C4: Disposal of the remaining part of the prod-\nuct in a landfill\n\n\n\u2022\n\n\nModule A3: Manufacturing of steel products and\nmanagement of production waste\n\n\n\u2022\n\n\nModule D: Benefits from recycling the steel\n\n\n\u2022\n\n\nSweden and for Finland has been applied (corresponding\nto a GWP-GHG impact of 0.07 kg CO2eq per kWh, and\n0.5 kg CO\u2082eq per kWh, respectively), however 30% from\nH\u00e4meenlinna electricity is fossil free with GWP-GHG\nimpact of 0.0046 kg CO\u2082eq per kWh.\n\n\nAllocation:\n\n\nPre-consumer scrap is used in the production of steel.\nThe environmental burden from the use of this scrap is\nallocated based on economic value by making a\nconservative assumption equal to 5% of virgin (blast\nfurnace-based) steel. This corresponds to a value of\n0.1 kg CO,eq per kg of pre-consumer scrap.\nCo-products from blast furnace and coke making\noperations have been allocated based on economic\nvalue as per PCR 2019:14. Similarly, impact associated\nwith internal energy generation have been allocated\nbased on economic value.\n\n\nScenario for module C1:\n\n\nThe product is being deconstructed by a machine pow-\nered by diesel.\n\n\nScenario for module C2:\n\n\nThe waste is transported 150 km by truck to waste pro-\ncessing (C3) and disposal (C4).\n\n\nCut-off criteria:\n\n\nScenario for module C3:\n\n\n98% of the product is assumed to be processed in order\nto be sent for recycling.\n\n\nThe maximum cut-off criteria established by the PCR and\nEN 15804:2012+A2:2019 standard is 1% of all material\nand energy flows to a single unit process and 5% of total\ninflows (mass and energy) to the upstream and core\nmodule. No cut-offs exceeding this limit have been made.\n\n\nScenario for module C4:\n\n\n2% of the product is assumed to be disposed of as waste\nat a landfill.\n\n\nInclusion of infrastructure and capital goods:\nInfrastructure and capital goods are not included in any\nof the modules covered in this EPD. For the electricity\nsources of renewable origin (within the residual mix),\nthe infrastructure of the power plant is included.\n\n\nScenario for module D:\n\n\nThe environmental benefit of the recycled steel is gained\nthrough the avoided production of primary steel. This\nbenefit corresponds to -1.7 kg CO2eq per kg of scrap in\nmodule D. The net flow of the recycled steel being credit-\ned in module D corresponds to 0.86 kg and is based on an\nassumed recycling rate of 98% and an assumption of yield\nlosses in the steel recycling process.\n\n\nElectricity information:\n\n\nAt SSAB Oxel\u00f6sund, SSAB Lule\u00e5, and SSAB Raahe\nsome of the electricity used is produced internally\n(corresponding to a GWP-GHG impact of 0.62, 0.36 and\n2.03 kg CO\u2082eq per kWh, respectively). At SSAB Borl\u00e4nge\nand SSAB H\u00e4meenlinna, only external electricity is used.\nFor external electricity, the residual electricity mix for\n\n\nWeighted average for the EPD:\n\n\nThe results represent a weighted average based on the\nproduction volumes for the product group.\n\n\nModules declared, geographical scope, share of specific data (in GWP-GHG indicator) and data variation\n\n\nLife cycle stage\n\n\nSpecific\ndata used\n\n\nVariation -\nproducts\n\n\nVariation -\nsites\n\n\nModule\n\n\nModules\ndeclared\n\n\nGeography\n\n\nEU, SE & FI\n\n\nRaw material supply\nTransport\n\n\nA1\n\n\nX\n\n\n+ 8%\n- 1%\n\n\nProduct stage\n\n\nEU, SE & FI\n\n\n<10%\n\n\nA2\n\n\n76%\n\n\nX\n\n\nSE & FI\n\n\nManufacturing\n\n\nA3\n\n\nX\n\n\nTransport\n\n\nA4\n\n\nND\n\n\nConstruction\nprocess stage\n\n\nConstruction installation\n\n\nA5\n\n\nND\n\n\nUse\n\n\nB1\n\n\nND\n\n\nB2\n\n\nMaintenance\n\n\nND\n\n\nRepair\n\n\nB3\n\n\nND\n\n\nUse stage\n\n\nReplacement\n\n\nB4\n\n\nND\n\n\nRefurbishment\n\n\nB5\n\n\nND\n\n\nOperational energy use\n\n\nB6\n\n\nND\n\n\nOperational water use\n\n\nB7\n\n\nND\n\n\nDe-construction demolition\n\n\nC1\n\n\nX\n\n\nEU\n\n\nTransport\n\n\nC2\n\n\nX\n\n\nEU\n\n\nEnd of life stage\n\n\nWaste processing\n\n\nC3\n\n\nX\n\n\nEU\n\n\nDisposal\n\n\n\u0f59\u0f74\u3002\nC4\n\n\nX\n\n\nEU\n\n\nResource recovery stage Reuse-Recovery-Recycling-potential D\n\n\nX\n\n\nEU\n\n\nX: Module Declared\n\n\nND: Module not declared\n\n\n7\n\n\n4.2 PRODUCT CONTENT DECLARATION\n\n\nPre- and postconsumer scrap content is 4.1%.\nRecycled material content with internal scrap is 18.9%.\n\n\nContent declaration and average chemical composition of\ncold rolled steel sheets and slit coils per kg produced is:\n\n\nProduct Composition\n\n\nWeight (%)\n\n\nBiogenic carbon,\nweight (%)\n\n\nWeight (kg)\n\n\nBiogenic carbon,\nweight (kg)\n\n\nPre-consumer scrap\n\n\n0.021\n\n\n0%\n\n\n0\n\n\n2.1%\n\n\nPost-consumer scrap\n\n\n0.020\n\n\n2.0%\n\n\n0%\n\n\n0\n\n\nInternal scrap\n\n\n0.148\n\n\n0%\n\n\n0\n\n\n14.8%\n\n\nPrimary steel\n\n\n81.2%\n\n\n0.812\n\n\n0%\n\n\n0\n\n\nAverage chemical composition*\n\n\nIron (Fe)\n\n\n> 97%\n\n\nManganese (Mn)\n\n\n0.6%\n\n\nSilicon (Si)\n\n\n0.3%\n\n\nCarbon (C)\n\n\n0.1%\n\n\n< 1.5%\n\n\nOther\n\n\n* The figures provided represent the best estimate at the time of publication.\n\n\nContent Declaration of renewable\npackaging material\n\n\nWeight (kg)\n\n\nWeight %\n(of product)\n\n\nBiogenic carbon, weight\n(kg/declared unit)\n\n\n0.0030\n\n\nWood\n\n\n0.30%\n\n\n0.0013\n\n\nThe production of the packaging materials has been\nomitted since it falls under the cut-off limit. The content\nof biogenic material in the packaging is 0.0013 kg per kg\nof steel.\n\n\n4.3 ENVIRONMENTAL PERFORMANCE\n\n\nINDICATOR RESULTS\n\n\nThe estimated impact results are only relative statements,\nwhich do not indicate the endpoints of the impact\ncategories, exceeding threshold values, safety margins\nand/or risks. Usage of results from A1-A3 without\n\n\nCold rolled steel sheets and slit coils do not contain\nany of the substances of very high concern (SVHC)\nregulated by Regulation (EC) No 1907/2006 (REACH)\nor Regulation (EC) No 1272/2008 of the European\nParliament and of the Council.\n\n\nconsidering the results of module C is not encouraged.\n\n\n80\n\n\nPotential environmental impact - mandatory indicators according to EN 15804+A2 (version EF 3.1)\n\n\nResults per declared unit: 1 kg of product\n\n\nIndicator\n\n\nA1-A3\n\n\nUnit\n\n\nC2\n\n\nD\n\n\nC1\n\n\nC3\n\n\nC4\n\n\nClimate Change - fossil\n\n\nkg CO\u2082 eq\n\n\n2.17\n\n\n4.39E-04\n\n\n1.01E-02 2.71E-03 2.99E-04\n\n\n-1.49\n\n\n8.16E-04 1.34E-06 2.68E-05 1.01E-05 9.52E-07 3.16E-04\n\n\nClimate Change - biogenic\n\n\nkg CO\u2082 eq\n\n\nGlobal warming\npotential (GWP)\n\n\nClimate Change - land use and land\n\n\n7.31E-06 8.61E-05 3.66E-05 1.80E-06 -1.98E-04\n\n\nkg CO\u2082 eq\n\n\n4.81E-04\n\n\nuse change (LULUC)\n\n\nClimate Change - total\n\n\n2.17\n\n\nkg CO\u2082 eq\n\n\n4.47E-04\n\n\n1.03E-02 2.76E-03 3.02E-04\n\n\n-1.49\n\n\nDepletion potential of the stratospheric ozone layer (ODP) kg CFC-11 eq\n\n\n1.34E-18 4.89E-15 8.08E-16 2.00E-12\n\n\n6.00E-12\n\n\n4.39E-17\n\n\nAcidification potential (AP)\n\n\nmole H+ eq\n\n\n1.18E-05 1.36E-05 2.13E-06 -3.64E-03\n-3.47E-07\n1.35E-03 1.49E-06 3.93E-06 6.24E-06 5.47E-07 -5.85E-04\n\n\n4.76E-03\n\n\n3.01E-06\n\n\nkg P eq\n\n\nFreshwater\n\n\n7.71E-07\n\n\n1.86E-09\n\n\n3.12E-08 1.05E-08 6.80E-10\n\n\nEutrophication\npotential (EP)\n\n\nMarine\n\n\nkg N eq\n\n\nTerestrial\n\n\nmole N eq\n\n\n-5.24E-03\n1.67E-06 -2.38E-03\n\n\n1.43E-02 1.65E-05\n\n\n4.74E-05\n\n\n6.90E-05 6.03E-06\n\n\nFormation potential of tropospheric ozone (POCP)\n\n\nkg NMVOC eq\n\n\n2.89E-06 1.01E-05\n\n\n3.78E-03\n\n\n1.73E-05\n2.84E-09\n5.07E-02 3.95E-03\n\n\nkg Sb eq\n\n\n3.71E-11 8.05E-10\n\n\nAbiotic\n\n\nMinerals and metals*\n\n\n7.64E-06\n\n\n1.94E-11 -8.43E-06\n\n\ndepletion\n\n\nFossil resources*\n\n\n28.7\n\n\n0.140\n\n\nMJ\n\n\n5.68E-03\n\n\n-14.8\n\n\npotential (ADP)\n\n\nWater scarcity potential (WDP)*\n\n\nm\u00b3\n\n\n5.18E-04\n\n\n-0.100\n\n\n6.48E-06 9.13E-05\n\n\n0.179\n\n\n3.43E-05\n\n\n*Disclaimer: The results of this environmental impact indicator shall be used with care as the uncertainties of these results are high or as there is limited\nexperience with the indicator.\n\n\nNote: Biogenic carbon in packaging is balanced in A1-A3.\n\n\nAdditional mandatory and voluntary impact category indicators\n\n\nResults per declared unit: 1 kg of product\n\n\n72\nC2\n\n\nIndicator\n\n\nUnit\n\n\nA1 - A3\n\n\nC4\n\n\nD\n\n\nC1\n\n\nC3\n\n\nGlobal warming potential (GWP) GWP-GHG (1)\n\n\nkg CO\u2082 eq\n\n\n2.17\n\n\n4.47E-04\n\n\n1.03E-02\n\n\n2.76E-03\n\n\n-1.49\n\n\n3.02E-04\n\n\n(1) This indicator accounts for all greenhouse gases except biogenic carbon dioxide uptake and emissions and biogenic carbon stored in the product.\nAs such, the indicator is identical to GWP-total except that the characterization factor for biogenic CO2 is set to zero.\n\n\nResource use indicators\n\n\nResults per declared unit: 1 kg of product\n\n\nIndicator\n\n\nA1-A3\n\n\nUnit\n\n\nC2\n\n\nD\n\n\nC1\n\n\nC3\n\n\nC4\n\n\n4.80E-04 7.81E-03 5.41E-03 6.89E-04\n0.00E+00 0.00E+00\n4.80E-04 7.81E-03\n0.140\n\n\nUsed as energy carrier (PERE)\n\n\n1.28\n\n\n0.584\n\n\nMJ\n\n\nPrimary energy\n\n\nUsed as raw materials (PERM)\n\n\n0.00E+00\n\n\n0.00E+00 0.00E+00 0.00E+00\n\n\nMJ\n\n\nresources -\nRenewable\n\n\nTotal (PERT)\n\n\n5.41E-03 6.89E-04\n\n\n0.584\n\n\nMJ\n\n\n1.28\n\n\nUsed as energy carrier (PENRE)\n\n\n5.07E-02 3.95E-03\n0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n5.07E-02 3.95E-03\n\n\n-14.8\n\n\n28.7\n\n\n5.68E-03\n\n\nMJ\n\n\nPrimary energy\n\n\nUsed as raw materials (PENRM)\n\n\n0.00E+00\n\n\nMJ\n\n\nresources -\n\n\nNon-renewable\n\n\nTotal (PENRT)\n\n\n28.7\n\n\n5.68E-03\n\n\nMJ\n\n\n0.140\n\n\n-14.8\n\n\nUse of secondary material (SM)\n\n\n4.40E-02 0.00E+00 0.00E+00\n\n\n0.00E+00 0.00E+00 0.00E+00\n0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n4.18E-25 0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n8.18E-03 5.39E-07 8.94E-06 1.51E-05 1.05E-06 -0.151\n\n\nkg\n\n\nUse of renewable secondary fuels (RSF)\n\n\n3.56E-26\n\n\nMJ\n\n\nUse of non renewable secondary fuels (NRSF)\n\n\nMJ\n\n\nm\u00b3\n\n\nNet use of fresh water (FW)\n\n\nNote: Primary energy calculated using PCR option B.\n\n\n9\n\n\nWaste indicators\n\n\nResults per declared unit: 1 kg of product\n\n\nUnit\n\n\nA1-A3\n\n\nIndicator\n\n\nD\n\n\nC1\n\n\nC2\n\n\nC3\n\n\nC4\n\n\nHazardous waste disposed (HWD)\n\n\n1.19E-06 1.84E-13 7.06E-12\n8.84E-07 2.08E-05 1.39E-05 2.00E-02\n7.34E-09 1.70E-07 6.38E-07 4.15E-08 1.62E-06\n\n\n7.33E-12 9.84E-13 -1.11E-07\n\n\nkg\n\n\nNon-hazardous waste disposed (NHWD)\n\n\nkg\n\n\n0.179\n\n\n3.73E-02\n\n\nRadioactive waste disposed (RWD)\n\n\nkg\n\n\n9.85E-04\n\n\nOutput indicators\n\n\nResults per declared unit: 1 kg of product\n\n\nIndicator\n\n\nUnit\n\n\nA1 - A3\n\n\nD\n\n\nC1\n\n\nC2\n\n\nC3\n\n\nC4\n\n\n0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n0.00E+00 0.00E+00 0.00E+00 0.980 0.00E+00 0.00E+00\n0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00 0.00E+00\n\n\nComponents for re-use (CRU)\n\n\nkg\n\n\nMaterials for recycling (MFR)\n\n\nkg\n\n\nMaterial for energy recovery (MER)\n\n\nkg\n\n\nExported electrical energy (EEE)\n\n\nMJ\n\n\nExported thermal energy (EET)\n\n\nMJ\n\n\nDisclaimer\n\n\nILCD classification\n\n\nDisclaimer\n\n\nIndicator\n\n\nGlobal warming potential (GWP)\n\n\nNone\n\n\nDepletion potential of the stratospheric ozone layer (ODP)\n\n\nILCD Type 1\n\n\nNone\n\n\nPotential incidence of disease due to PM emissions (PM)\n\n\nNone\n\n\nAcidification potential, Accumulated Exceedance (AP)\n\n\nNone\n\n\nEutrophication potential, Fraction of nutrients reaching freshwater\nend compartment (EP-freshwater)\n\n\nNone\n\n\nEutrophication potential, Fraction of nutrients reaching marine\nend compartment (EP-marine)\n\n\nNone\n\n\nILCD Type 2\n\n\nEutrophication potential, Accumulated Exceedance (EP-terrestrial)\n\n\nNone\n\n\nFormation potential of tropospheric ozone (POCP)\n\n\nNone\n\n\nPotential Human exposure efficiency relative to U235 (IRP)\n\n\n1\n\n\nAbiotic depletion potential for non-fossil resources (ADP-minerals&metals)\n\n\n2\n\n\nAbiotic depletion potential for fossil resources (ADP-fossil)\nWater (user) deprivation potential, deprivation-weighted\nwater consumption (WDP)\n\n\n2\n\n\n2\n\n\nILCD Type 3\n\n\nPotential Comparative Toxic Unit for ecosystems (ETP-fw)\nPotential Comparative Toxic Unit for humans (HTP-C)\nPotential Comparative Toxic Unit for humans (HTP-nc)\nPotential Soil quality index (SQP)\n\n\n2\n\n\n2\n\n\n2\n\n\n2\n\n\nDisclaimer 1 - This impact category deals mainly with the eventual impact of low dose ionizing radiation on human health of the nuclear fuel\ncycle. It does not consider effects due to possible nuclear accidents, occupational exposure nor due to radioactive waste disposal in underground\nfacilities. Potential ionizing radiation from the soil, from radon and from some construction materials is also not measured by this indicator.\nDisclaimer 2 - The results of this environmental impact indicator shall be used with care as the uncertainties on these results are high or as there\nis limited experienced with the indicator.\n\n\n10\n\n\nVariation in environmental indicators\n\n\nThe table below shows the variation for modules A-C where the difference between products is greater than 10%.\n\n\nCold rolled steel sheets and slit coils\n\n\nDifference (%)\n\n\nEnvironmental impact indicator\n\n\nGWP-biogenic\n\n\n21%\n\n\n43%\n\n\nODP\n\n\nAP\n\n\n16%\n\n\nEP-fresh\n\n\n36%\n\n\nEP-marine\n\n\n14%\n\n\nEP-terrest\n\n\n14%\n\n\nPOCP\n\n\n14%\n\n\nADP-elements\n\n\n41%\n\n\nADP-fossil\n\n\n16%\n\n\n5. References\n\n\nGeneral Programme Instructions of the International EPD\u00ae System. Version 4.0\n\n\n\u2022\n\n\nPCR 2019:14 Construction products. Version 1.3.4 (2024-04-30)\n\n\n\u2022\n\n\nCEN European Committee for Standardisation (2021). EN15804:2012+A2:2019/\nAC:2021 (CEN 2021), Sustainability of construction works - Environmental product\ndeclarations - Core rules for the product category of construction products\n\n\n\u2022\n\n\nLCA for experts Software System and database for Life Cycle Engineering, sphera,\nLeinfelden-Echterdingen, Germany\n\n\n\u2022\n\n\n\u26ab Hallberg, L., LCA methodology report - SSAB Blast Furnace steel, as basis for the\npublication of EPDs within the International EPD\u24c7 System, June 2025\n\n\n11\n\n\nSSAB\n\n\nwww.ssab.com\n"}, "expected_output": {"claims": [{"unit": "%", "value": 2, "evidence": ["Post-consumer scrap", "2.0%"]}, {"unit": "%", "value": 2.1, "evidence": ["Pre-consumer scrap", "2.1%"]}, {"unit": "%", "value": 4.1, "evidence": ["Pre- and postconsumer scrap content is 4.1%.", "The system boundaries are cradle-to-gate with modules\nC1-C4 and module D.\n"]}, {"unit": "%", "value": 18.9, "evidence": ["Recycled material content with internal scrap is 18.9%.", "The system boundaries are cradle-to-gate with modules\nC1-C4 and module D.\n"]}]}, "metadata": {"product_category": "Metal, mineral, plastic & glass products", "request_id": "req_792a548b616ebed2"}}
{"id": "b2234323ef5e7afcafdc3c15", "input": {"query": "What nitrogen fertilizer application rates in kg N per hectare were used for zero-tillage barley production in Queensland dryland conditions?", "source_url": "https://era.dpi.qld.gov.au/id/eprint/1355/1/ThomasZeroTillage-sec.pdf", "document_text": "CSIRO PUBLISHING\n\n\nwww.publish.csiro.au/journals/ajea\n\n\nAustralian Journal of Experimental Agriculture, 2007, 47, 965\u2013975\n\n\nZero tillage and nitrogen fertiliser application in wheat\nand barley on a Vertosol in a marginal cropping area\nof south-west Queensland\n\n\nG. A. Thomas,F, R. C. Dalal\u00b3, E. J. Weston C, D, C. J. Holmes^,E, A. J. King^,\nD. N. Orange and K. J. Lehane\n\n\nAQueensland Department of Natural Resources and Water, Natural Resource Sciences,\nSouth West Region, PO Box 318, Toowoomba, Qld 4350, Australia.\nBQueensland Department of Natural Resources and Water, Natural Resource Sciences,\n\n\n80 Meiers Road, Indooroopilly, Brisbane, Qld 4068, Australia.\n\n\nCQueensland Department of Primary Industries and Fisheries, PO Box 102, Toowoomba,\n\n\nQld 4350, Australia.\n\n\nDPresent address: 50 Broadwater Terrace, Redland Bay, Qld 4165, Australia.\n\n\nEPresent address: MS 172, Allora, Qld 4362, Australia.\n\n\nFCorresponding author. Email: gregory.thomas@nrw.qld.gov.au\n\n\nAbstract. Winter cereal cropping is marginal in south-west Queensland because of low and variable rainfall and\ndeclining soil fertility. Increasing the soil water storage and the efficiency of water and nitrogen (N) use is essential for\nsustainable cereal production. The effect of zero tillage and N fertiliser application on these factors was evaluated in wheat\nand barley from 1996 to 2001 on a grey Vertosol. Annual rainfall was above average in 1996, 1997, 1998 and 1999 and\nbelow average in 2000 and 2001. Due to drought, no crop was grown in the 2000 winter cropping season.\nZero tillage improved fallow soil water storage by a mean value of 20 mm over 4 years, compared with conventional\ntillage. However, mean grain yield and gross margin of wheat were similar under conventional and zero tillage. Wheat\ngrain yield and/or grain protein increased with N fertiliser application in all years, resulting in an increase in mean gross\nmargin over 5 years from $86/ha, with no N fertiliser applied, to $250/ha, with N applied to target \u226513% grain protein.\nA similar increase in gross margin occurred in barley where N fertiliser was applied to target malting grade. The highest\nN fertiliser application rate in wheat resulted in a residual benefit to soil N supply for the following crop.\nThis study has shown that profitable responses to N fertiliser addition in wheat and barley can be obtained on long-term\ncultivated Vertosols in south-west Queensland when soil water reserves at sowing are at least 60% of plant available water\ncapacity, or rainfall during the growing season is above average. An integrative benchmark for improved N fertiliser\nmanagement appears to be the gross margin/water use of ~$1/ha.mm. Greater fallow soil water storage or crop water use\nefficiency under zero tillage has the potential to improve winter cereal production in drier growing seasons than\nexperienced during the period of this study.\n\n\nIntroduction\n\n\nWinter cereal cropping, with wheat as the main rainfed crop,\nis carried out over 200000 ha in south-west Queensland\n(Australian Bureau of Statistics 2001). Winter cereals are\nusually sown between mid April and late June, depending on\nsowing opportunity, following a 6-8 month fallow period after\nharvest of the preceding crop in October-November. Reduced\ntillage practices during the fallow period, including zero tillage,\nhave been shown to result in better soil water storage and grain\nyield of wheat and barley than conventional tillage in other parts\nof southern Queensland (Thomas et al. 1997). These practices\ntherefore offer potential to improve the reliability of cropping in\nthe south-west area. Nitrogen fertiliser application is an option\nfor compensating for N fertility decline under winter cereal\ncropping. However, when the experiments discussed in this\npaper commenced in 1996, N fertiliser use in the south-west\n\n\nThe south-west Queensland cropping region, which includes the\nRoma, St George and Dirranbandi districts, is marginal for\nrainfed crop production. The region receives an average annual\nrainfall of 500-600 mm with 20-30% year-to-year variability\nand has annual potential evapotranspiration of ~2000 mm,\nresulting in a rainfall: potential evapotranspiration ratio of <0.3\n(Weston et al. 1975). Moreover, as the period of cereal cropping\nhas increased, grain yield and grain protein have decreased as\nsoil fertility, particularly soil nitrogen (N) supply, has declined\n(Dalal and Mayer 1986a). On a Vertosol, which is the main\ncropping soil in the region, the mean rate of net N loss from the\nsoil profile was 35.8 kg N/ha.year (Dalal and Mayer 1986b).\nTherefore, it is essential to address these issues of water and N\nsupply to crops in developing sustainable farming systems for\nthis area if it is to continue to support grain cropping.\n\n\n\u00a9 CSIRO 2007\n\n\n10.1071/EA05253\n\n\n0816-1089/07/080965\n\n\nAustralian Journal of Experimental Agriculture\n\n\n966\n\n\nG. A. Thomas et al.\n\n\nTable 1.\n\n\nSoil chemical and physical characteristics, Nindigully trial\n\n\nSoil depth\n(cm)\n\n\nElectrical\nconductivity\n(1:5 H2O) (dS/m)\n\n\nCation exchange Exchangeable\n\n\nClay (%)\n\n\nChloride\n(mg/kg)\n\n\npH\n(1:5 H\u2082O)\n\n\nPlant available\n\n\nsodium (%)\n\n\ncapacity\n[cmol(+)/kg]\n\n\nwater capacity\n(mm)\n\n\n22222\n25\n26\n24\n20\n\n\n0.10\n\n\n0-10\n\n\n3.8\n\n\n24\n\n\n8.6\n\n\n10\n\n\n52\n\n\n11\n\n\n10-30\n\n\n0.11\n\n\n53\n\n\n9.0\n\n\n6.2\n\n\n40\n\n\n0.18\n\n\n11.7\n\n\n30-60\n\n\n36\n\n\n55\n\n\n60\n\n\n9.2\n\n\n60-90\n\n\n8.8\n\n\n64\n\n\n0.59\n\n\n144\n\n\n15.5\n\n\n54\n\n\n90-120\n\n\n1.58\n\n\n305\n\n\n57\n\n\n8.1\n\n\n17.0\n\n\n54\n\n\narea was none or low and little was known about the economics\nof N fertiliser application in the area.\nThe aim of this study was to determine whether zero tillage\nfallow management and N fertiliser application improve grain\nyield and quality and gross margin in wheat and barley on a\nVertosol in the marginal cropping region of south-west\nQueensland.\n\n\nand physical characteristics of the soil profile are given in\nTable 1. Details of the analytical methods used for these\nanalyses are given by Bruce and Rayment (1982).\nAt St George, long-term (1943\u201396) mean daily maximum\nand minimum temperatures range from 19\u00b0C and 5.4\u00b0C,\nrespectively, in July, to 34.5\u00b0C and 21.5\u00b0C, respectively, in\nJanuary (Bureau of Meteorology). Frosts (mean screen\ntemperature \u22642.2\u00b0C) may occur between early May and mid\nSeptember. Mean daily evaporation ranges from 3.1 mm in June\nto 11.0 mm in December. Long-term mean monthly rainfall at\nSt George and actual monthly rainfall at the Nindigully trial site\nfrom 1996 to 2001 are shown in Table 2.\n\n\nMaterials and methods\n\n\nField trial site\n\n\nWe established a farming systems trial at Nindigully, on the\nproperty 'Dunkerry South' (28\u00b030'S, 148\u00b045\u2032E), ~50 km south\nof St George in south-west Queensland, in May 1996. The site\nI was on a grey Vertosol soil (Isbell 1996) that had been under\ncultivation since 1956. Native vegetation at the site was\nprimarily coolibah (Eucalyptus microtheca) open woodland and\nMitchell grass (Astrebla lappacea). The site was chosen as\nbeing representative of a land system in the region with a large\narea of cropping (~100000 ha, Robinson 2005) and in which\nsoil fertility decline had been identified (Dalal and Mayer\n1986a, 1986b). At the start of the trial, soil organic carbon, total\nnitrogen and bicarbonate-extractable phosphorus at 0-10 cm\nwere 0.65%, 0.07% and 9 mg/kg, respectively. Other chemical\n\n\nField trial establishment and crop management\n\n\nThe field trial commenced in May 1996 following a wheat crop\nin 1995 and three tillage operations with a chisel plough during\nthe fallow period. Several experiments were conducted within\nthe trial area. In this paper, we report on studies on tillage\nmethod (conventional and zero) (experiment 1) and N fertiliser\napplication (experiment 2). The experiments were sown and\nharvested in 1996, 1997, 1998, 1999 and 2001. Due to drought,\nno crop was grown in the 2000 winter cropping season.\nIn experiment 1, following harvest of wheat in 1996, main\nplots, 40 m long and 36 m wide and replicated three times, were\n\n\nTable 2. Long-term mean monthly, in-crop and fallow rainfall for St George (1881\u20131996), and monthly,\nin-crop and fallow rainfall recorded at the Nindigully trial site from 1995 to 2001\nn.m., not measured\n\n\nMean rainfall\n\n\nRainfall (mm) at Nindigully trial site\n1998\n\n\nMonth\n\n\nfor St George,\n1881-1996 (mm)\n\n\n1996\n\n\n1995\n\n\n1997\n\n\n1999 2000 2001\n\n\n194\n573-376-346\u00d722\n95\n44\n1\n1\n18\n\n\n2020*232265\n24\n64\n27\n51\n\n\n35\n\n\nJan.\n\n\n74\n\n\n238\n\n\n118\n\n\n45\n\n\nFeb.\n\n\n62\n\n\n50\n\n\n61\n\n\n22\n\n\n13\n\n\n0\n\n\nMar.\n\n\n54\n\n\n13\n\n\n189\n\n\n61\n\n\n33\n\n\n63\n\n\n22\n\n\nApr.\nMay\nJune\n\n\n24\n\n\n4\n\n\n39\n\n\n87\n\n\n72\n\n\n11\n\n\n18\n\n\n33\n\n\n33\n\n\n9\n\n\n66\n\n\n60\n\n\n33\n\n\nJuly\nAug.\nSept.\n\n\n43\n\n\n124\n\n\n31\n\n\n2\n\n\n25\n\n\n9\n\n\n54\n\n\n108\n\n\n41\n\n\n27\n\n\n23\n\n\n78\n\n\n32\n\n\n0\n\n\n48\n\n\n94\n\n\n73\n\n\nOct.\n\n\n39\n\n\n48\n\n\n55\n\n\n57\n\n\n32\n\n\n93\n\n\nNov.\n\n\n46\n\n\n68\n\n\n52\n\n\n32\n\n\n3\n\n\nDec.\n\n\n62\n\n\n203\n\n\n58\n\n\n8\n\n\nTotal\n\n\n730\n\n\n786\n\n\n372\n\n\n516\n\n\n700\n\n\n616\n\n\nn.m.\n\n\n269 111A\n286\n\n\nMay Oct. (usual in-crop)\nNov.-Apr. (usual fallow)\n\n\n196\n\n\n288\n\n\n181\n\n\n496\n\n\nn.m.\n\n\n320\n\n\n339\n\n\n274\n\n\n410\n\n\n384\n\n\n426\n\n\nANo crop sown due to lack of sowing rainfall.\n\n\nAustralian Journal of Experimental Agriculture 967\n\n\nTillage and N supply in cereals on a Vertosol in south-west Queensland\n\n\neach split into two treatments, conventional tillage (CT) and\nzero tillage (ZT). The main plots comprised 16 runs of a 2.25-m\nwide seeder. The tillage subplot width was 18 m (eight seeder-\nruns). Tillage treatments were imposed after the harvest of 1996\ncereal crops and maintained until the harvest of 2001. The CT\ntreatment had 3-4 cultivations with a chisel plough or scarifier\nfor weed control during the fallow period. A similar number of\nherbicide applications were used for weed control during the\nfallow period under the ZT tillage system. At sowing of wheat\nin 1997, 1998, 1999 and 2001, the CT and ZT treatments were\nsplit further into two sub-subplots, each 9 m wide (four seeder-\nruns), one without and one with N fertiliser application.\nIn experiment 2, four N fertiliser treatments (one without and\nthree with N fertiliser application) were applied at sowing in\nwheat under the ZT system in all years. The main plots were 40 m\nlong and 36 m wide (16 runs of a 2.25-m wide seeder) with three\nreplications and were split into four N fertiliser subplots that were\neach 9 m wide (four seeder-runs). From 1996 to 1999, N fertiliser\ntreatments in wheat were applied to the same subplots each year.\nTwo N fertiliser treatments (one without and one with N\nfertiliser application) were applied at sowing in barley under ZT\nin all years. The main plots of barley were 40 m long and 18 m\nwide (eight seeder-runs) with three replications and were split\ninto two N fertiliser subplots that were each 9 m wide (four\nseeder-runs). The location of the main plots of barley varied\nfrom year to year within the trial area.\nIn experiments 1 and 2, a basal application of 40 kg/ha of\nmixed fertiliser (9.4% N, 20.5% P, 2.2% S, 2.5% Zn) was\nbanded with the seed at sowing in all years except 1998.\nIn both experiments, N fertiliser application rates were\ndetermined from consideration of soil water and soil nitrate-N\ncontent obtained from soil sampling before sowing. Using the\nequation of Dalal et al. (1997), relating grain protein\nconcentration to the ratio between available soil water (mm) and\nN supply (soil nitrate-N + N fertiliser) (kg/ha) at 0-120 cm\ndepth at sowing, the amount of N fertiliser required to target a\ngiven grain protein concentration was calculated. The required\nratios were 2.5 for malting barley (10.2 \u00b1 1% protein on oven-\ndry basis) and for wheat grain at 12% moisture, 2 for 10.5 \u00b1 1%\nprotein, 1.5 for 11.5 \u00b1 1% protein and \u22641 for \u226513% protein.\nWheat protein concentrations of 10.5%, 11.5% and 13% are the\nminimum requirements for the Australian Premium White\n(APW), Australian Hard (AH) and Australian Prime Hard\n(APH) grades, respectively. A retrospective assessment of N\nfertiliser requirements to achieve grain protein concentration of\n10.2% in barley and 13% in wheat was also carried out using the\ncomputer program HOWWET? (Freebairn et al. 1997). In the N\nrequirement section of the program, measured available soil\nwater and soil nitrate-N at sowing from experiment 2 were\nentered with both a commonly used estimate for water use\nefficiency (WUE) (10 kg/ha.mm, 100-mm threshold soil water)\nand a mean measured value for WUE (9.5 kg/ha.mm) from the\nexperiment from 1996 to 2001. The N fertiliser requirements\nestimated from HOWWET? were then compared with those\nusing the available soil water/N supply method.\nIn experiment 1, the N fertiliser application rate varied from\nyear to year, depending on soil water and nitrate-N content, and\nwas chosen to target wheat grain protein of 11.5 \u00b1 1% at 12%\nmoisture content in 1997 and 13 \u00b1 1% at 12% moisture content\n\n\nin 1998, 1999 and 2001. In experiment 2, application rates of N\nfertiliser also varied from year to year, depending on soil water\nand nitrate-N content (Tables 5 and 6). In each year, increasing\nN fertiliser rates were chosen to target grain protein of 10.5%,\n11.5% and 13% (\u00b11%) at 12% moisture in wheat, and a single\nrate was chosen to target malting barley of 10.2 \u00b1 1% grain\nprotein on an oven-dried basis.\nIn 1996, 1999 and 2001, wheat under CT and ZT and barley\nunder ZT were sown with a small-plot seeder which had nine\nrigid, spear point tines followed by solid, centre-ribbed press\nwheels, at a row spacing of 25 cm. Sowing depth was 5-8 cm. In\nN fertiliser treatments, urea was applied at sowing at 5-8 cm\ndepth in the centre of alternate seed rows. In 1997 and 1998,\nwheat and barley were sown with a small-plot air seeder of similar\nconfiguration, but with sowing tines modified to distribute N\nfertiliser in bands 3-4 cm on either side of the seed row.\n\n\nWheat (cv. Hartog) and barley (cv. Tallon) were both sown\non 24 May 1996 and on 29 May 1997. In 1998, wheat\n(cv. Sunco) was sown on 23 May and barley (cv. Tallon) was\nsown on 14 May. Wheat (cv. Kennedy) and barley (cv. Tallon)\nwere both sown on 22 June 1999 and 21 June 2001. Both crops\nwere sown at 40 kg/ha in 1996, 1999 and 2001 and at 30 kg/ha\nin 1997 and 1998.\n\n\nSoil and plant sampling and analysis\n\n\nSoil profile nitrate-N concentration and water content were\nmeasured from soil samples collected in April, 4-8 weeks\nbefore sowing and in October-November, shortly after harvest,\neach year. At each sampling time, four soil cores were taken in\na line across two adjacent crop rows and inter-rows at a\nrandomly selected location in each plot. Two samples were taken\nin the crop rows to a depth of 120 cm in 1996 and 150 cm in\nsubsequent years, and two samples were taken in the centre of\nthe inter-row space to a depth of 30 cm. Cores were 50 mm in\ndiameter in 1996 and 1997 and 38 mm in following years. Cores\nwere divided into 0-10, 10-30, 30-60, 60-90, 90-120 and\n120\u2013150 cm layers, according to depth of sampling. For each\nplot, the samples from corresponding layers were bulked, sealed\nin plastic bags and stored at 4\u00b0C until analysed for nitrate-N\nconcentration and moisture content.\nSoil was dried at 35 \u00b1 5\u00b0C in a forced draught oven and\nground <2 mm for colorimetric determination of nitrate-N (Best\n1976) after extraction of 10 g of soil in 100 mL of 2 mol/L KCl.\nNitrate-N in kg/ha was calculated from the nitrate-N\nconcentration for each layer, depth of the layer and its bulk\ndensity, and then summed for the whole profile.\nGravimetric soil water content (g/g) was determined by\ndrying soil samples at 105\u00b0C for 48 h. Volumetric soil water\ncontent (mm) for each layer was then calculated using bulk\ndensity values determined from soil cores taken when the soil\nprofile was fully wet. Plant available soil water content was\ndetermined for each layer by subtracting the lowest soil water\ncontent recorded during the trial for that layer in a mature,\nstressed crop (lower limit of water extraction) (Gardner 1985)\nfrom the measured soil water content, and was then expressed as\nequivalent depth of water (mm) for the whole profile.\nThe number of ears/m\u00b2 was determined from one or two\n0.25-m\u00b2 samples taken from each subplot at crop maturity. Crop\ngrain yield at maturity was determined from machine harvesting\n\n\n968 Australian Journal of Experimental Agriculture\n\n\nG. A. Thomas et al.\n\n\nmethod of Reynolds and Elrick (1991) to measure hydraulic\nconductivity and pore size distribution. Rainwater (electrical\nconductivity <3 \u00b5S/m) was also used for these measurements.\nFine bedding sand (mean diameter <0.001 m) was used to\nensure good contact between the disc permeameter and the soil.\nThe disc permeameters did not measure flow in pores >3 mm,\nso large cracks and micropores were not characterised.\nMeasurements were made from four disc permeameters per\ntreatment in each replicate.\nUsing a linear relationship between infiltration and surface\ncover (Freebairn et al.1993), infiltration measurements for 0%\nand 100% cover were used to estimate values for existing cover\nlevels under CT and ZT.\n\n\n1.75 m x 36 m of each of the inner two seeder-runs of each\nsubplot. Grain samples were taken from each subplot for grain\nN and grain weight determination. Grain N concentration was\ndetermined from Kjeldahl digests using automated ammonium\nanalysis (Crooke and Simpson 1971). Grain yield, grain protein\nconcentration and grain weight were calculated at 12% moisture\ncontent, except for barley grain protein concentration, which\nwas expressed at 0% moisture content.\n\n\nWater use efficiency and nitrogen use efficiency\n\n\nWater use efficiency for grain yield (kg/ha.mm) was determined\nby dividing grain yield by crop water use (mm). Crop water use\nwas calculated as profile soil water content in April before\nsowing, minus soil water content in October-November after\nharvest, plus in-crop rainfall received between these sampling\ntimes, mostly during crop growth.\nNitrogen use efficiency (NUE) for grain N yield was\ncalculated as follows: NUE (%)\nNUE (%) \n= 100 \u00d7 grain N yield\n(kg/ha)/plant utilisable N (kg/ha). Plant utilisable N was equal to\nsoil nitrate-N before sowing, plus N fertiliser (if applied), minus\nsoil nitrate-N after harvest, plus soil nitrate-N mineralised during\nthe cropping period. Soil nitrate-N mineralised during the\ncropping period was estimated from the change in soil nitrate-N\nduring this period in an unplanted area adjacent to the trial.\n\n\nStatistical analysis\n\n\nSignificant differences (P = 0.05) between treatments for the\nvarious measurements in each experiment were determined\nusing standard analyses of variance (Snedecor and Cochran\n1967). Experiment 1 was analysed as a factorial design (two\ntillage treatments, each split for two N fertiliser treatments \u00d7\nthree replications). In experiment 2, N fertiliser treatments in\nwheat and barley were analysed as a randomised block design\nwith three replications.\n\n\nResults\n\n\nGross margin\n\n\nSeasonal conditions\n\n\nTo compare the profitability of treatments, gross margins were\ncalculated as the on-farm income from the grain produced, less\nthe variable or operating costs involved in fallow and crop\nmanagement. Gross margins do not take into account fixed or\noverhead expenses such as rates, taxes, insurance, interest and\ndepreciation on machinery and buildings. The on-farm income\nfrom the grain was calculated from the Australian Wheat Board\nprices for grain at the nearest grain depot (Thallon), at the grain\nprotein concentration recorded for each treatment. Variable costs\nwere estimated from a record of operations and materials and their\ncosts (Department of Primary Industries 2000). Gross margins\nwere calculated using grain prices and costs pertaining to each\nyear and also using mean grain prices between 1996 and 2004.\nProfitability of water use ($/ha.mm) was determined by\ndividing the gross margin by the crop water use.\n\n\nAnnual rainfall was above average in 1996, 1997, 1998 and\n1999 and below average in 2000 and 2001 (Table 2). Rainfall\nwas insufficient to enable sowing of the crop in 2000. In-crop\nrainfall was below average in 1997. In 1998, when in-crop\nrainfall was well above average, there were severe infestations\nof yellow spot (Pyrenophora tritici-repentis) in wheat and net\nblotch (Pyrenophora teres) in barley.\n\n\nEffects of tillage practice on wheat grain yield, grain protein\nconcentration, gross margin and water use efficiency\nBecause there were no significant interactions between tillage\npractice and N fertiliser application, mean effects of CT and ZT\nacross N fertiliser treatments are presented. Grain yield of wheat\nwas greater under ZT than under CT in 1997, but was similar\nunder CT and ZT in all other years and for the mean values over\n4 years (Table 3). There were no differences in grain protein\nconcentration between CT and ZT.\nGross margins calculated using grain prices and costs\npertaining to each year were greater under ZT than under CT in\n1997, but greater under CT than under ZT in 1998 and 1999\n(Table 3). Gross margins in 2001 and mean gross margins over\nyears were similar under CT and ZT.\nThere were no differences in crop water use between CT\nand ZT (data not shown), but efficiency of water use for\ngrain production was greater under ZT than under CT in 1997\n(Table 3).\n\n\nInfiltration measurements\n\n\nIn April 1999, infiltration of water into the soil was measured in\nthe field in CT and ZT treatments using a portable rainfall\nsimulator based on the design of Bubenzer and Meyer (1965).\nOne hundred mm of simulated rainfall was applied\nsimultaneously over 1 h to two 100-cm wide and 160-cm long\nplots in each of the three replicates of CT and ZT treatments.\nCrop stubble was cut and removed from one of these plots to\napproximate 0% cover and was added to the other plot to give\n100% cover. Runoff water was collected at the downslope edge\nand routed by vacuum through calibrated tipping buckets. Tip\nrate was logged at 1-min intervals. Water used for simulations\nwas rainwater with electrical conductivity <3 \u00b5S/m.\nAfter rainfall simulation, soil hydraulic conductivity was\nmeasured with disc permeameters (Perroux and White 1988) in\nthe same area of each CT and ZT treatment that received\nsimulated rainfall. Disc permeameters were applied using the\n\n\n4\n\n\nEffects of tillage practice on infiltration and soil water and\nnitrate-N at sowing\nAt the end of the fallow period in 1999, cumulative infiltration\nafter 100 mm of simulated rainfall over 1 h was greater under\nZT than under CT at existing stubble cover levels (60% and\n\n\nAustralian Journal of Experimental Agriculture\n\n\nTillage and N supply in cereals on a Vertosol in south-west Queensland\n\n\n969\n\n\nTable 3. Effects of conventional tillage (CT) and zero tillage (ZT) on plant available soil water\nat sowing (0-120 cm), wheat grain yield, grain protein, gross margin and water use efficiency\nn.s., not significant\n\n\nGrain yield\n(t/ha)\n\n\nTillage\nmethod\n\n\nAvailable soil\nwater at\n\n\nGrain protein Gross margin\n(%)\n\n\nWater use\n\n\n($/ha)\n\n\nefficiency\n(kg/ha.mm)\n\n\nsowing (mm)\n\n\n1997 crop\n\n\nCT\n\n\n106\n\n\n8.3\n\n\n140\n\n\n2.11\n\n\n9.6\n\n\nZT\n\n\n150\n\n\n2.59\n\n\n9.8\n\n\n154\n\n\n9.8\n\n\n1.s.d. (P = 0.05)\n\n\n0.13\n\n\n1.0\n\n\n32\n\n\nn.s.\n\n\nn.s.\n\n\n1998 crop\n\n\n1.62\n\n\n-2\n225\n\u0f56\u0f66\u0fb3\u0f56\u0f0b\u0f66\u0fa4\u0f58\n27\n21\n163\n29\n\n\nCT\n\n\n82\n\n\n13.0\n\n\n2.9\n\n\n105\n\n\nZT\n\n\n1.54\n\n\n13.0\n\n\n2.8\n\n\n1.s.d. (P = 0.05)\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\n1999 crop\n\n\nCT\n\n\n165\n\n\n3.11\n\n\n10.7\n\n\n8.0\n\n\n199\n\n\nZT\n\n\n3.08\n\n\n10.6\n\n\n7.4\n\n\n1.s.d. (P=0.05)\n\n\n27\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\n2001 crop\n\n\n2.27\n\n\n145\n\n\n145\n\n\n12.7\n\n\nCT\n\n\n6.4\n\n\n160\n\n\n2.40\n\n\n55\n\n\n6.5\n\n\nZT\n\n\n12.4\n\n\n1.s.d. (P = 0.05)\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nMean of 1997-2001 crops\n\n\n133\n\n\n2.23\n\n\nCT\n\n\n11.4\n\n\n125\n\n\n6.4\n\n\n\u13c3\u13a2\n\n\n2.36\n\n\n11.4\n\n\n153\n\n\n92\n\n\n6.7\n\n\n1.s.d. (P = 0.05)\n\n\n18\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\n10%, respectively) and with 0% cover, but was greater under CT\nthan under ZT with 100% cover (Table 4). Mean final\ninfiltration rate after 1 h was greater with 100% cover\n(17 mm/h) than with existing cover (5 mm/h) and 0% cover\n(2 mm/h) [1.s.d. (P 0.05) 11]. Saturated hydraulic\nconductivity of surface soil after simulated rainfall was greater\nunder ZT than under CT at existing stubble cover levels\n(Table 4). Mean pore density in the 1.5\u20133.0 mm diameter pore\nrange in surface soil was greater under CT (mean of\n9.3 pores/m\u00b2) than under ZT (mean of 6.8 pores/m\u00b2) [1.s.d.\n(P=0.05)=1.7]. There were no significant differences between\nCT and ZT in saturated hydraulic conductivity and pore density\nin the 1.5\u20133.0 mm diameter pore range in subsurface soil (13 cm\ndepth) (data not shown).\nMean plant available water in the soil profile at sowing over\n4 years was 20 mm greater under ZT than CT (Table 3).\nHowever, in individual years, plant available soil water at\nsowing was greater under ZT than under CT only in 1999, when\nthis effect was not reflected in improved grain yield (Table 3).\n\n\nMean fallow water storage efficiency (percentage of rainfall\nstored in the soil during the fallow period between crops) for the\n1996-97, 1997-98, 1998-99 and 1999-2001 fallow periods was\n16% under ZT and 13% under CT.\nThere were no differences between CT and ZT in soil\nnitrate-N at sowing (data not shown).\n\n\n=\n\n\n=\n\n\nEffects of N fertiliser application on wheat grain yield,\ngrain protein, yield components, gross margin and\nwater use efficiency\n\n\nSince grain yield responses to N fertiliser were similar under CT\nand ZT practices in wheat, only data for ZT practice are\npresented. Wheat grain yield increased with increasing rates of\nN fertiliser addition and N supply in all years except 2001\n(Table 5). However, grain protein concentration increased with\nincreasing rates of N fertiliser addition and N supply in all years.\nThe first increment of N application generally resulted in\nsignificantly greater grain yield and protein than when no N was\napplied. Although increases in grain yield with subsequent\n\n\nTable 4.\n\n\nCumulative infiltration and saturated hydraulic conductivity of surface soil after simulated\nrainfall (100 mm rain over 1 h), for conventional tillage (CT) and zero tillage (ZT) with stubble removed (0%\ncover), at existing stubble cover levels (10% and 60%, respectively) and with stubble added to give 100% cover\n\n\nCumulative infiltration (mm)\nExisting cover\n\n\nTillage method\n\n\nSaturated hydraulic conductivity (mm/h)\nExisting cover\n\n\n100% cover\n\n\n100% cover\n\n\n0% cover\n\n\n0% cover\n\n\n1740\n1720\n\n\nCT\n\n\n19\n\n\n23\n\n\n65\n\n\n195\n\n\n340\n\n\n1260\n\n\n35\n\n\nZT\n\n\n42\n\n\n47\n\n\n440\n\n\n1.s.d. (P = 0.05)\n(tillage x cover)\n\n\n10\n\n\n435\n\n\nAustralian Journal of Experimental Agriculture\n\n\n970\n\n\nG. A. Thomas et al.\n\n\nincrements of N addition were not generally significant, grain\nprotein concentration often continued to increase significantly\nas N fertiliser addition and N supply increased. As a result, gross\nmargins calculated using costs and returns pertaining to each\nyear increased with increasing N supply in 1996, 1997 and 1999\nand for the mean values over 5 years. In a wet growing season\nand disease-affected crop in 1998, there was no effect of N\nfertiliser application on gross margins. When gross margins\nwere calculated using mean grain prices between 1996 and\n2004, mean gross margin increased from $30/ha, with no N\nfertiliser applied, to $150/ha, with the highest rate of N addition\n(data not shown).\nExcept for 1998, grain weight decreased with increasing N\nsupply. The decrease in grain weight was compensated by\n\n\nincreasing tiller numbers, as shown by the increasing number of\nears (Table 5).\nPlant available soil water at sowing and crop water use were\nnot significantly affected by N fertiliser application in\nindividual years (data not shown). However, mean crop water\nuse over 5 years was greater with mean annual N fertiliser\napplication rate of 64 kg N/ha (399 mm) than with no N\nfertiliser application (374 mm) [1.s.d. (P = 0.05) = 23]. Water\nuse efficiency for grain yield of wheat increased with N\nfertiliser application in all years except 2001 and for mean\nvalues over 5 years (Table 5). As for grain yield, the greatest\nincrease in WUE was generally with the first increment of N\naddition. In the wet growing season of 1998, WUE was below\n4 kg/ha.mm, indicating the adverse effects of periodic\n\n\nTable 5. Effects of N supply (N fertiliser applied at sowing + soil nitrate-N before sowing, 0-120 cm) on\ngrain yield, grain protein, grain weight, number of ears, gross margin and water use efficiency for continuous\nzero tillage (ZT) wheat\nn.s., not significant\n\n\nGrain\nyield\n(t/ha)\n\n\nNumber\nof ears/m\u00b2\n\n\nWater use\nefficiency\n(kg/ha.mm)\n\n\nN fertiliser\n(kg N/ha)\n\n\nSoil\nnitrate-N\n\n\nGrain\n\n\nGrain\n\n\nGross\n\n\nprotein\n(%)\n\n\nweight\n(mg)\n\n\nmargin\n($/ha)\n\n\n(kg N/ha)\n\n\n1996 crop\n\n\n3.01\n\n\n8.5\n\n\n8.1\n\n\n75\n\n\n0\n\n\n40.3\n\n\n313\n\n\n189\n\n\n30\n\n\n75\n\n\n3.50\n\n\n38.5\n\n\n306\n\n\n239\n\n\n9.2\n\n\n8.9\n\n\n75\n\n\n355\n\n\n9.8\n\n\n60\n\n\n3.66\n\n\n10.7\n\n\n35.6\n\n\n289\n\n\n12.3\n\n\n90\n\n\n75\n\n\n3.82\n\n\n33.2\n\n\n372\n\n\n353\n\n\n10.0\n\n\n1.s.d. (P=0.05)\n\n\n93\n\n\n0.29\n\n\n0.9\n\n\n2.5\n\n\n46\n\n\n0.9\n\n\nn.s.\n\n\n1997 crop\n\n\n39\n\n\n0\n\n\n2.20\n\n\n7.7\n\n\n38.5\n\n\n178\n\n\n79\n\n\n7.5\n\n\n40\n\n\n9.3\n\n\n2.70\n\n\n9.9\n\n\n35.8\n\n\n281\n\n\n161\n\n\n2.81\n\n\n11.8\n\n\n32.9\n\n\n344\n\n\n190\n\n\n9.3\n\n\n70\n\n\n54\n\n\n2.93\n\n\n333\n\n\n10.3\n\n\n100\n\n\n13.1\n\n\n32.0\n\n\n224\n\n\n1.s.d. (P=0.05)\n\n\n15\n\n\n0.7\n\n\n0.17\n\n\n1.9\n\n\n76\n\n\n48\n\n\n1.4\n\n\n1998 crop\n\n\n33\n90\n35508\n55\n55\n18\n\n\n230\n\n\n1.12\n\n\n2.6\n\n\n11.8\n\n\n0\n\n\n-26\nn.s.\n201712\n57\n47\n\n\n20.7\n\n\n30\n\n\n20.3\n\n\n1.51\n\n\n12.6\n\n\n282\n\n\n3.4\n\n\n40\n\n\n1.83\n1.79\n\n\n13.4\n\n\n20.9\n\n\n309\n\n\n3.8\n\n\n3.7\n\n\n13.8\n\n\n50\n\n\n20.6\n\n\n296\n\n\n1.s.d. (P=0.05)\n\n\n0.27\n\n\n0.7\n\n\n0.4\n\n\nn.s.\n\n\nn.s.\n\n\n1999 crop\n\n\n\u0f66\u0f58\u0fa8\u0f62\u0f9e\u0f67\u0f72\n198\n220\n344\n52\n\n\n0\n\n\n155\n\n\n2\n\n\n36\n\n\n2.58\n\n\n8.0\n\n\n40.9\n\n\n6.2\n\n\n60\n\n\n32\n\n\n11.3\n\n\n3.84\n\n\n38.2\n\n\n234\n\n\n9.4\n9.1\n9.2\n\n\n38.6\n\n\n12.2\n\n\n90\n\n\n37\n\n\n4.01\n\n\n297\n\n\n150\n\n\n37.3\n\n\n49\n\n\n4.02\n\n\n13.4\n\n\n240\n\n\n1.s.d. (P= 0.05)\n\n\n0.28\n\n\n0.9\n\n\n1.8\n\n\n115\n\n\n1.3\n\n\nn.s.\n\n\n2001 crop\n\n\n10.0\n\n\n36.5\n\n\n186\n\n\n8.6\n\n\n0\n\n\n90\n\n\n2.92\n\n\n255\n255\n\n\n40\n\n\n90\n\n\n3.13\n\n\n33.5\n\n\n238\n\n\n11.7\n\n\n8.8\n\n\n90\n\n\n32.2\n\n\n333\n\n\n60\n\n\n3.00\n\n\n13.4\n\n\n256\n\n\n8.0\n\n\n100\n\n\n31.7\n\n\n269\n\n\n280\n\n\n90\n\n\n2.99\n\n\n14.9\n\n\n7.9\n\n\n1.s.d. (P= 0.05)\n\n\n3.6\n\n\n77\n\n\n1.8\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nMean of 1996-2001 crops\n\n\n72237\n6.7\n8.0\n7.9\n8.3\n0.7\n\n\n0\n\n\n54\n\n\n2.36\n\n\n9.2\n\n\n35.4\n\n\n226\n\n\n86\n\n\n40\n\n\n59\n\n\n33.3\n\n\n274\n\n\n168\n\n\n2.94\n\n\n10.9\n\n\n328\n\n\n32.0\n\n\n60\n\n\n12.3\n\n\n64\n\n\n3.06\n\n\n202\n\n\n3.11\n\n\n98\n\n\n71\n\n\n13.5\n\n\n30.9\n\n\n302\n\n\n250\n\n\n1.s.d. (P=0.05)\n\n\n9\n\n\n0.13\n\n\n0.4\n\n\n0.9\n\n\n47\n\n\n20\n\n\nAustralian Journal of Experimental Agriculture\n\n\nTillage and N supply in cereals on a Vertosol in south-west Queensland\n\n\n971\n\n\nwaterlogging and plant disease on crop growth and yield in\nthis year.\nGross margin/water use increased with N fertiliser addition\nand greater N supply in 3 out of 5 years (1996, 1997 and 1999)\n(data not shown) and for mean values over 5 years (Fig. 1). The\nhighest value recorded for gross margin/water use in individual\nyears was $0.93/ha.mm, for wheat with the highest N fertiliser\napplication and the greatest N supply in 1996.\n\n\n0.8\n\n\nMean gross margin/water use\n\n\nWheat\n\n\n\u2610 Barley\n\n\n\u0627\u0633\n\n\n0.6\n\n\n($/ha.mm)\n\n\n0.4\n\n\nEffects of N fertiliser application on barley grain yield,\ngrain protein concentration, yield components, gross\nmargin and water use efficiency\n\n\n0.2\n\n\n0\n\n\nSignificant increases in barley grain yield resulted from the\naddition of N fertiliser in 1996 and 1999 (Table 6). In 1996,\n1997, 1998 and 1999 seasons, grain protein concentration with\nN fertiliser application was in the range to qualify for malting\ngrade barley (9-12% at 0% moisture content) and, except for the\n1998 crop, resulted in similar gross margins to those for wheat\nat 12.3% grain protein in 1996 and APH grade wheat (\u226513%\ngrain protein) in 1997 and 1999. Barley was of feed grain\nquality and of lower value because of weather damage in 1998\nand high grain protein in 2001, due to high residual nitrate-N in\nsoil from the unavoidable long fallow (no crop in 2000 due to\nlack of sowing rain). Where gross margins were calculated\nusing costs and returns pertaining to each year, mean gross\n\n\n0\n\n\n40\n\n\n98\n\n\n0\n\n\n29\n\n\n64\n\n\nMean N fertiliser application rate (kg N/ha.year)\n\n\nFig. 1. Mean effects of N fertiliser application on gross margin/water use\nfor wheat and barley, 1996\u20132001. Vertical line indicates the 1.s.d. value\n(P = 0.05).\n\n\nmargin over 5 years, where malting grade barley was targeted,\nwas almost double that where it was not. Using mean grain\nprices between 1996 and 2004, mean gross margin where\nmalting barley was targeted was $180/ha (data not shown).\nGrain weight of barley was greater with N fertiliser\napplication than without in 1999, and lower with N fertiliser\n\n\nTable 6. Effects of N supply (N fertiliser applied at sowing + soil nitrate-N before sowing, 0-120 cm) on grain\nyield, grain protein (0% moisture content), grain weight, number of ears, gross margin and water use efficiency\nfor zero tillage (ZT) barley\nn.s., not significant\n\n\nN fertiliser\n\n\nWater use\nefficiency\n(kg/ha.mm)\n\n\nGrain\nyield\n(t/ha)\n\n\nGrain\n\n\nNumber\nof ears/m\u00b2\n\n\nSoil\nnitrate-N\n(kg N/ha)\n\n\nGrain\nweight\n(mg)\n\n\nGross\nmargin\n($/ha)\n\n\n(kg N/ha)\n\n\nprotein\n(%)\n\n\n1996 crop\n\n\n3.15\n\n\n174\n\n\n75\n\n\n47.5\n\n\n0\n\n\n9.1\n\n\n40\n\n\n75\n\n\n3.68\n\n\n11.6\n\n\n38.9\n\n\n1036\n\n\n448\n\n\n10.2\n\n\n1.s.d. (P=0.05)\n\n\n2.5\n\n\n0.29\n\n\n46\n\n\n1.0\n\n\nn.s.\n\n\n1997 crop\n\n\n2.48\n\n\n11.1\n\n\n39.1\n\n\n0\n\n\n67\n\n\n245\n\n\n461\n\n\n10.9\n\n\n226\n\n\n9.5\n\n\n20\n\n\n68\n\n\n2.48\n\n\n12.4\n\n\n39.2\n\n\n511\n\n\n1.s.d. (P =\n\n\n= 0.05)\n\n\n0.8\n\n\n1.4\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\n1998 crop\n\n\n27.4\n\n\n0\n\n\n1.92\n\n\n49\n\n\n10.5\n\n\n394\n\n\n20\n\n\n4.5\n\n\n379\n\n\n2.25\n\n\n11.4\n\n\n5.2\n\n\n20\n\n\n49\n\n\n28.4\n\n\n33\n\n\n1.s.d. (P=0.05)\n\n\n0.5\n\n\n0.7\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\n1999 crop\n\n\n0\n\n\n36\n\n\n2.96\n\n\n7.2\n\n\n37.0\n39.2\n\n\n476\n\n\n62\n\n\n7.1\n\n\n9.8\n\n\n533\n\n\n45\n\n\n36\n1\nn.s.\n\n\n4.06\n\n\n360\n\n\n9.2\n\n\n1.s.d. (P=0.05)\n\n\n1.3\n\n\n0.28\n\n\n1.0\n\n\n1.8\n\n\nn.s.\n\n\nn.s.\n\n\n2001 crop\n\n\n0\n\n\n95\n\n\n3.17\n\n\n13.2\n\n\n33.8\n\n\n464\n\n\n155\n\n\n8.6\n\n\n3.15\n\n\n376\n\n\n20\n\n\n95\n\n\n14.4\n\n\n31.1\n\n\n133\n\n\n8.9\n\n\n1.s.d. (P=0.05)\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nn.s.\n\n\nMean of 1996-2001 crops\n\n\n2.74\n\n\n65\n\n\n9.3\n\n\n131\n\n\n0\n\n\n37.0\n\n\n431\n\n\n7.9\n\n\n35.3\n\n\n65\n\n\n29\n\n\n3.12\n\n\n10.8\n\n\n567\n\n\n240\n\n\n8.5\n\n\n1.s.d. (P = 0.05)\n\n\n0.4\n\n\n0.9\n\n\n0.13\n\n\n47\n\n\n20\n\n\nn.s.\n\n\nn.s.\n\n\nAustralian Journal of Experimental Agriculture\n\n\n972\n\n\nG. A. Thomas et al.\n\n\nthan without in 1996 and for mean values over 5 years (Table 6).\nTiller numbers were not affected by N application in any\nindividual year, but mean tiller numbers over 5 years were\ngreater with N fertiliser application than without.\nThere were no significant differences in crop water use\nbetween N fertiliser treatments in barley in any year or across\nyears (data not shown). Water use efficiency was significantly\nlower in N fertilised treatment than without N application in\n1997, but increased with N fertiliser application in 1998 and\n1999 seasons (Table 6). Mean gross margin/water use over\n5 years for barley increased significantly with N fertiliser\napplication (Fig. 1). Mean gross margin/water use values were\ngreater in barley than in wheat where no N fertiliser was applied,\nwhile maximum values were similar in wheat and barley. The\nhighest gross margin/water use in barley in individual years\n($1.24/ha.mm) was obtained for malting grade barley in 1996\n(data not shown).\n\n\n70\n\n\ny = 0.41x\nR2 = 0.90\n\n\n60\n\n\nIncreased grain N (kg/ha)\n\n\n50\n\n\n40\n\n\n30\n\n\n20\n\n\n10\n\n\n50\n\n\n100\n\n\n150\n\n\n200\n\n\nFertiliser N applied (kg N/ha)\n\n\nFig. 3. Relationship between increased grain N (grain N yield in the N\ntreatment - grain N yield in the control) in wheat and barley and N fertiliser\napplications. The slope of the regression line \u00d7 100 is the apparent N\nrecovery in grain.\n\n\nResidual benefit of applied N fertiliser\n\n\nIn all years, the N fertiliser application to wheat resulted in a\nresidual benefit to soil N supply for the following crop, as\nmeasured by soil nitrate-N (0-120 cm depth) before sowing\n(Fig. 2). Thus, in this semiarid, marginal crop growing\nenvironment, a substantial amount of N fertiliser that is unused\nby a crop is carried over to the next season.\n\n\nPrediction of grain protein concentration from the ratio of\navailable soil water/N supply before sowing\nMeasured grain protein concentrations of wheat and barley were\ncompared with the predicted values using the equation derived\nby Dalal et al. (1997):\n\n\nNitrogen use efficiency for grain N yield\n\n\nApparent N fertiliser use efficiency for wheat and barley grain\n[(grain N yield of N treatment - grain N yield of control)/N\nfertiliser applied] was 41% (Fig. 3), with no significant tillage\nor N fertiliser application rate effects. There were no\nsignificant differences between N fertiliser treatments in NUE\nfor grain N yield. Accounting for residual N fertiliser and in-\ncrop N mineralisation increased NUE for grain N yield, which\nranged from a mean of 43% in 1996 to 59% in 1999 (data not\nshown).\n\n\nGrain protein (%) = 6.52 + 10.0 \u00d7 exp(\u22120.44 available\nwater/available N)\n\n\nwhere plant available water (mm) and available N (soil nitrate-N\n+ N fertiliser) (kg/ha) in 0-120 cm depth of soil were measured\nat or near sowing.\n\n\n16\n\n\nPredicted grain protein concentration (%)\n\n\ny= 1.00 (\u00b1 0.02 s.e.)x\nR\u00b2 = 0.75\n\n\n14\n\n\n140\n\n\n120\n\n\nResidual fertiliser N (kg/ha)\n\n\n12\n\n\ny=0.74x-22.36\nR\u00b2 =\n\n\n100\n\n\n-= 0.68\n\n\n0\n\n\n80\n\n\n60\n60\n\n\n40\n40\n\n\n20\n\n\n6\n\n\n8\n\n\n10\n\n\n12\n\n\n14\n\n\n16\n\n\nMeasured grain protein concentration (%)\n\n\n-20\n\n\n0\n\n\n50\n\n\n100\n\n\n150\n\n\n200\n\n\nPrevious year fertiliser N application (kg N/ha)\n\n\nFig. 4. Relationship between the predicted (Dalal et al. 1997) (y) and\nmeasured (x) grain protein concentrations for wheat and barley. The\nequation of regression line (full) through the origin is: y = 1.00 (\u00b10.02 s.e.)x\n(R2 = 0.75). The dotted lines show \u00b11% range in grain protein about the\nregression line. The dashed line is the 1:1 line.\n\n\nFig. 2.\nSoil nitrate-N (0-120 cm) at sowing following N fertiliser\napplication in the previous year, in excess of that with no N fertiliser\napplication.\n\n\nAustralian Journal of Experimental Agriculture 973\n\n\nTillage and N supply in cereals on a Vertosol in south-west Queensland\n\n\nwater at sowing in 1999. The lower gross margin under ZT than\nunder CT in 1998 and 1999 was due to higher fallow weed\ncontrol costs with herbicide under ZT than tillage costs under\nCT in these years.\nOther experiments in southern Queensland have shown that\nthere is a greater advantage in yield with ZT over CT in drier\ngrowing seasons, as a result of greater soil water storage or\nwater use efficiency (Freebairn et al. 1986; Marley and Littler\n1989; Radford et al. 1992; Thomas et al. 1995; Strong et al.\n1996a). Retention of crop residues on the soil surface, as occurs\nunder ZT, has also been shown to extend crop sowing time and\nfacilitate timeliness of sowing during dry weather (Radford and\nNielsen 1983), thus providing more sowing opportunities and\nenhancing potential yield benefits associated with optimum\nsowing time (Woodruff and Tonks 1983). Therefore, ZT may\nprovide yield and economic benefits to longer-term crop\nproduction under the drier conditions more likely to be\nexperienced in this environment than those that occurred from\n1996 to 2001.\nWe observed significant grain yield responses to N fertiliser\napplication in all cropping years except the 2001 season, when\nthe crop followed a long fallow of 18 months (due to lack of\nsowing rains in 2000) and 80 kg N/ha was accumulated in the\nsoil profile. Except for a wet growing season in 1998, gross\nmargins significantly increased from fertiliser application. We\nsuggest that profitable responses to N fertiliser addition can be\nobtained in the marginal cropping area of south-west\nQueensland on deep grey clay soil that has been cropped for\nseveral years and is of relatively low fertility. However, all trials\nthat produced grain yields commenced with soil water storage\n>60% of the plant available water capacity, or received above\naverage rainfall over the growing season. In drier growing\nseasons, crops were either not sown (2000, 2002) or failed to\nproduce harvestable grain yield (2003, data not shown).\nRobinson et al. (1999) concluded from simulation modelling\nthat positive responses to applying 60 kg N/ha at sowing\noccurred in 97% of years under conditions similar to those\nexperienced in 1996, but in less than 40% of years when soil\nwater at sowing was less than 50% plant available water\ncapacity. However, on a Vertosol at Warra, ~300 km north-east\nof Nindigully, N application increased gross returns from five of\nthe seven wheat crops grown, because of increasing grain yield\nand/or grain protein responses (Strong et al. 1996a). These\nresults occurred over the 1987-94 period, which included a\n4-year period of exceptionally dry seasonal conditions. At\nWarra, although grain yield responses were inconsistent in the\nfirst year of fertiliser application where no N fertiliser had been\napplied to preceding crops, grain protein usually increased with\nincreasing N application. Holford and Doyle (1992) also\nshowed profitable grain yield responses with wheat to varying\nrates of N application during a dry growing season in northern\nNew South Wales. Similar effects to those at Warra and in\nnorthern New South Wales may also occur under drier seasonal\nconditions in the south-west Queensland region.\nThis study has also shown that unused available soil and\nfertiliser N generally remains in the soil profile to benefit the\nfollowing crop. This benefit results in a lower N fertiliser\nrequirement in the following crop. Similar effects were also\nrecorded at Warra (Strong et al. 1996b). However, unused\n\n\nN fertiliser requirement using available\nsoil water/N supply (kg N/ha)\n\n\n160\n\n\n120\n\n\n80\n\n\nHOWWET? with\nestimated WUE\nHOWWET? with\nmeasured WUE\n1:1 line\n\n\n=\n\n\n0\n\n\n120\n\n\n40\n\n\n80\n\n\n160\n\n\nN fertiliser requirement using HOWWET?\n(kg N/ha)\n\n\nFig. 5. Relationship between the N fertiliser requirement to target grain\nprotein concentration of 10.2% in barley (0% moisture content) and 13% in\nwheat (12% moisture content) using the available soil water/N supply\nmethod (Dalal et al. 1997) to calculate y, and HOWWET? (Freebairn et al.\n1997) to calculate x with (i) a commonly used estimate of water use\nefficiency (WUE) (y = 1.41x + 13.31, R2 = 0.91) and (ii) a measured mean\nvalue for WUE from this experiment (y = 1.07x - 7.08, R\u00b2 = 0.92).\n\n\nThere was a significant and a 1:1 relationship between the\npredicted (y) and measured (x) grain protein values (Fig. 4).\nIncluding a commonly used estimate of WUE, HOWWET?\n(Freebairn et al. 1997) under-predicted grain yield and therefore\nthe N fertiliser requirements to achieve grain protein\nconcentration of 10.2% in barley and 13% in wheat, compared\nwith the available soil water/N supply method (Fig. 5). Using a\nmeasured mean value for water use efficiency from the\nexperiment in HOWWET? resulted in better agreement between\nthese two methods.\n\n\nDiscussion\n\n\nAt the existing levels of soil cover during the 1998-99 fallow\nperiod, the soil under ZT had four times greater hydraulic\nconductivity and two times greater water infiltration than the\nsoil under CT, resulting in 27 mm extra water in the soil profile\nunder ZT practice at the end of the fallow period. In Vertosols,\nDalal (1989) showed a much higher rate of Cl leaching under ZT\nthan CT, indicating greater water infiltration under ZT, and\nTurpin et al. (1999) measured much faster water infiltration and\nbromide movement under ZT than CT, especially when stubble\nwas retained.\nAlthough mean plant available water in the soil profile at\nsowing over 4 years was 20 mm greater under ZT than under CT,\nthe mean grain yields and gross margins were similar under CT\nand ZT over this period. ZT resulted in higher grain yield, gross\nmargin and water use efficiency than CT in 1997, when plant\navailable soil water levels at sowing were similar for the two\ntillage practices, indicating that ZT made more efficient use of\nstored soil water and in-crop rainfall in this near-average\ngrowing season. CT and ZT resulted in similar grain yields in\n1998, 1999 and 2001, even though available soil water before\nsowing was higher under ZT than under CT in 1999. Above\naverage in-crop rainfall may have negated differences in soil\n\n\nAustralian Journal of Experimental Agriculture\n\n\n974\n\n\nG. A. Thomas et al.\n\n\nnitrate-N in the soil profile may be lost occasionally by leaching\nbelow the root zone and by denitrification, depending on\nseasonal conditions. Its benefit (and that of fresh N fertiliser\napplication) may also be lost to the following crop if factors\nsuch as plant disease limit the response to the additional\nN supply.\nApparent N fertiliser use efficiency of 41% (Fig. 3) is\ncomparable to that obtained by Strong et al. (1996b) in the\nWarra experiment (45%) over a 6-year period that had a range\nof seasonal conditions. Values for NUE for grain N yield\n(43-59%) were similar to those recorded by Huggins and Pan\n(1993) (51-72%) and Dalal et al. (1996) (43\u201365%).\nWe found that the relationship between the ratio of plant\navailable water and available N supply at sowing and grain\nprotein concentration developed by Dalal et al. (1997) for wheat\nand barley is applicable in this environment. As concluded by\nDalal et al. (1997), the amount of plant available water (mm) to\na depth of 120 cm at sowing should be equally matched by\navailable N (kg N/ha), to produce approximately the 13% protein\nconcentration required for the APH grade of wheat. Malting\ngrade barley (10.2% protein concentration at 0% moisture\ncontent) would require ~40% of the available N needed for APH\nwheat. This knowledge enables the protein concentration for\nAPH grade wheat or malting grade barley to be targeted by\nmeeting any deficit in N supply from nitrate-N present in the\nsoil with that from fertiliser application at sowing. In 1996,\ngrain protein concentration did not reach the level for APH\nclassification, even with the highest rate of N addition\n(90 kg N/ha), indicating that the top N fertiliser application rate\nshould have been higher for this relatively high-yielding crop.\nAt sowing in 1996, soil water characteristics of the site (crop\nlower limit) had not been clearly defined, and available soil\nwater at sowing was apparently underestimated. Subsequent\ndeterminations showed that plant available water/nitrogen ratio\nwas greater than 1 (1.3) and, therefore, APH grade wheat would\nnot have been expected (estimated grain protein, 12.2% v.\nobserved value of 12.3%). However, grain protein concentration\nwas >13% with the highest rate of N addition in all other\ncropping years.\nIn this example, the ability of HOWWET? (Freebairn et al.\n1997) to predict N fertiliser requirements was greatly improved\nwhen used with measured, rather than estimated, water use\nparameters. This indicates the importance of on-farm records in\nvalidating such decision support programs.\nThe highest WUE for grain yield recorded in wheat was ~10.0\nkg grain/ha.mm of water used, which was lower than the\nmaximum value recorded for wheat by Radford et al. (1992)\n(18 kg grain/ha.mm) and Thomas et al. (1995) (15 kg\ngrain/ha.mm) in south-west Queensland. The considerably lower\nvalues recorded in 1998 (2.6-3.7 kg grain/ha.mm) were\nassociated with above average rainfall during the growing period,\nwhich resulted in periodic waterlogging, and the severe\noccurrence of the leaf diseases, such as yellow spot in wheat and\nnet blotch in barley. The lower values for WUE for grain yield\nobtained at Nindigully may have been associated with the timing\nand effectiveness of rainfall in relation to stage of growth of the\ncrop and higher potential evaporative demand. Other factors,\nsuch as soil-borne disease, may also have limited the efficiency\nof water use under the continuous winter cereal system.\n\n\nWildermuth (2001) found that the incidence and severity of\ncrown rot (Fusarium pseudograminearum) in continuous wheat\nincreased at a faster rate and to a higher level with the application\nof N fertiliser in this experiment. However, higher wheat yield\nwith N fertiliser application was associated more consistently\nwith greater water use efficiency than with greater water use.\nGross margin/water use could be a useful concept for\ncomparing the relative value and economic efficiency of crops\nand management practices, as it integrates grain yield, grain\nquality and water use. Although it is influenced by the relative\nprices of crops within and between years, a value of ~$1/ha.mm,\nwhich was the highest value recorded in these experiments,\nappears to be a potential benchmark in this region.\n\n\nConclusions\n\n\nZero tillage has potential to improve winter cereal production in\nsouth-west Queensland, particularly in drier growing seasons\nthan experienced in this study, by resulting in greater fallow soil\nwater storage or crop water use efficiency than conventional\ntillage.\nProfitable responses to N fertiliser addition in wheat and\nbarley can be obtained on long-term cultivated Vertosols in\nsouth-west Queensland when soil water reserves at sowing are\nat least 60% of plant available water capacity or rainfall during\nthe growing season is above average. Presowing soil profile\nwater and nitrate-N levels can be used to determine the N\nfertiliser rate required to target grain protein concentration in\nthis environment. An integrative benchmark for improved N\nfertiliser management appears to be gross margin/water use of\n~$1/ha.mm.\n\n\nAcknowledgements\n\n\nThe authors thank the Grains Research and Development Corporation\n(GRDC) for financial assistance and Mr and Mrs D. Hill, 'Dunkerry South'\nNindigully, who provided land and other assistance for the trial. We\ngratefully acknowledge the involvement and advice of members of the trial\nmanagement committee. We also thank Dr J. Standley, Mr J. Hagedoorn,\nMr K. Spann, Mrs J. Glasby and Mrs A. Pumfrey for soil and plant chemical\nanalyses, Mr G. McNamara, Ms J. Riddell and Ms M. Fraser for technical\nassistance, Messrs, J. Henderson, G. Pauli, G. Kedzlie, D. Cooper, R. Norris,\nD. Baills, M. Weston and P. Marsh for field operations, and Mrs A. Kelly and\nMs K. Bell for advice on trial design and statistical analysis.\n\n\nReferences\n\n\nAustralian Bureau of Statistics (2001) 'Queensland Year Book 2001. No. 58.\nABS Catalogue No. 1301.3.' (Australian Bureau of Statistics: Canberra)\nBest EK (1976) An automated method for determining nitrate nitrogen in\nsoil extracts. Queensland Journal of Agricultural and Animal Sciences\n33, 161-166.\nBruce RC, Rayment GE (1982) Analytical methods and interpretations used\nby the Agricultural Chemistry Branch for soil and land use surveys.\nQueensland Department of Primary Industries, Bulletin QB2004.\nBubenzer GD, Meyer LD (1965) Simulation of rainfall and soils for\nlaboratory research. Transactions of the American Society of\nAgricultural Engineers 8, 73\u201375.\nCrooke WM, Simpson WE (1971) Determination of ammonium in Kjeldahl\ndigests of crops by an automated procedure. Journal of the Science of\nFood and Agriculture 22, 9-10. doi:10.1002/jsfa.2740220104\nDalal RC (1989) Long-term effects of no-tillage, crop residue, and nitrogen\napplication on properties of a Vertisol. Soil Science Society of America\nJournal 53, 1511\u20131515.\n\n\nAustralian Journal of Experimental Agriculture 975\n\n\nTillage and N supply in cereals on a Vertosol in south-west Queensland\n\n\nDalal RC, Mayer RJ (1986a) Long-term trends in fertility of soils under\ncontinuous cultivation and cereal cropping in southern Queensland. I.\nOverall changes in soil properties and trends in winter cereal yields.\nAustralian Journal of Soil Research 24, 265-279. doi:10.1071/\nSR9860265\n\n\nnitrogen accumulation and wheat performance in south west\nQueensland. Soil and Tillage Research 22, 73-93. doi:10.1016/0167-\n1987(92)90023-5\n\n\nReynolds WD, Elrick DE (1991) Determination of hydraulic conductivity\nusing a tension infiltrometer Soil Science Society of America Journal 55,\n633-639.\nRobinson JB (2005) Understanding and applying decision support systems\nin farming systems research in Australia. PhD thesis, University of\nWestern Sydney, Australia.\nRobinson JB, Freebairn DM, Dimes JP, Dalal RC, Thomas GA, Weston EJ\n(1999) Modelling wheat production from low-rainfall farming systems\nin northern Australia. Environment International 25, 861-870.\ndoi:10.1016/S0160-4120(99)00061-6\nSnedecor GW, Cochran WG (1967) \u2018Statistical methods.' 6th edn. (Iowa\nState University Press: Ames, IA)\nStrong WM, Dalal RC, Weston EJ, Cooper JE, Lehane KJ, King AJ, Chicken\nCJ (1996a) Sustaining productivity of a Vertisol at Warra, Queensland,\nwith fertilisers, no-tillage or legumes. 2. Long-term fertiliser nitrogen\nneeds to enhance wheat yields and grain protein. Australian Journal of\nExperimental Agriculture 36, 665-674. doi: 10.1071/EA9960665\nStrong WM, Dalal RC, Cahill MJ, Weston EJ, Cooper JE, Lehane KJ,\nKing AJ, Chicken CJ (1996b) Sustaining productivity of a Vertisol at\nWarra, Queensland, with fertilisers, no-tillage or legumes. 3. Effects of\nnitrate accumulated in fertilised soil on crop response and profitability.\nAustralian Journal of Experimental Agriculture 36, 675\u2013682.\ndoi:10.1071/EA9960675\n\n\nDalal RC, Mayer RJ (1986b) Long-term trends in fertility of soils under\ncontinuous cultivation and cereal cropping in southern Queensland. V.\nRate of loss of total nitrogen from the soil profile and changes in\ncarbon:nitrogen ratios. Australian Journal of Soil Research 24,\n493-504. doi: 10.1071/SR9860493\n\n\nDalal RC, Strong WM, Weston EJ, Cooper JE, Lehane KJ, King AJ (1996)\nComparison of legumes and fertilizer nitrogen for wheat production in\nsubtropical Australia. In \u2018Nitrogen economy in tropical soils'. (Ed.\nN Ahmed) pp. 363\u2013369. (Kluwer Academic Publishers: Dordrecht,\nNetherlands)\nDalal RC, Strong WM, Weston EJ, Cooper JE, Thomas GA (1997)\nPrediction of grain protein in wheat and barley in a subtropical\nenvironment from available water and nitrogen in Vertisols at sowing.\nAustralian Journal of Experimental Agriculture 37, 351-357.\ndoi: 10.1071/EA96126\nDepartment of Primary Industries (2000) 'Winter crop management notes.'\n(Kingswood Press: Underwood, Qld)\nFreebairn DM, Ward LD, Clarke AL, Smith GD (1986) Research and\ndevelopment of reduced tillage systems for Vertisols in Queensland,\nAustralia. Soil and Tillage Research 8, 211\u2013219. doi:10.1016/0167-\n1987(86)90335-1\nFreebairn DM, Loch RJ, Cogle AL (1993) Tillage methods and soil and\nwater conservation in Australia. Soil and Tillage Research 27, 303-325.\ndoi: 10.1016/0167-1987(93)90074-Y\nFreebairn DM, Hamilton NA, Cox PG, Glanville SF, Dimes JP,\nHolzworth D (1997) 'HOWWET? - estimates how much rain is stored\nin the soil during a fallow and how much soil nitrate has accumulated.\nA computer program, Version Number: 2.10.' (Agricultural Production\nSystems Research Unit, QDPI/DNR/CSIRO: Toowoomba, Qld)\nGardner EA (1985) Soil water. In \u2018Identification of soils and interpretation\nof soil data'. (Ed. GE Rayment) pp. 197\u2013234. (Australian Society of Soil\nScience Incorporated, Queensland Branch: Brisbane)\nHolford ICR, Doyle AD (1992) Yield responses and nitrogen fertiliser\n\n\nThomas GA, Gibson G, Nielsen RGH, Martin WD, Radford BJ (1995)\nEffects of tillage, stubble, gypsum, and nitrogen fertiliser on cereal\ncropping on a red-brown earth in south-west Queensland. Australian\nJournal of Experimental Agriculture 35, 997-1008. doi:10.1071/\nEA9950997\n\n\nThomas GA, Felton WL, Radford BJ (1997) Tillage and crop residue\nmanagement. In \u2018Sustainable crop production in the sub-tropics: an\nAustralian perspective'. Queensland Department of Primary Industries\nInformation Series QI97035. (Eds AL Clarke, PB Wylie) pp. 195\u2013213.\n(Queensland Department of Primary Industries: Brisbane)\nTurpin JE, Bridge BJ, Orange D, Thompson JP (1999) Water and bromide\nmovement in a Vertosol under four fallow management systems.\nAustralian Journal of Soil Research 37, 75-89. doi:10.1071/S97108\nWeston EJ, Nason CN, Armstrong RDH (1975) Resources study and\nproblem analysis for primary industries in the Condamine-Maranoa\nBasin of southern Queensland. Queensland Department of Primary\nIndustries, Division of Plant Industry Bulletin No. 731.\nWildermuth GB (2001) Crown rot - an important disease in the Western\nDowns. In 'Western Farming Systems Project Results Booklet 2001'.\n(Ed. KJ Lehane) pp. 30-31. Queensland Government (Department of\nPrimary Industries, Natural Resources and Mines) and Grains Research\nand Development Corporation, QI 01093. (Greenridge Press:\nToowoomba, Qld)\nWoodruff DR, Tonks J (1983) Grain yield variation in relation to time of\nflowering of wheat genotypes. Australian Journal of Agricultural\nResearch 34, 1-11. doi: 10.1071/AR9830001\n\n\nrequirements of wheat in relation to soil nitrate levels at various depths.\nAustralian Journal of Soil Research 30, 683-694. doi:10.1071/\nSR9920683\n\n\nHuggins DR, Pan WL (1993) Nitrogen efficiency component analysis: an\nevaluation of cropping system differences in productivity. Agronomy\nJournal 85, 898-905.\nIsbell RF (1996) \u2018The Australian soil classification.' (CSIRO Publishing:\nMelbourne)\nMarley JM, Littler JW (1989) Winter cereal production on the Darling\nDowns an 11 year study of fallowing practices. Australian Journal of\nExperimental Agriculture 29, 807\u2013827. doi: 10.1071/EA9890807\nPerroux KM, White I (1988) Designs for disc permeameters. Soil Science\nSociety of America Journal 52, 1205-1215.\nRadford BJ, Nielsen RGH (1983) Extension of crop sowing time during dry\nweather by means of stubble mulching and water injection. Australian\nJournal of Experimental Agriculture and Animal Husbandry 23,\n302-308. doi: 10.1071/EA9830302\nRadford BJ, Gibson G, Nielsen RGH, Butler DG, Smith GD, Orange DN\n(1992) Fallowing practices, soil water storage, plant-available soil\n\n\nManuscript received 1 June 2005, accepted 6 February 2007\n\n\nhttp://www.publish.csiro.au/journals/ajea\n"}, "expected_output": {"claims": [{"unit": "kg N/ha", "value": 29, "evidence": ["29", "Table 6. Effects of N supply (N fertiliser applied at sowing + soil nitrate-N before sowing, 0-120 cm) on grain yield, grain protein (0% moisture content), grain weight, number of ears, gross margin and water use efficiency for zero tillage (ZT) barley"]}, {"unit": "kg N/ha", "value": 20, "evidence": ["20", "Table 6. Effects of N supply (N fertiliser applied at sowing + soil nitrate-N before sowing, 0-120 cm) on grain yield, grain protein (0% moisture content), grain weight, number of ears, gross margin and water use efficiency for zero tillage (ZT) barley"]}, {"unit": "kg N/ha", "value": 40, "evidence": ["40", "Table 6. Effects of N supply (N fertiliser applied at sowing + soil nitrate-N before sowing, 0-120 cm) on grain yield, grain protein (0% moisture content), grain weight, number of ears, gross margin and water use efficiency for zero tillage (ZT) barley"]}, {"unit": "kg N/ha", "value": 45, "evidence": ["45", "Table 6. Effects of N supply (N fertiliser applied at sowing + soil nitrate-N before sowing, 0-120 cm) on grain yield, grain protein (0% moisture content), grain weight, number of ears, gross margin and water use efficiency for zero tillage (ZT) barley"]}]}, "metadata": {"product_category": "Chemical products", "request_id": "req_416307f05da44f81"}}
{"id": "05c4008ab8bcf76cfdaecee3", "input": {"query": "What is the annual maintenance lime application rate in kg per hectare or tonnes per hectare for Australian grain farming systems?", "source_url": "https://cdn.environment.sa.gov.au/landscape/docs/ki/liming-stanton-fact2-2017.pdf", "document_text": "The dollars and sense of liming: The Stantons' story\n\n\nSecond and third generation Island farmers, the Stantons,\n\n\nrun a mixed farming enterprise of cropping, sheep and\ncattle.\n\n\nMost of the farm land is inherently acidic and soil testing\nover the last 10 years has shown a steady decline in pH. The\nStantons place a high priority on maintaining soil fertility\nlevels, especially in cropping paddocks. In the past, paddocks\nwith low pH were limed prior to being cropped. This means\nthat only some paddocks on each property have been limed,\nwhile test results show that nearly all paddocks will need to\nbe limed at some stage.\n\n\n\"The current strategy is to set up a rotation of liming\napproximately one fifth of the total land base or about 300\nha per year to assist in budgeting. It's envisaged that possibly\ntwo full rotations will be needed to get all paddocks up to\nideal levels of close to 5.5 pHCaCl. The only deviation to this\nplan is that paddocks being sown to lupins will not be limed\nthe year the lupins are sown. A budget of $20,000 has been\nallocated per year to enable this to occur\" explained Jenny\nStanton.\n\n\nThe Stock Jo\n\n\nFARM FACTS\n\n\nPrimary Industries and Regions SA (PIRSA) staff with\nassistance from the GRDC and Department of Environment,\nWater and Natural Resources (DEWNR), have developed\nthree computerised decision support tools to assist\nlandholders and advisers to make better decisions in treating\nsoil acidity. They are:\n\n\nOperators:\n\n\nRichard and Kate Stanton with sons and daughters-in-law\nWill and Jenny and Michael and Sarah.\n\n\nLocation:\n\n\nStokes Bay and on the central plateau\n(SE of Parndana).\n\n\n\u00bb Acid $Cost\n\n\nused to estimate the impact of\n\n\nacidification on production (i.e. the cost of not liming)\n\n\nSize:\n\n\nLimeCheque (Lime Sources Cost Comparison) \u2014 a tool\nfor calculating lime application rates for acidic soils and\ncomparing the costs of lime from different suppliers\n\n\n\u00bb\n\n\n1,500 arable ha (spread over four farms).\nCropping rotation:\n\n\nLegume (lupins on sandy soils and beans on ironstone\nsoils), canola, cereal (wheat, barley or oats).\n\n\nMaintenance Liming Rate Calculator \u2014 a tool for\n\n\n>>>\n\n\ncalculating the replacement lime requirement to offset\n\n\nLivestock:\n\n\nannual acidification and maintain the current pH of the\nsoil.\n\n\nMerinos, 120 cattle, Poll Dorset stud.\n\n\nPastures:\n\n\nAll the tools can be found at\n\n\nSubclovers and annual grasses plus 250 ha kikuyu.\n\n\nhttp://agex.org.au/project/soil-acidity/\n\n\nSoil Type:\n\n\nJenny took one cropping paddock \"Middle of the Road\" as\nan example to test the calculators.\n\n\nIronstone to deep sands. Main soil issues are acidity,\n\n\nnon-wetting sandy soils and waterlogging.\n\n\nAverage rainfall:\n\n\n582 mm.\n\n\nSOUTH\n\n\nNatural Resources\n\n\nAUSTRA\n\n\nKangaroo Island\n\n\nGovernment\n\n\nof South Australia\n\n\nThe Cost of Not Treating Soil Acidity tool (Figure 1) shows\nthat without liming there will be a $1,545 loss in production\nover the next four years of crop; run that over 10 years and\nthe cost will be almost $3,900!\n\n\n\"We put 100 kg/ha of Calciprill down the tube to give a\nquick pH lift and boost calcium levels as legumes, especially\nbeans, love calcium just like grasses love nitrogen. The beans\nyielded well, averaging 2.5 t/ha, compared to an average\nof 1.8 t/ha pre-liming\" Jenny said. \"Since then I've noticed\na distinct difference to the paddock: feed stays greener for\nlonger once the soil profile starts to dry out, compared to\nadjacent paddocks that haven't been limed and the actual\nsoil structure has changed, becoming more friable\".\n\n\nThe Lime Sources Cost Comparison tool (Figure 2) was used\nto calculate the cost of moving the pH from 4.5 to 5.5 in\none lime application. This requires a lime application rate of\n4.3 t/ha at a total paddock cost of $2,260. In a nutshell a\nlime expenditure of approximately $2,300 will save $3,900\nin lost production or a 70% return on their money. Due to\nthe potential for inducing manganese deficiencies at this\nhigh lime application rate, the Stantons would probably\napply two applications of about 2.5 t/ha spread over the\n10 years.\n\n\nThe paddock results that really convinced the Stantons of\nthe benefits of liming was when they purchased a piece of\nL and N Deer's property \u2014 Caledonia.\n\n\n\"In the year before we purchased the property, Lloyd had\nlimed a paddock before sowing it to barley as barley is quite\nintolerant of low pH. As we cropped the whole property, we\nremoved all the internal fences and the old barley paddock\nplus surrounding paddocks were sown to wheat. The old\nbarley paddock stood out with an increase in yield of about\n20-30% more than the surrounding paddocks. In 2007\nthe paddock was tested and the old barley paddock was\npHCaC 5.8 and surrounding areas only 4.6. For many years,\nuntil the surrounding areas were limed, the yield difference\nwas extremely noticeable. This very visible response to lime\nmade us aware of the real economic benefits of maintaining\nour soil pH\" Will said.\n\n\nOnce paddocks have been limed the soil will still acidify. The\nMaintenance Liming Rate Calculator (Figure 3) calculates\nhow much lime is needed to maintain soil pH levels. It also\nhelps identify he key drivers of acidity. In this example, a key\ncause of low pH is the acidifying effects of nitrogen inputs,\neither via the legume component of the hay/cropping\nprogram or the application of nitrogenous fertilisers. On\naverage almost 0.5 t of lime is required per year just to\nbalance the acidification.\n\n\nThe Middle of the Road paddock was actually limed at 3t/ha\nin 2013 and the paddock was sown to beans. In hindsight,\nthe paddock should have been limed the year before to\nenable the limesand to fully react with the soil.\n\n\nOnce the Stantons had seen the results from the tools, they\nfelt a lot more comfortable about their budgeted annual\nlime expenditure.\n\n\n\"Basically we will recoup that expenditure within a year or\ntwo of liming\", Jenny said.\n\n\nCENTRAL WEST\n\n\nENTRAL WEST\n\n\nSpreading\n\n\n1828\n\n\nTAKE HOME MESSAGES\n\n\nUse the tools - they are easy to use and real eye opener on what acidity can be\ncosting you.\n\n\nSignificant yield improvements have been seen when cropped paddocks are limed.\nMake liming a part of your annual program.\n\n\nFigure 1: The cost of not treating soil acidity\n\n\nThis year\n\n\nCurrent pH value (CaCl2)\n\n\nNext year\n\n\nThree years time\n\n\nFour years time\n\n\nSelect crop type\n\n\nSelect crop type\n\n\nSelect crop type\nwheat (tol)\n\n\nSelect pH value\n\n\nSelect crop type\ncanola\n\n\nbeans\n\n\n4.5\n\n\nbeans\n\n\nEnter expected crop yield t/ha\n2\n\n\nEnter expected crop yield t/ha\n\n\nEnter expected crop yield t/ha\n\n\nEnter expected crop yield t/ha\n3.5\n\n\n2\n\n\n2\n\n\nEnter crop pasture value $/ha\n$625\n\n\nEnter crop pasture value $/ha\n$200\n\n\nEnter crop pasture value $/ha\n$400\n\n\nEnter crop pasture value $/ha\n$625\n\n\nEstimated loss of production\n\n\nEstimated loss of production\n0.7 t\n\n\nEstimated loss of production\n\n\nEstimated loss of production\n0.8 t\n\n\n0.8 t\n\n\n1 t\n\n\nEstimated value of lost production Estimated value of lost production Estimated value of lost production Estimated value of lost production\n$500\n$407\n\n\n$138\n\n\n$500\n\n\nCumulative loss/ha $1,545\n\n\nCumulative loss\n\n\n$ Value\n\n\nYield\n\n\n4\n\n\n\u0f72\u0f14 \u0f14 \u0f14 \u0f14 \u0f14 \u0f0b\n\n\nexpected\nactual\n\n\nexpected\n\n\nactual\n\n\n3\n\n\n2\n\n\nt/ha\n\n\n0\n\n\n0\n\n\nwheat (tol)\n\n\nwheat (tol)\n\n\ncanola\n\n\nbeans\n\n\nbeans\n\n\ncanola\n\n\nbeans\n\n\nbeans\n\n\nFigure 2: Lime sources cost comparison\n\n\nLime rate calculator\n\n\nHow\n\n\nPaddock middle froud\n\n\n\ub2c8 \ub2e4\n\n\nTarget Jme rat\u0186ng C & Adjusted\n\n\nZone Area\n\n\nha H (C&Cl, Texture +[CaCl,]\n\n\nLime rate\n\n\nApplication\n\n\nEnter\n\n\nSelect\n\n\nEnter\n\n\nEnter\n\n\nSelect\nsandy loam\n\n\nDG\n\n\nhigh\n\n\n1\n\n\n5.0\n\n\n5.5\n\n\n5.0\n\n\ninformation\n\n\n2\n\n\n00\n\n\n\u0395\n\n\nDO\n\n\n0.0\n\n\nLime Sources Comparison\n\n\nSource 1\n\n\nSource\n\n\nSelect\n\n\nNeutralizing\nValue (NV) or\n\n\nSANJAN\n\n\nNeutraliving\n\n\nTonnequined\n\n\n42 km\n\n\nom\n\n\nEither\n\n\n5/km/t\n\n\n$6.30 St\n\n\nSt\n\n\n517.50 5\n\n\nEnter\n\n\nSpreading cot\n\n\nOr\n\n\nContract Scot\n\n\nSit\n\n\nSit\n\n\n5\n\n\n\u5373\n\n\nTotal Sje\n\n\nProduct & cart\n\n\n575\n\n\nZone1\nJone\n\n\n50\n\n\nJone\n\n\n575\n\n\nPadlock\n\n\n190\n\n\nQuantity of lim\n\n\nFigure 3: Maintenance liming rate calculator\n\n\nTop Soil\n\n\nPaddock name:\n\n\nEnter\nMiddle of Road\n\n\nSelect\n\n\nLoamy Sand\n\n\nTexture:\n\n\nYour\n\n\n+\n\n\nH\n\n\n+\n\n\nH\n\n\nYear (20.....\n\n\n2016\n\n\n2019\n\n\n2014\n\n\n2010\n\n\n2015\n\n\n2012\n\n\nEnter\n\n\n2011\n\n\nAnnual Rainfall\n\n\n555\n\n\n854\n\n\n518\n\n\nEnter\n\n\n771\n\n\n625\n\n\nSaturated Soil\n\n\nTes\n\n\nYes\n\n\nSelect\n\n\nYes\n\n\nYes\n\n\nYes\n\n\nLeaching S\n\n\n35%\n\n\n85%\n\n\n85%\n\n\n85%\n\n\n85%\n\n\nProduct Removal\n\n\nClover hay\n\n\nOilseed\n\n\nGlower hay\n\n\nSelect\n\n\nGrain legume\n\n\nCrop Type\n\n\nCereal grain\n3.3\n\n\nCerea\n\n\nYield (tonnanha||\n\n\n54\n\n\n2.5\n\n\n19\n\n\nEnter\n\n\n2.1\n\n\nLime replacement/tonne grain yield (ba/ha)\n\n\n40\n\n\n40\n\n\n20\n\n\n9\n\n\n2\n\n\n9\n\n\n2\n\n\nLime required due to product removal [kg/ha)\n\n\nTotal (kg/ha)\n\n\n144\n\n\n258\n\n\n29.7\n\n\n46\n\n\n526.2\n\n\n4.2\n\n\n8.8\n\n\n40.5\n\n\nLegume Fixed Nitrogen\n\n\nkg N fixed per tonne legume production\n\n\n30\n\n\n60\n\n\n0\n\n\n0\n\n\nC\n\n\n0\n\n\n188\n\n\nLegume fixed nitrogen (kuha]\n\n\n109\n\n\n1935\n\n\n0\n\n\n0\n\n\n0\n\n\nG\n\n\nLeaching adjusted lime requirement (kg/ha/kg N\nLime required due to legume fixed N (kg/h)\n\n\n5.1\n\n\nG.O\n\n\n3.1\n\n\n0.0\n\n\n51\n\n\nG.D\n\n\nTotal (kg/h)\n\n\n4227.20\n\n\n550.4\n\n\n502.11\n\n\n0\n\n\n+424.7\n\n\nFertilisers inputs\n\n\nFertiliser 1. (Seeding)\n\n\nFertiliser Type\n\n\nSingle Super\n\n\n\u041e\u0414\u0420\n\n\nSelect\n\n\nMAP\n\n\nDIAP\n\n\nMAP\n\n\nMAP\n\n\n|Rate fertiliser (kg/ha)\n\n\nEnter\n\n\nGC\n\n\n40\n\n\n50\n\n\nBO\n\n\na\n\n\nProduct N (\n\n\n10.0\n\n\n10.0\n\n\n100\n\n\n18.0\n\n\nC.O\n\n\n18.0\n\n\nRate of nitrogen\n\n\n9\n\n\n0\n\n\nB\n\n\n8.6\n\n\n9\n\n\nLeaching adjusted lime requirement (ke/ha/kN\n\n\n0.0\n\n\n4.9\n\n\n6.7\n\n\n6.7\n\n\n6.7\n\n\n4.9\n\n\nLime required due to seeding fertiliser applications (kg/h)\n\n\n43.7\n\n\nTotal (kg/ha)\n\n\n166\n\n\n57.3\n\n\n0.0\n\n\n58.3\n\n\n43.7\n\n\n282.0\n\n\nFertiliser 2.\n\n\nFertiliser Type\n\n\nSelect\n\n\nUrea\n\n\nUreda\n\n\nUnea\n\n\nUrea\n\n\nPate fertilizer (kg/ha)\n\n\nEnter\n\n\n200\n\n\n150\n\n\nProduct N (\n\n\n0.0\n\n\n460\n\n\n46.0\n\n\n460\n\n\n0.0\n\n\n460\n\n\nRate of nitrogen\n\n\n345\n\n\n92\n\n\n0\n\n\n10\n\n\n69\n\n\n92\n\n\nLeaching adjusted lime requirement (kg/ha/kg N\nLime required dun to in-crop fortificer applications [kg/ha\n\n\n00\n\n\n5.1\n\n\n5.1\n\n\n5.1\n\n\n3.1\n\n\nTotal (kp/ha)\n\n\n105.6\n\n\n1815\n\n\n211.1\n\n\n0.0\n\n\n0.0\n\n\n1090.0\n\n\n281.5\n\n\nAnnual replacement lime required (kg/ha)\n\n\n355.0\n\n\n474.5\n\n\n982.3\n\n\n5216\n2598.3\n\n\n272.Z\n\n\nCumulative lime required over\"n\" years (kg/ha)\n\n\n474.5\n\n\n1456\n\n\nINTL\n\n\n26035\n\n\nSummary of lime requirement to address annual acidification\n\n\nInfluence of management practise on total lime required\n\n\nNumber of years of management data (n)\n\n\n7\n\n\nCumulative replacement required by management practise (kg/ha)\n\n\nLime required due to product\nremoval (kg/ha)\n\n\nLime required due to product removal (kg/ha)\n\n\n526.2\n\n\n34%\n\n\n9%\n\n\nLime required due to legume fixed N (kg/ha)\n\n\n1344.9\n\n\nLime required due to legume\nfixed N (kg/ha)\n\n\nLime required due to seeding fertiliser applications (kg/ha)\n\n\n282\n\n\nLime required due to in-crop fertiliser applications (kg/ha)\n\n\n1090.9\n\n\nLime required due to seeding\nfertiliser applications (kg/ha)\n\n\nTotal cumulative lime required over \"n\" years (kg/ha)\nAverage annual replacement lime required (kg/ha)\n\n\n3243.9\n\n\n16%\n\n\n41%\n\n\n463.4\n\n\nLime required due to in-crop\nfertiliser applications (kg/ha)\n\n\nRecommended lime application rate for 10 year period (t/ha)\n\n\n4.6\n\n\nLime maintenance rate\n\n\nFOR MORE INFORMATION\n\n\nAnnual replacement lime required (kg/ha)\n\n\nNatural Resources Kangaroo Island\n\n\nLime required due to legume fixed N (kg/ha)\n\n\nLime required (kg/ha)\n\n\n37 Dauncey Street Kingscote SA 5223\n\n\nP 08 8553 4444 E kinrc@sa.gov.au\n\n\nwww.naturalresources.sa.gov.au/kangarooisland\n\n\nThe project is supported by Natural Resources Kangaroo\nIsland through funding from the Australian Governments\nNational Landcare Program.\n\n\nNatural Resources\n\n\nNational\nLandcare\nProgramme\n\n\nKangaroo Island\n\n\n4\n\n\n2\n\n\n3\n\n\n5\n\n\n7\n\n\n6\n\n\nAustralian Government\n\n\nYear\n"}, "expected_output": {"claims": [{"unit": "kg/ha", "value": 463.4, "evidence": ["Once paddocks have been limed the soil will still acidify. The\nMaintenance Liming Rate Calculator (Figure 3) calculates\nhow much lime is needed to maintain soil pH levels. It also\nhelps identify he key drivers of acidity. In this example, a key\ncause of low pH is the acidifying effects of nitrogen inputs,\neither via the legume component of the hay/cropping\nprogram or the application of nitrogenous fertilisers. On\naverage almost 0.5 t of lime is required per year just to\nbalance the acidification.\n", "On average almost 0.5 t of lime is required per year just to balance the acidification.", "Total cumulative lime required over \"n\" years (kg/ha)\nAverage annual replacement lime required (kg/ha)", "463.4"]}]}, "metadata": {"product_category": "Chemical products", "request_id": "req_af88d019da0b9460"}}
{"id": "85c5fa9be53502c21e8ce464", "input": {"query": "What is the residual formaldehyde content (wt% or ppm) in Mannich-reaction tannin-based coagulant products?", "source_url": "https://oulurepo.oulu.fi/bitstream/handle/10024/31039/nbnfi-fe2021121761374.pdf?sequence=1&isAllowed=y", "document_text": "Industrial Crops & Products 176 (2022) 114336\n\n\nContents lists available at ScienceDirect\n\n\nINDUSTRIAL\nCROPS\nAND PRODUCTS\n\n\nIndustrial Crops & Products\n\n\nELSEVIER\n\n\njournal homepage: www.elsevier.com/locate/indcrop\n\n\nCheck for\nupdates\n\n\nSustainable tannin-based coagulants synthesized through Mannich reaction\nusing melamine as an amine source for water treatment applications\n\n\n*\nAdedayo Bello\u201d, Tiina Leivisk\u00e4\n\n\nUniversity of Oulu, Chemical Process Engineering, P.O. Box 4300, FI-90014 Oulu, Finland\n\n\nARTICLE INFO\n\n\nABSTRACT\n\n\nKeywords:\nMelamine\n\n\nThis study explored the potential of melamine, as the nitrogenating source for the cationization of pulverized\nquebracho and spruce tannins. The influence of modification conditions on the properties of the coagulants was\nstudied by varying the formaldehyde and melamine ratios with different activation times and temperatures.\nBased on considerations of charge density and shelf life, the most viable modifications were established as co-\nagulants synthesized with a 1:0.52 formaldehyde to melamine molar ratio at 70 \u00b0C. At optimal conditions, the\ncharge density of the quebracho and spruce coagulants was 2.22 meq/g and 1.04 meq/g, respectively, and the\nresidual formaldehyde content in the coagulants was low. The developed synthesis of this study demonstrated a\nclear advantage over previous methods due to the rapid modification step. X-ray photoelectron spectroscopy\n(XPS) confirmed the emergence of an amine which signifies a successful Mannich reaction in the coagulant.\nAlthough optimal modification conditions for the coagulants were established at 70 \u00b0C (5-min activation time),\nthe average molecular weight could not be determined for these conditions. Nevertheless, electrospray ionization\nmass spectrometry (ESI-MS) and matrix-assisted laser desorption/ionization time-of-flight (MALDI-TOF) mass\nspectrometry measurements revealed that the tannin-based coagulants obtained at lower temperatures (23 \u00b0C\nand 45 \u00b0C) possessed low average molecular weight (approx. 800-900 Da). Furthermore, ESI-MS and MALDI-\nTOF spectra showed that the Mannich modification resulted in the depolymerisation of the quebracho tannin,\nleading to a reduction in units of higher mass fractions in the synthesized coagulant, which contrasted with the\nspruce tannin. Jar test experiments with surface and industrial process waters demonstrated that the tannin\ncoagulants enhanced particle settling effectively.\n\n\nResidual formaldehyde\n\n\nMannich reaction\n\n\nBiocoagulants\nProcess waters\n\n\nmechanisms used for this process are mainly charge neutralization and\nsweep coagulation (Hussain et al., 2019; Zhu et al., 2011) while\nadsorption, patch coagulation and bridging flocculation are also\nemployed depending on the coagulant (Roussy et al., 2005; Leivisk\u00e4 and\nR\u00e4m\u00f6, 2008). For a chemical to be considered adequate for coagulation,\nit should be able to achieve a significant turbidity reduction within a\nshort period of time (Sincero and Gregoria, 2003). However, the effi-\nciency of the coagulation-flocculation is dictated by several factors,\nwhich include coagulant type and dosage, pH, concentration of organic\nmatter and dissolved solids, ionic strength and temperature of the\neffluent (Zhu et al., 2011; Sher et al., 2013).\nAluminium and iron salts are the most frequently used coagulants in\nwastewater treatment, and their efficiency in improving the aesthetic\ncondition of water and wastewaters are well reported in the literature\n(Alexander et al., 2012; Zhao et al., 2011). However, the major draw-\nbacks in the use of these salts include the generation of a large amount of\n\n\n1. Introduction\n\n\nWater is a key commodity for many industrial processes, and one\nmain concern is the remediation of industrial effluents, which are usu-\nally characterised by high turbidity. Waters with elevated turbidity\ninduced mostly by the presence of colloids and suspended solids cannot\nbe released directly into the environment without pre-treatment. This\nconcern has been further exacerbated by the changing and more strin-\ngent water quality standards that individual point source dischargers\nsuch as industrial facilities need to comply with. Coagulation-\nflocculation is considered the most efficient pre-treatment technique\nfor turbidity removal (Daifa et al., 2019; Teh et al., 2016).\nThe coagulation-flocculation process is initiated by the addition of\nchemicals to alter the physical state of pollutants, facilitating the\nagglomeration of the colloids and subsequent removal through a sedi-\nmentation or flotation process (Alexander et al., 2012). The removal\n\n\nCorresponding author.\n\n\nE-mail address: adedayo.bello@oulu.fi (A. Bello).\n\n\nhttps://doi.org/10.1016/j.indcrop.2021.114336\n\n\nReceived 23 August 2021; Received in revised form 22 November 2021; Accepted 25 November 2021\n\n\nAvailable online 9 December 2021\n\n\n0926-6690/\u00a9 2021 The Authors. Published by Elsevier B.V. This is an open access article under the CC BY license (http://creativecommons.org/licenses/by/4.0/).\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\ntoxic sludge, residual metal in treated waters, ineffectiveness over a\nwide range of pH values and poorer performance during the treatment of\ncold waters, detrimental effects to human health and the environment\n(Bratby, 2016; Haarhoff and Cleasby, 1988). Another widely used type\nof coagulant is synthetic polymers. In comparison to aluminium and iron\nsalts, synthetic polymers counteract some of the disadvantages associ-\nated with the use of inorganic salts, such as the generation of smaller\nvolumes of sludge and the absence of residual metals in sludge and\ntreated waters. Nevertheless, the applicability of these synthetic poly-\nmers has been marred by high cost, a low degree of biodegradability,\nand the fact that little is known of the toxicity of their monomers (Per-\neira et al., 2018; Harford et al., 2011). This has necessitated the need for\nan efficient and environmentally friendly coagulant to replace these\nconventional coagulants.\nOrganic coagulant derived from both animal- and plant-based re-\nsources provides a viable solution to the problems associated with fully\nsynthetic coagulants. Chitosan, which is obtained by the deacetylation\nof chitin extracted from crustacean shells, is an example of animal-based\ncoagulants whose impressive coagulative performance has been exten-\nsively reviewed (Lichtfouse et al., 2019; Yang et al., 2016). However, the\nlimited availability of chitin resources seems to have reduced their po-\ntential for replacing chemical coagulants. On the other hand,\nplant-based coagulants are largely non-hazardous, biodegradable and\nrenewable coagulants, which are naturally derived from various plant\ncomponents. In recent years, a group of plant-based coagulants that has\nattracted increasing interest is tannin-based coagulants (Arismendi\net al., 2018; Beltr\u00e1n-Heredia et al., 2010; S\u00e1nchez-Mart\u00edn et al., 2014;\nGraham et al., 2008; Grenda et al., 2020; Bello et al., 2020). This is\nbecause tannins are widely available in almost every part of a plant and\nare water-soluble, non-toxic polyphenols (Chung et al., 1998),\nTannins can be easily extracted with water and then cationized\nthrough the Mannich reaction to produce an effective coagulant. The\nMannich reaction is a process that involves the condensation of an\naldehyde with a primary or secondary amine to produce an iminium ion,\nwhich then introduces a positive charge into the polyphenolic matrix of\na tannin (Teh et al., 2016; Grenda et al., 2020). However, formaldehyde\nand the amines used during the cationization process are toxic. For\nexample, the International Agency for Research on Cancer classified\nformaldehyde as a human leukemogenic (IARC, 2012), and several\nstudies have found positive statistical significance between formalde-\nhyde exposure and different types of cancer (IARC, 2012; Hauptmann\net al., 2009; Zhang et al., 2010). Despite this, no study was found to have\nconducted a residual analysis of formaldehyde in tannin-based co-\nagulants. Furthermore, previous studies on the synthesis of tannin-based\ncoagulants have performed the Mannich reaction with the aid of amines\nsuch as diethanolamine (Arismendi et al., 2018; Beltr\u00e1n-Heredia et al.,\n2010; Bello et al., 2020), ethanolamine (Arismendi et al., 2018;\nBeltr\u00e1n-Heredia et al., 2010; Bello et al., 2020), dimethylamine hydro-\nchloric acid (Grenda et al., 2020) and ammonium chloride (Arismendi\net al., 2018; Beltr\u00e1n-Heredia et al., 2010). The amines mentioned above\nhave different toxicity levels with detrimental effects on human health\nand the environment and could only be employed in real applications\nunder stringent guidelines (Brooks, 2008; Poste et al., 2014). Melamine,\non the other hand, is a highly nitrogenous compound considered not to\nbe acutely toxic and to have no notable environmental effect (Rajpoot\net al., 2020; ECCC, 2016). However, due to the formation of urinary\nstones, which could lead to kidney failure, melamine is prohibited as a\ndirect additive in food products (FAO, 2019). Melamine reacts with\nformaldehyde to produce resins with impressive physical and chemical\nproperties and is used in many industrial and biotechnological appli-\ncations (Brettebauer and Schwarzinger, 2012). The reaction of mel-\namine and formaldehyde occurs mostly at high temperatures (>90 \u00b0C)\nand is achieved via a two-phase reaction mechanism of methylolation\nand condensation. During the reaction, up to six molecules of formal-\ndehyde can be attached to a melamine molecule to form two methylol\ncompounds in the initial methylolation phase, which is succeeded by a\n\n\ncondensation reaction to produce melamine-formaldehyde resins (Pizzi\nand Mittal, 2017). Some studies (Yubo et al., 2016; Kuo et al., 2019)\nhave reported the synthesis of resins with superior adhesive character-\nistics from a composite of tannin-melamine-formaldehyde compared to\nconventional melamine-formaldehyde resins. Also, Bucuroiu and\nco-workers (Bucuroiu et al., 2016) investigated the possibility of\napplying melamine-formaldehyde polymer directly as a coagulant. This\ncoagulant was claimed to be moderately effective in treating oily\nwastewaters, but limited information is available regarding its synthesis\nand performance. To the best of our knowledge, melamine has not been\nused in the Mannich modification of tannins into coagulants. The reac-\ntion of melamine with formaldehyde may provide additional benefits\nsuch as improved shelf-life of tannin coagulants and tight binding of\nformaldehyde.\nIn the present study, an investigation on the synthesis of sustainable\nMannich modification of tannin extracted from the bark of two tree\nspecies (quebracho and spruce) was conducted by replacing conven-\ntional toxic amines with melamine. The effects of modification param-\neters such as formaldehyde and amine concentrations, Mannich\nactivation time and temperature were evaluated with consideration to\nshelf life and charge density to optimize the modification. The residual\nformaldehyde content of the optimized products was analysed, and their\naverage molecular weights were estimated. A series of jar tests were\nconducted with effluents of different characteristics to determine the\nperformance of the biocoagulants in real applications. The jar test ex-\nperiments were evaluated using major water quality parameters, i.e.,\nturbidity, total surface charge (TSC), dissolved organic carbon (DOC)\nand specific ultra-violet absorbance (SUVA 254). Electrospray ionization\n(ESI), matrix-assisted laser desorption ionization time-of-flight (MALDI-\nTOF) and X-ray photoelectron spectroscopy (XPS) characterization were\nused to elucidate the changes induced by Mannich modification on the\npristine tannin samples.\n\n\n2. Materials and methods\n\n\n2.1. Tannins\n\n\nTwo tannin raw materials were aminomethylated through the\nMannich reaction. Spruce tannin (designated hereinafter as ST) was\nsupplied by VTT, the Technical Research Centre of Finland. The spruce\ntannin was obtained through hot water extraction from the bark of\nsummer harvested spruce trees (Picea abies) and was pulverized into a\npowdered tannin product through spray-drying. A detailed description\nof the spruce tannin extraction conditions and procedures has been re-\nported in previous studies (Bello et al., 2020; Kemppainen, 2015).\nQuebracho (Schinopsis balansae) tannin (QT), a commercial tannin\nproduct, was provided by Silvachimica Srl (Italy).\n\n\n2.2. Chemicals and reagents\n\n\nAll chemicals and reagents used during this study were of analytical\ngrade. Milli-Q ultrapure water was used throughout the experiments\nexcept when indicated. The melamine used as an amine source in cat-\nionization was manufactured by Sigma-Aldrich Chemical Corporation\n(USA), and formaldehyde (37% v/w) was supplied by VWR interna-\ntional (France). HCl (Merck KGaA) and NaOH (VWR Chemicals) were\nemployed for pH adjustment. The ferric sulphate (PIX-322) that was\nused as reference material in the coagulation experiments was provided\nby Kemira Oyj, Helsinki (Finland). BTG Instrument AB, (Sweden), sup-\nplied the sodium polyethylene sulphonate (PesNa) and poly-\ndiallyldimethylammonium chloride (PolyDadmac) solutions employed\nto measure the charge density of coagulants and surface charge of water\nsamples.\n\n\n2\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\n2.3. Water samples\n\n\ndensities of modified products were estimated in the equivalent of\ncharge per gram of solid tannin in the standardized volume of\ncoagulants.\nAll experiments were conducted in two repeats, and the most viable\nconditions for synthesis were evaluated based on the charge density and\nshelf life of the coagulants. To evaluate the shelf life, the tannin-based\ncoagulants were stored at room temperature and evaluated after three\nand eight months by visually checking their solubility (no formation of\nprecipitate) and measuring the charge density.\nFor comparison, melamine-formaldehyde coagulants (M-FA) were\nsynthesized with the same method used for tannin-based coagulant,\nexcept that tannin solution was replaced with MQ water. The M-FA\ncoagulant was used as a reference during the jar test experiment for\nkaolin/river water to study the advantage of tannin in the tannin-based\ncoagulant.\n\n\nOne natural water and two industrial process waters with different\ncharacteristics were selected to evaluate the applicability of the tannin-\nbased coagulant to remove turbidity. The natural water sample was\ncollected from the Oulu river (Oulu, Finland) and it had high humic\ncontent and low turbidity. The river water was then spiked with kaolin\nto obtain a kaolin/river water with relatively high turbidity (~40 NTU).\nThree grams of fine kaolin clay (0.063-0.5 mm) was dispersed in a batch\nof 121 of river water to produce a kaolin suspension of 250 mg/l. Other\nprocedures undertaken to obtain a stable kaolin dispersion in the river\nwater have been explained in detail in a previous publication (Bello\net al., 2019). The two process wastewaters (PW1 and PW2) were\nsampled from a metallurgical factory located in Finland. PW1, con-\ntaining oil and grease, was taken from the neutralization plant of a cold\nrolling mill before the flotation stage. PW2 was water from a gas\nscrubber obtained from a ferrochrome smelter. In contrast to the river\nwater, the industrial wastewaters were characterized by high turbidity.\nThe water quality of the water samples is presented in Table 1.\n\n\n2.5. Characterization\n\n\nBefore XPS analysis, all the samples were dried overnight in an oven\nat 40 \u00b0C to produce a solid product. The products were then ground to\nobtain particles of uniform size. XPS characterization was performed\nwith a Thermo Fisher Scientific ESCALAB 250xi equipped with a\nrotating monochromatic Al anode, generating an X-ray beam at 1486.6\neV. The XPS data analysis was performed with Avantage software, and\nall spectra were fitted using the Shirley-type background. Charge\ncorrection was performed on the binding energy (BE) with the\nassumption that the binding energy of adventitious carbon is 284.8 eV.\nX-ray diffraction (XRD) characterization was performed with a\nRigaku SmartLab diffractometer equipped with a Co rotating anode\nlamp (40 kV, 135 mA). The speed of acquisition was four degrees per\nminute at 0.02 degrees per step. The charge density of the tannin co-\nagulants was determined via titration with PesNa using a M\u00fctek particle\ncharge detector PCD 03 pH (Hersching, Germany). The average of three\ntitrations was calculated. The pH measurement of the modified tannin\nwas performed using a VWR pHenomenal\u00ae 1100 L instrument.\nThe weight-average molecular weight of the tannin samples and\ntannin-based coagulants were estimated with the following mass spec-\ntrometry methods: ESI and MALDI-TOF. Prior to molecule weight\nmeasurement by MALDI-TOF and ESI, the samples were dissolved in\n50% acetonitrile (ACN)/0.1% formic acid at 1000-fold dilution. For the\nESI method, the analytes were infused at 10 \u03bcl/min into a Q-extractive\nplus orbitrap mass spectrometer equipped with a standard ESI source.\nThe orbitrap mass spectrometer was calibrated with a Thermo mass\nstandard dedicated for Orbi-trap instruments. Mass spectra were recor-\nded in positive mode using the standard setting for source parameters\nand data acquisition over a mass range of 160-4000 m/z. The molecular\naverage weight of the tannin fractions from the ESI spectra was calcu-\nlated with the classical polymer parameters presented in Eq. (1).\n\n\n2.4. Mannich modification\n\n\nTannin-based coagulants were synthesized through the amino-\nmethylation of tannin using formaldehyde (FA) and melamine as the\naldehyde and amine sources. The coagulant modification was achieved\nthrough a dual-step system. The first step involved the preparation of\ntannin solution and Mannich solution. For the tannin solution, 2.5 g of\npulverized tannin was dissolved in 10 ml of MQ water at ambient tem-\nperature. The tannin was gradually added in batches, and the complete\ndissolution of tannin in MQ water was ensured by the continuous mixing\nof the solution with a laboratory spoon. In the Mannich solution (MS),\nthe melamine was allowed to react with formaldehyde at different\ndosages to generate the imine ion. To ensure an optimal reaction in MS,\nthe mixture was kept at 120 \u00b0C for 2 h. After that, the MS was acidified\nby injecting 1 ml of HCl (37%) into the solution, which reduced the pH\nto ~1. Next, the tannin solution and the MS were mixed and reacted over\nspecific periods and temperatures under continuous agitation. The ob-\ntained product was standardized by transferring it to a 50 ml volumetric\nflask and filled up to the mark with MQ water. The solid contents of the\ncoagulant, before and after standardization were ~10% (w/v) and 5%\n(w/v), respectively. The modified products were within a pH range of\n1.7\u00b10.2 after standardization. The dosage of tannin coagulants was\ndetermined as the active coagulant content of the standardized product.\nChemical dosages and experimental conditions were selected ac-\ncording to preliminary trials, which provided information about the\nconditions under which undesirable instantaneous gelation of QT co-\nagulants (T > 45 \u00b0C and high FA to melamine ratio) or the formation of\nwhite crystals (low FA to melamine ratio) occurred. The effects of\nchemical dosages and experimental conditions on the properties of the\ntannin-based coagulants were then studied in two parts. First, the effect\nof activation times (1, 5 and 20 min) was studied with varying melamine\ndosages (9.0, 10.4 and 11.4 mmol) while the FA dosage (26.9 mmol) and\ntemperature (45 \u00b0C) were kept constant. This gave FA to melamine\nmolar ratios of 1:0.33, 1:0.39 and 1:0.42, respectively. Next, the effect of\nactivation temperature (23, 45 and 70 \u00b0C) with different FA dosages\n(20.1, 26.9 and 40.4 mmol) was studied with a fixed melamine dosage\n(10.4 mmol) and using an activation time of five minutes. The FA to\nmelamine molar ratios were 1:0.52, 1:0.39, and 1:0.26, and charge\n\n\n\u03a3(m/z) 11\n\u03a3(m/z);11\n\n\nMw =\n\n\n(1)\n\n\nm/z is the mass to charge of the oligomer.\nIt is the intensity of the oligomer.\n\n\nThe MALDI-TOF spectra were measured with a Bruker Ultra-\nflextreme instrument in reflectron mode with protonation in the mass\nrange of 500-6000 m/z. MALDI was externally calibrated before the\nmeasurement with Bruker's peptide calibration standard II. The MALDI-\n\n\nTablee 1\n\n\nCharacteristics of the kaolin/river water and process waters.\n\n\nDOC (mg/l)\n\n\nWater sample\n\n\nTSC (\u03bceq/1)\n\n\nSUVA (l/mg..m)\n3.284\n\n\nTurbidity (NTU)\n\n\nUV 254 (1/cm)\n\n\npH\n\n\nConductivity (\u00b5S/cm)\n\n\nKaolin/river water\n\n\n36\n\n\n7.5\n\n\n43\n\n\n0.312\n\n\n9.5\n\n\n-45\n-18.6\n\n\n0.181\n\n\nPW1\n\n\n50\n152\n\n\n6.3\n\n\n448\n\n\n11\n\n\n1.645\n0.904\n\n\n0.235\n\n\nPW2\n\n\n8.4\n\n\n1735\n\n\n26\n\n\n-11.5\n\n\n3\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nTOF spectra of tannin samples and tannin coagulants were run with the\nsame volume of matrix solution (50% v/v) while the matrix solution was\nprepared by dissolving 2,5-dihydroxy benzoic acid (DHB) in 50% ACN/\nwater to a concentration of 10 mg/ml. Subsequently, the matrix and\nanalyte were subjected to brief ultrasonic mixing, and 1 \u00b5l aliquots were\nplaced on a steel sample stage to air dry. After drying, the steel sample\nstage was introduced into the spectrometer. The weight-average mo-\nlecular weight of the tannin fraction in the spectra produced from\nMALDI-TOF was computed using Eq. (2).\n\n\n3. Results and discussion\n\n\n3.1. Effect of activation time and melamine dosage\n\n\nThe effect of activation time (1-20 min) and melamine dosage\n(9.0-11.4 mmol) on charge density when the FA dosage (26.9 mmol) and\nactivation temperature (45 \u00b0C) were kept constant is shown in Fig. 1. A\nvisible charge density progression could be observed as the activation\ntime and melamine dosage increased in most QT modifications (Fig. 1a).\nQT coagulants having activation times of 1 min and 20 min resulted in a\nsignificant increase in charge density when the melamine dosage was\nincreased. On the other hand, the charge densities of the QT coagulants\nhaving an activation time of 5 min deviated from the trend, and a logical\nreason for this deviation could not be established. When comparing the\neffect of activation time on recorded charge densities, there was only a\nminimal gain in charge density with 10.4 mmol melamine dosage within\nthe studied timeframe. Coagulants modified with 9.0 and 11.4 mmol\nmelamine dosages recorded a perceptible 25% (0.24 meq/g) and 8%\n(0.14 meq/g) increase in charge densities, respectively, when the acti-\nvation time was increased from 1 to 20 min. Nonetheless, the activation\ntime increase was confirmed not to be statistically significant with all\nmelamine dosages (t-test, p < 0.05, two-tailed).\nThe ST coagulant graph displayed in Fig. 1b indicates that there was\nno notable correlation between the charge density and melamine dosage,\nas coagulant modification with a higher melamine dosage did not result in\na significantly higher charge density. The ST coagulants modified with\nshorter activation times (1 and 5 min) clearly showed higher charge\ndensity values at higher melamine dosages as compared to the lowest\nmelamine dosage, while this was not the case with the 20 min activation\ntime. It is noteworthy that the most significant increase in charge density\ndue to melamine dosage variation was recorded when the dosage was\nraised from 9.0 to 10.4 mmol (0.29 \u00b1 0.17-0.48 \u00b1 0.18 meq/g) when\ncompared with 11.4 mmol (0.41 \u00b1 0.14 meq/g). Interestingly, the charge\ndensities of the modified coagulants increased progressively with the\nactivation time for coagulants modified with 9.0-10.4 mmol melamine\ndosages. Charge densities increased from an average of 0.21 \u00b1 0.11 meq/\ng to 0.41 \u00b1 0.17 meq/g when activation time was increased from 1 min\nto 5 min, and further increased to 0.55 \u00b1 0.11 meq/g when the activation\ntime was raised to 20 min.\n\n\n\u03a3N;M\u00b2\n\u03a3\u039d \u039c\n\n\nMw\n\n\n(2)\n\n\n=\n\n\nMi is the mass of the oligomer.\n\n\nN\u0118 is the number of molecules of a specific molecular weight.\nThe residual formaldehyde of the tannin coagulants was analysed\nusing a high-performance liquid chromatograph with diode-array\ndetection (HPLC-DAD) from Waters (Autosampler: 717 Plus; Quater-\nnary pump: 600 Controller; 2998 Detector). The samples were diluted in\nwater and mixed with 2,4-dinitrophenylhydrazine, whereby aldehydes\nand ketones react to give the respective hydrazones. These were sepa-\nrated through a reversed-phase HPLC method, detected at 355 nm and\nquantified.\n\n\n2.6. Water analysis\n\n\nThe turbidity of water samples was measured with a Hach 2100Q\nportable turbidity metre. The total surface charge of water samples was\nmeasured with a M\u00fctek particle charge detector (PCD 03) by titrating\n10 ml of water samples with PesNa or Poly-Dadmac titrant (0.001 N). A\nMetrohm 744-pH metre equipped with an epoxy electrode and a Mettler\nToledo conductivity metre were employed to measure the pH and con-\nductivity of the water samples. The ultraviolent absorbance (UVA) of the\nwater samples was measured at 254 nm with a Shimadzu ultraviolet\nspectrometer (model UV-1800), while dissolved organic carbon (DOC)\nmeasurement was performed with a SIEVERS 900 portable TOC analy-\nser. The DOC and UVA measurement was performed after filtration with\n0.2-\u03bcm filtered (VWR, polyethersulphone membrane) water samples,\nand specific ultraviolet absorbance (SUVA) values were calculated from\nthe UVA and DOC values in accordance with EPA methods (EPA, 2009).\nA detailed analysis of water samples was performed in accordance with\nstandardized methods: NH4-N, NO2-N, NO3-N and PO4-P were measured\nwith a continuous flow analyser using the SFS-EN ISO 11732:2005, SFS-\nEN ISO 13395:1997 and SFS-EN ISO 15681-2:2005 methods, respec-\ntively. Cl, F and SO4\u00b2 were analysed using ion chromatography (SFS-EN\nISO 10304-1:2009), and the concentration of the elements in the water\nsamples was analysed using an inductively coupled plasma optical\nemission spectrometer (ICP-OES) (SFS-EN ISO 17294-2:2016).\n\n\nOn a general note, coagulants with relatively stable high charge\ndensities were obtained in both tannin modifications when the mel-\namine dosage was kept at 10.4 mmol. When considering the charge\ndensities obtained from the shortest activation time (1 min), this study\nshowed a rapid crosslinking between the flavonoid (tannin) units and\niminium ions. A short activation timeframe was chosen for this study\nbecause a prolonged activation time resulted in the gelation of QT co-\nagulants modified with 9.0 mmol of melamine at 45 \u00b0C. However,\nearlier researchers of this topic (Arismendi et al., 2018; Beltr\u00e1n-Heredia\net al., 2010; Grenda et al., 2020) have ascertained that crosslinking\nprogresses over time and that a longer time is required to obtain a\ncomplete crosslinking of the two constituents (tannin and MS). A direct\ncomparison is difficult because different amines were used during\nMannich modification in previous studies. Grenda et al. (2020) pro-\nduced tannin coagulants from Acacia mearnsii and quebracho with\ndimethylamine as the amine source and reported a 100 and 180-minutes\nactivation times for an optimal coagulant modification. In a separate\nstudy, Grenda et al. (2018) also recommended a reaction time of 90 min\nfor Acacia mearnsii based on the products shelf life. In their studies,\nBeltr\u00e1n-Heredia et al. (2010), synthesized a tannin-based coagulant\nfrom quebracho and Acacia mearnsii with diethanolamine, ethanolamine\nand ammonium chloride and suggested a 24-hour timeframe for optimal\nmodification. The synthesis developed in this study has a clear advan-\ntage over previous methods due to the rapid modification step. How-\never, it is essential to state that all QT coagulants synthesized during this\nphase of the study gelated after one month of storage, whereas the ST\ncoagulants retained their liquid state.\n\n\n2.7. Coagulation experiments\n\n\nCoagulation experiments were performed in jar tests using a Kemira\nflocculator 2000 jar test apparatus equipped with one-litre glass bea-\nkers. For the jar tests, the beakers were filled to the 800 ml mark with\nwater samples and dosed with different dosages of coagulants. The so-\nlution was then subjected to rapid mixing (150 rpm) for one minute,\nslow mixing (40 rpm) for 20 min and sedimentation for 30 min. The\nsedimentation period was followed by the extraction of 200 ml of the\nsupernatant from 3 cm below the surface of the test water sample for\nvarious analyses. Jar test residues were recovered by first decanting the\nexcess supernatant, centrifuging the flocs at 2500 rpm for 10 min and\ndecanting the supernatant again. The flocs were then oven-dried over-\nnight at 40 \u00b0C to produce jar test residues.\n\n\n4\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nST Coagulant\n\n\na)\n\n\nQT Coagulant\n\n\nb)\n\n\n2.0\n\n\n2.0\n\n\nMelamine 9.0 mmol\nMelamine 10.4 mmol\n\n\nCharge density (meq/g)\n\n\nMelamine 11.4 mmol\n\n\n1.5\n\n\n1.5\n\n\n1.0\n\n\n1.0\n\n\n0.5\n\n\n0.5\n\n\n0.0\n\n\n-0.0\n\n\n5\nTime (min)\n\n\n20\n\n\n20\n\n\n1\n\n\n5\nTime (min)\n\n\nFig. 1. Effects of activation time and melamine dosage on charge densities of a) QT and b) ST coagulants. Range of charge densities represents the deviation in two\nrepeats: temperature 45 \u00b0C and FA dosage 26.9 mmol. Charge densities were measured immediately after Mannich modification.\n\n\n3.2. Effect of activation temperature and formaldehyde dosage\n\n\ntemperatures (23, 45 and 70 \u00b0C) and formaldehyde dosages (20.1, 26.9\nand 40.4 mmol) were investigated with a constant melamine dosage of\n10.4 mmol and a five-minute activation time. The result for the QT co-\nagulants, displayed in Fig. 2a, indicates that the activation temperature\nhad a significant effect only when the lowest FA dosage was used, as the\ncharge density increased from 0.88 to 1.78 meq/g with increasing\ntemperature. With higher FA dosages, the activation temperature did\n\n\nAs described in section 3.1, coagulants with considerably high\ncharge densities can be obtained with short activation times. Besides,\nsome preliminary experiments in this study had shown that QT Co-\nagulants were quite unstable when modified at high temperature and\nwhen using a longer activation time. Next, the effects of the activation\n\n\nST Coagulant\n\n\nQT Coagulant\n\n\na)\n\n\nb)\n\n\n2.0\n\n\n2.0\n\n\nFA 20.1 mmol\n\n\nFA 26.9 mmol\n\n\nCharge density (meq/g)\n\n\nFA 40.4 mmol\n\n\n1.5\n\n\n1.5\n\n\n10\n\n\n1.0\n\n\n0.5\n\n\n0.5\n\n\n0.0\n\n\n-0.0\n\n\n23\n\n\n70\n\n\n23\n\n\n70\n\n\n45\nTemperature (\u00b0C)\n\n\n45\nTemperature (\u00b0C)\n\n\nFig. 2. Effects of temperature and FA-melamine dosages on charge densities of a) QT and b) ST coagulants. Range of charge densities represents the deviation in two\nrepeats: activation time (5 min) and melamine dosage (10.4 mmol). Charge densities were measured immediately after Mannich modification.\n\n\n5\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nnot affect the charge density values. The QT coagulant resulted in a high\ncharge density even when modified at room temperature and using a\nmedium FA dosage of 26.9 mmol. The reduction in charge density with\nthe highest FA dosage can be attributed to a reduced iminium ion for-\nmation due to an excess of formaldehyde in the reaction. Excess form-\naldehyde to amine has been reported to inhibit the formation of\nhemiaminals: an intermediate compound produced before iminium ions\nare formed during Mannich reactions (Iwasawa et al., 2007; Short et al.,\n1992). Earlier studies on the synthesis of tannin-based coagulants have\nnoted that a higher formaldehyde dosage is detrimental to the\nphysico-chemical properties of the biocoagulants, as it results in the\ninstant gelation of products or coagulants with a relatively low shelf life\n(Ibrahim et al., 2021). Also, a physical observation of the QT coagulants\nmodified with the highest formaldehyde dosage (40.4 mmol) revealed\nthat they underwent a colour change from a homogeneous dark brown\nsolution to light brown products characterized by precipitates, during\nmodification at 70 \u00b0C.\nIn contrast to the QT coagulants, the charge densities of the ST co-\nagulants increased with an increase in temperature in most cases\n(Fig. 2b), except for the coagulant modified with 20.1 mmol of FA at\n45 \u00b0C. Also, increasing the formaldehyde dosage did not enhance the ST\ncoagulant charge densities. High temperature might have distinctively\nfavoured the crosslinking process in the spruce tannin compared to the\nQT tannin. Nevertheless, the range of charge densities obtained from all\nthe individual modifications for the ST coagulants (0.32\u20130.90 meq/g)\nwas still lower than that of the QT coagulants (0.82\u20131.95 meq/g). This\nwas probably related to the presence of fewer phenolic functional groups\nin the spruce tannin to accept the iminium ion generated in the Mannich\nsolution (Kemppainen, 2015). Previous studies (Bello et al., 2020) have\ncorroborated this claim and further elaborated that the content of\nproanthocyanidins (condensed tannin) is over 300% higher in\nquebracho tannin than in its spruce counterpart. Another factor\nconsidered to have influenced the performance of spruce tannin is the\npresence of impurities such as ash and glucosides. These impurities\nusually lead to the formation of undesired and competitive by-products\nduring aminomethylation (Arismendi et al., 2018).\n\n\n3.3. Evaluation of shelf life of tannin-based coagulants\n\n\nA viable coagulant in water treatment applications should have a\nreasonable shelf life. To evaluate the stability of the Mannich-modified\ncoagulants, they were stored at room temperature for three months.\nAfter the three-month storage period, it was discovered that all the QT\ncoagulants modified at the higher formaldehyde dosage (\u226526.9 mmol)\nhad gelated. Surprisingly, gelation was not recorded with any ST bio-\ncoagulants during the reviewed timeframe. The reason for the selective\ncoagulant gelation could not be validated during this study. However, it\nis widely known that one of the reaction mechanisms behind the ami-\nnomethylation of tannin-based polymers is condensation\n(S\u00e1nchez-Mart\u00edn et al., 2014; Grenda et al., 2020). This infers that\ncrosslinking progresses over time, which leads to continuous thickening\nand eventual gelation of the modified products (Grenda et al., 2020).\nFurther studies were performed only on both tannin coagulants\nmodified with 10.4 mmol melamine dosage and 20.1 mmol FA dosage\n(FA to melamine molar ratio 1:0.52). The variations in charge densities\nobtained for the coagulants after the three- and eight-month storage\nperiods are shown in Fig. 3 and Table 2. Both tannin coagulants wit-\nnessed an average increase of about 45% in charge densities after the\nthree-month storage period. However, it is important to note that the\ncharge density variations were more pronounced in the QT samples. A\ncomparison between charge densities obtained from the samples after\nthe three- and eight-month storage periods reveals that the crosslinking\nreaction had been completed within three months, resulting in products\nwith more stable charge densities. Interestingly, coagulants with rela-\ntively stable charge densities could be obtained by employing higher\ntemperature during the activation phase. Thus, higher temperature was\nconcluded to have accelerated the crosslinking process. It is essential to\nnote that the liquid-equivalent charge densities (meq/1) of QT co-\nagulants were close to those of M-FA coagulants modified with exactly\nthe same melamine and formaldehyde dosages and subjected to similar\nmodification conditions (Table S1).\nFormaldehyde is categorized as a carcinogen by the International\nAgency for Research on Cancer (IARC, 2012), and thus a lower amount\n\n\nQT Coagulant\n\n\na)\n\n\nST Coagulant\n\n\nb)\n\n\n3.0\n\n\n3.0\n\n\n23\u00b0C\n\n\n45\u00b0C\n\n\nCharge density (meq/g)\n\n\n2.5\n\n\n70\u00b0C\n\n\n2.0\n\n\n2.0\n\n\n1.5\n\n\n1.5\n\n\n1.0\n\n\n1.0\n\n\n0.5\n\n\n0.5\n\n\n0.0\n\n\n-0.0\n\n\n3\nTime (month)\n\n\n0\n\n\n8\n\n\n3\n\n\nTime (month)\n\n\nFig. 3. Changes in charge densities of a) QT coagulants and b) ST coagulants monitored over time (0, 3 and 8 months) as a function of reaction temperature (23 \u00b0C,\n45 \u00b0C and 70 \u00b0C). Range of charge densities represents the deviation in two repeats (two modifications) (FA 20.1 mmol, melamine 10.4 mmol, activation time 5 min).\n\n\n6\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nin some fishes (European Food and safety Authority, 2014) to\n406 mg/kg in dried mushrooms (Yeh et al., 2013). An estimation of the\nconcentration of FA content in a treated water sample, if 100 mg/1\ndosages of coagulants are applied and assuming all residual FA stays in\nthe treated water, shows a significantly low residual FA of 1.8 mg/1\n(0.00018 wt%) and 4.2 mg/1 (0.00042 wt%) for the QT and ST co-\nagulants modified at 70 \u00b0C, respectively.\n\n\nTable 2\n\n\nResidual formaldehyde, charge densities, and molecular weights of coagulants\nafter 8 months' storage. Activation time 5 min, FA 20.1 mmol, melamine\n10.4 mmol.\n\n\nCoagulants\n\n\nFormaldehyde\ncontent (wt% FA/v\n(coagulant solution))\n\n\nMolecular\nweight (Da)\n\n\nCharge density (meq/g\nof solid tannin in\n\n\ncoagulant solution)\n\n\nMALDI ESI\n\n\nQT 23 \u00b0C\n\n\n2.133 0.005 (pH\n4.4)\n2.174 0.008 (pH\n\n\n916.4\n\n\n928.7\n\n\n0.08\n\n\n3.4. XPS analysis of tannin samples and tannin-based coagulants\n\n\nQT 45 \u00b0C\n\n\n880.1\n\n\n920.5\n\n\n0.09\n\n\n4.4)\n\n\nXPS analysis was employed to elucidate the elemental change\ninduced by the Mannich reaction on the tannin samples. XPS revealed\nthat the tannin samples (QT and ST) consisted mainly of carbon and\noxygen peaks located at binding energy ~285 eV (C1s) and ~533 eV\n(01s) in the wide spectra (Fig. S1). Additionally, the C1s peak could be\nfitted into two and three carbon components for the quebracho and\nspruce tannin, respectively. These were in agreement with a previous\nstudy (Bello et al., 2020). The first two carbon components at 284.8 eV\nand 286.4 present in both tannin samples were assigned to\nC-C/C=C/C-H (C1 component) and C-OH/C-O (C2 component) bonds\n(Bello et al., 2020; Kl\u00e9bert et al., 2017). The third peak at 288.1 eV\ndetected only in the spruce tannin was assigned to the C=O bond (C3\ncomponent) (Awada et al., 2012). The C=O bond might be attributed to\npolymeric sugars in the spruce tannin sample (Bello et al., 2020;\nKemppainen, 2015).\nFig. 4a and b show the XPS high-resolution Cls and N1s spectra for\nthe pristine melamine sample. The two peaks in the Cls spectrum\nlocated at 284.8 eV and 287.8 eV were assigned to C-C/C-H (C1\ncomponent) and C-N/ C=N (C2 component) bonds (Yao et al., 2005).\nThe C1 component is formed due to surface contamination by adventi-\ntious carbon (Dementjev et al., 2000). The C-N bonds in the melamine\nare located on the low BE side of the C2 component and the C=N bonds\n\n\nQT 70 \u00b0C\n\n\n2.2170.004 (pH\n\n\n0.09\n\n\nN/A\n\n\nN/A\n\n\n4.5)\n\n\n1.227 0.002 (pH\n\n\nST 23 \u00b0C\n\n\n0.14\n\n\n736.4\n\n\n918.5\n\n\n4.5)\n\n\n1.253 0.001 (pH\n4.5)\n1.040 0.002 (pH\n4.5)\n\n\nST 45 \u00b0C\n\n\n0.13\n\n\n830.2\n\n\n858.6\n\n\nST 70 \u00b0C\n\n\n0.21\n\n\nN/A\n\n\nN/A\n\n\nN/A: Not available (The molecular weight of coagulants modified at 70 \u00b0C was\ndisregarded because ESI and MALDI-TOF consistently produced spectra with\nlow intensities).\n\n\nof residual formaldehyde is preferred in the modified products. The\namount of residual formaldehyde in the coagulants after eight months is\npresented in Table 2. The result indicates a low percentage of residual\nformaldehyde in both tannin coagulants. Nonetheless, the formaldehyde\namount was slightly lower in the QT coagulants than that in the ST\ncoagulants. The residual FA amounts were low when considering the\nbackground level of formaldehyde in many food items. Formaldehyde\nexists in the environment due to natural processes and anthropogenic\nactivities (IPCS, 1989), and a reasonable amount of formaldehyde is\nfound as a natural constituent in food products ranging from 200 mg/kg\n\n\na)\n\n\nb)\n\n\nC2\n\n\nN1\n5.\n\n\nN2\n\n\nC1\n\n\n396\n\n\n402\n\n\n400\n\n\n300\n\n\n295\n\n\n290\n\n\n285\n\n\n280\n\n\n404\n\n\n398\n\n\nBE (eV)\n\n\nBE (eV)\n\n\nFig. 4. High resolution a) Cls and b) N1s XPS spectra of melamine.\n\n\n7\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\non the high BE side. In the N1s spectrum, the lower energy component\n(N1), located at 398.0 eV, was ascribed to C=N-C (imine), while the N2\ncomponent positioned at 399.0 eV was attributed to C-NH2, a type of\nnitrogen bond (amine) (Lee et al., 2016).\nThe XPS survey for the commercial ferric sulphate used as a reference\ncoagulant confirmed the presence of carbon, oxygen, iron and sulphur\n(Fig. S2). In the case of the Mannich-modified tannin coagulants, addi-\ntional peaks for Cl and N (besides the peaks observed in tannin spectra)\nwere observed in the survey spectra at binding energy 198 eV (Cl2p) and\n~400 eV (N1s) (Fig. S2). The Cl2p was believed to have originated from\nthe HCl used for acidification, and the N1s peak was related to mel-\namine. The Cls spectra of the QT and ST coagulant samples (Fig. 5a and\nb) were fitted to three and four carbon peaks, respectively. Both tannin-\nbased coagulants have peaks at 284.8 eV and 286.4 eV, which is\nconsistent with the peaks of C1 and C2 components identified in the\ntannin samples. The third peak for the QT and ST coagulants was located\nat 288.3 eV and 287.5 eV, respectively. The ST coagulant was observed\nto possess an additional peak at 288.8 eV. The emergence of the peak at\n287.5 eV (C3 component) in the ST coagulant is related to the C-N and\nC=N bonds of melamine (Son et al., 2020) associated with a successful\nMannich modification (Sahoo et al., 2006). The fourth peak in the ST\ncoagulant located at 288.8 eV (C4 component) can most likely be\nattributed to the emergence of a new carbon bond associated with the\ncarboxyl groups (O=C-OH, O=C-O) (Awada et al., 2012; Feng et al.,\n2017), which might indicate the formation of D-glucosamine derivatives\n(Sharma and Peddinti, 2018; Swilem et al., 2017). This peak further\ncorroborates the formation of competitive by-products during spruce\ntannin aminomethylation, hence reducing the regioselectivity of Man-\nnich modification, as suggested in previous research studies (Arismendi\n\n\net al., 2018; Bello et al., 2020). In the QT coagulant, the third broader\npeak (C3/C4 component) on the low BE side included the C-N and C=N\nbonds of melamine and possibly carboxyl groups (0=C-0, O=C-OH)\non the high BE side. These O=C-O/O=C-OH bonds could not be\njustified in the QT coagulant based on the Cls spectrum of the pristine\nquebracho tannin (Fig. S3). However, Kemppainen et al. (2014) char-\nacterized the same commercial quebracho tannin with \u00b3\u00b9p nuclear\nmagnetic resonance and reported a 0.22 mmol/g carboxylic content,\nindicating the presence of some impurities.\nFig. 5c and d illustrate the chemical state of nitrogen in the modified\ncoagulants. The weight percentage of nitrogen in the samples after\nMannich modification was 19.0% and 16.4% for the QT and ST co-\nagulants, respectively (Table S3). The N1s high-resolution spectrum of\nboth tannin modifications were fitted to three peaks, representing imine\n(C=N-C in melamine structure, N1 component), deprotonated amine\n(N2) and protonated amine (N3) components. For the QT coagulant, the\nimine, deprotonated, and protonated amine peaks were located at\n398.5 eV (N1), 399.5 eV (N2) and 400.3 eV (N3) binding energies,\nrespectively, while those of the ST coagulant were located on almost the\nsame binding energies (398.4 eV (N1), 399.4 eV (N2) and 400.4 eV\n(N3)). The presence of the nitrogen peak signifies the success of the\nMannich reaction.\n\n\n3.5. Molecular weight of tannin samples and modified coagulants\n\n\nThe positive mode ES1 spectra of the pristine tannin samples and the\nMannich-modified coagulants are shown in Fig. 6. The quebracho tannin\n(Fig. 6a) possessed polyphenols of higher mass fractions than the spruce\ntannin (Fig. 6b). Furthermore, the data indicates that fragmented or\n\n\nb)\n\n\na)\n\n\nC1\n\n\nC1\n\n\nC2\n8\u2192\n\n\n-3\n\n\nC3/C4\n\n\n295\n\n\n300\n\n\n290\n\n\n285\n\n\n280\n\n\n300\n\n\n295\n\n\n290\n\n\n285\n\n\n280\n\n\nBE (eV)\n\n\nBE (eV)\n\n\nc)\n\n\nd)\n\n\nA\n\n\nN1\n\n\nN2\n\n\nN2\n\n\nN1\n\n\nN3\n\n\nN3\n\n\nT\n396\n\n\n404\n\n\n402\n\n\n400\n\n\n398\n\n\n404\n\n\n402\n\n\n400\nBE (eV)\n\n\n398\n\n\n396\n\n\nBE (eV)\n\n\nFig. 5. High resolution C1s XPS spectrum for a) QT and b) ST coagulants and N1s XPS spectrum for c) QT and d) ST coagulants (FA 20.1 mmol, melamine 10.4 mmol,\nactivation time 5 min, temperature 70 \u00b0C).\n\n\n8\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nsimple phenolic compounds dominate the polyphenol composition of\nthe spruce tannin. The quebracho tannin produced a spectrum with a\nseries of procyanidin peaks with a mass difference of 288 Da apart\n(1108-1396 m/z) and an oligomer series with a mass of repeated units of\n272 Da (563-835 and 683-956 m/z). The pattern of these oligomeric\nunits was in agreement with the literature for hot water extracted\nquebracho tannin (Venter et al., 2012; Pash et al., 2001). The quebracho\ntannin fractions were more widely distributed within a 350-2000 m/z\nmass range (Fig. 6a). The average molecular weight of the quebracho\ntannin was estimated as 1389 Da (by ESI). The fractions of the spruce\ntannin were narrowly distributed within the mass range of 150-900 m/z\nand possessed an average molecular weight of 612 Da (Fig. 6b).\nThe mass fractions displayed in Fig. 6c and d for the QT and ST co-\nagulants were quite different from their pristine tannin samples. The\nobtained mass fractions after Mannich modification were characterized\nby lower mass polymers (100-1400 m/z), compared with the tannin\nsample for quebracho. The increase in the proportion of lower mass\nfractions suggests that the aminomethylation process resulted in the\ndegradation of interflavonoid bonds in the quebracho tannin (Pash et al.,\n2001). This confirms previous findings that, when subjected to a low pH\nsuch as that required for the Mannich reaction, the proanthocyanidins in\nquebracho tannin depolymerize, releasing subunits as electrophilic\nflavan-3-ol intermediates (Kennedy and Jones, 2001). Also, the\nobserved reduction in the average molecular weight of the QT coagulant\nafter Mannich modification (920 Da) was of interest. In contrast, the\ndepolymerization effect due to acidic treatment witnessed with the QT\ncoagulant was not recorded in the ST coagulant spectrum. Although it\nhad a lower relative intensity, the ESI spectrum of the ST coagulant\nclearly showed an increase in the mass fractions compared to its pristine\nspruce tannin sample, and it produced a higher average molecular\nweight of 859 Da. This implies that an increase in polymer chain lengths\nmight have occurred in the ST coagulant due to the condensation and\ncrosslinking mechanism associated with the Mannich reaction (Grenda\net al., 2020). It is worth mentioning that both tannin coagulants\npossessed higher mass fractions (Fig. S4) and higher average molecular\n\n\nweights (Table S2) than M-FA coagulants modified under the same\nconditions.\n\n\nThe MALDI-TOF analysis displayed in Fig. 7 was in close agreement\nwith the molecular weight distribution pattern of the tannins and\nmodified coagulants produced by the ESI. The spectra displayed a series\nof peaks ranging from 500 to 2400 m/z. Quebracho and spruce tannin\nproduced an average molecular weight of 1089 and 769 Da in sequential\norder, whereas 880 and 830 Da were obtained for the modified QT and\nST coagulants, respectively. The average molecular weight estimated\nfrom the MALDI-TOF measurements was slightly lower than that ob-\ntained from the ESI measurement, except for the pristine spruce tannin.\nThe disparity in the estimated average molecular weight between the\nMALDI-TOF and ESI methods was attributed to the moderate mass ac-\ncuracy of MALDI-TOF, as signal intensity could differ significantly\nwithin one sample preparation (hot spot phenomenon) (Ole, 1994;\nStrupat, 2005).\n\n\n3.6. Coagulation performance\n\n\n3.6.1. Kaolin/river water\n\n\nFig. 8. displays the dosage curves for turbidity, total surface charge,\nultraviolet absorbance and specific ultraviolet absorbance of treated\nkaolin/river water samples. The tannin coagulants and the reference\ncoagulants (PIX-322 and M-FA) were all able to reduce the concentra-\ntion of suspended solids and colloids in the kaolin/river water sample\nimpressively (>90% turbidity reduction). A lower dosage was required\nwith the tannin coagulants (QT 25 mg/1 and ST 50 mg/l) and M-FA\ncoagulant (50 mg/l) for minimum residual turbidity when compared\nwith the PIX-322 (125 mg/l). It is noteworthy that the QT coagulant\nachieved clearly lower residual turbidity and slightly wider optimal\ndosage range than the ST and M-FA coagulants.\nThe pH of the kaolin/river water was close to neutral (7.5 \u00b1 0.2) and\nit is well known that ferric sulphate produces the positively charged\nspecies needed for charge neutralization at slightly acidic pH. However,\nif enough iron coagulant is added, then the solution pH becomes suitable\n\n\na) 60000000-\n\n\nb) 50\n\n\n50000000\n\n\n683 835\n\n\n295\n\n\n50000000-\n\n\nIntensity (au)\n\n\nIntensity (au)\n\n\n40000000-\n\n\n40000000-\n\n\n30000000-\n\n\n1108\n\n\n30000000\n\n\n956\n\n\n20000000-\n\n\n20000000-\n\n\n1398\n\n\n595\n\n\n165\n\n\n393 563\n\n\n1687\n\n\n10000000-\n\n\n10000000-\n\n\n835\n\n\n1944\n\n\n2331 2507\n\n\n1109\n\n\n0\n\n\n0\n\n\n500\n\n\n2500\n\n\n2500\n\n\n2000\n\n\n3000\n\n\n1000\n\n\n1500\nm/z\n\n\n500\n\n\n1500\n\n\n2000\n\n\n3000\n\n\n1000\n\n\nm/z\n\n\n0\n\n\nd) 16000000-\n\n\n12000000\n\n\n277\n\n\n287\n\n\n10000000-\n\n\nIntensity (au)\n\n\n14000000-\n\n\nIntensity (au)\n\n\n12000000\n\n\n415\n\n\n8000000\n\n\n10000000-\n\n\n6000000\n\n\n8000000\n\n\n835\n\n\n169\n\n\n664\n\n\n553\n\n\n6000000-\n\n\n4000000\n\n\n663\n\n\n169 415\n\n\n4000000-\n\n\n1108\n\n\n2000000-\n\n\n853\n\n\n962\n\n\n1057 1321 1453\n\n\n2000000\n\n\n1397 1670\n\n\n0\n\n\n0\n\n\n1500\nm/z\n\n\n500\n\n\n1500\nm/z\n\n\n2000\n\n\n2500\n\n\n3000\n\n\n500\n\n\n1000\n\n\n2000\n\n\n2500\n\n\n3000\n\n\n1000\n\n\nFig. 6. ESI mass spectrum for a) QT b) ST c) QT coagulant and d) ST coagulant. (FA 20.1 mmol, melamine 10.4 mmol, activation time 5 min, temperature 45 \u00b0C).\n\n\n9\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nb) 35000-\n\n\na)\n\n\n588\n\n\n683\n\n\n120000\n\n\n30000\n\n\nIntensity (au)\n\n\nIntensity (au)\n\n\n100000\n\n\n25000\n\n\n80000\n\n\n683\n\n\n20000-\n\n\n829 955\n\n\n60000\n\n\n15000\n\n\n759\n\n\n1145\n\n\n863\n\n\n40000\n\n\n1129\n\n\n10000\n\n\n557\n\n\n1401\n\n\n1963 2235\n\n\n1691\n\n\n20000\n\n\n5000\n\n\n0\n\n\n0\n\n\n600 800 1000 1200 1400 1600 1800 2000 2200 2400\n\n\n600 800 1000 1200 1400 1600 1800 2000 2200 2400\nm/z\n\n\nm/z\n\n\nd)\n\n\nc)\n\n\n553\n\n\n554\n\n\n400000\n\n\n200000-\n\n\nIntensity (au)\n\n\nIntensity (au)\n\n\n566\n\n\n150000-\n\n\n300000\n\n\n565\n\n\n100000-\n\n\n583\n\n\n200000\n\n\n985\n\n\n692\n\n\n584\n\n\n842\n\n\n50000-\n\n\n100000\n\n\n704\n\n\n1123\n\n\n992 1042\n\n\n(1273\n\n\n1273\n\n\n1493\n\n\n1423\n\n\n1280\n\n\n0\n\n\n0\n\n\n600 800 1000 1200 1400 1600 1800 2000 2200 2400\nm/z\n\n\n600 800 1000 1200 1400 1600 1800 2000 2200 2400\nm/z\n\n\nFig. 7. MALDI-TOF mass spectrum for a) quebracho tannin b) spruce tannin c) QT coagulant and d) ST coagulant. (FA 20.1 mmol, melamine 10.4 mmol, activation\ntime 5 min, temperature 45 \u00b0C).\n\n\nb)\n200-\n\n\na) 70\n\n\nM-FA Coagulant\nPIX-322\nQT Coagulant\nST Coagulant\n\n\nM-FA Coagulant\n\n\nPIX-322\n\n\n60\n\n\nQT Coagulant\nST Coagulant\n\n\nTotal surface charge (\u03bceq/l)\n\n\n150-\n\n\n50\n\n\nTurbidity (NTU)\n\n\n100\n\n\n40\n\n\n30\n\n\n50\n\n\n20\n\n\n10\n\n\n-50\n\n\n0\n\n\n150\n\n\n0\n\n\n50\n\n\n100\n\n\n200\n\n\n250\n\n\n300\n\n\n0\n\n\n50\n\n\n100\n\n\n150\n\n\n200\n\n\n250\n\n\n300\n\n\nDosage (mg/l)\n\n\nDosage (mg/l)\n\n\nc) 0.7\n\n\nd)\n\n\n0.6\n\n\n6-\n\n\n0.5\n\n\n5.\n\n\nSUVA (l/mg-m)\n\n\nUV 254 (1/cm)\n\n\n0.4\n\n\nM.FA Coagulant\nPIX-322\n\n\n0.3.\n\n\nM-FA Coagulant\n\n\nPIX-322\n\n\nQT Coagulant\nST Coagulant\n\n\nQT Coagulant\nST Coagulant\n\n\n0.2-\n\n\n0.1\n\n\n1\n\n\n0.0\n\n\n0-\n\n\n50\n50\n\n\n250\n\n\n300\n\n\n150\nDosage (mg/l)\n\n\n250\n\n\n0\n\n\n100\n\n\n150\n200\nDosage (mg/l)\n\n\n50\n\n\n100\n\n\n200\n\n\n300\n\n\nFig. 8. Effects of coagulant dosage on a) turbidity b) total surface charge c) UV254 absorbance and d) SUVA on the kaolin/river water using QT, ST, M-FA and PIX-\n\n\n322 coagulants.\n\n\n10\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nsignificantly at extreme coagulant dosage due to the evident increase in\nUV 254 absorbance, while the increment recorded with PIX-322 was\nattributed to low DOC values (data not shown).\n\n\nfor the generation of the required iron species (Bratby, 2016). This\nphenomenon was evident during the experiment as the pH of the treated\nwater was reduced from 7.5 to ~4.8, which was obtained after applying\nan iron coagulant dosage of 125 mg/1 before any significant turbidity\nreduction was recorded. On the other hand, the tannin-based and co-\nagulants were observed to work well at the initial pH of the kaolin/river\nwater sample and only moderately decreased the pH values to 6.9 and\n6.7 at the optimal dosage for the QT and ST coagulants, respectively. The\neffectiveness of tannin-based coagulants over a wide pH range has been\npreviously reported (Grenda et al., 2018). Compared to the tannin co-\nagulants, a much lower pH (5.9) was recorded for the M-FA coagulant at\noptimal dosage.\nThe TSC displayed in Fig. 8b indicates that increasing the coagulant\ndosage resulted in a decrease in surface charge (TSC becomes less\nnegative) with all of the tested coagulants. At the optimal dosages of\ntannin and M-FA coagulants based on the residual turbidites, the TSC\nvalues were close to zero, while overdosing increased turbidity and\nsurface charge. The lack of increased turbidity with the PIX-322 coag-\nulant even at excess coagulant concentrations supports the formation of\nelectroneutral iron species at above optimal coagulant dosage (Bratby,\n2016). In such an instance, the increase of coagulant dosage has little or\nno effect on the turbidity of the treated water. Fig. 8b shows that the TSC\ndosage curve of water samples treated with PIX-322 was more stable\nthan the other coagulants even at an excess dosage. TSC describes the\ncharge of the treated water. For PIX-322 coagulant, once the positive\ncharge required to neutralize the colloid is generated at optimal pH,\noverdosing of the iron salt results in the formation of Fe-hydroxide\nprecipitates, which settles with the flocs. On the other hand, tannin\nand M-FA coagulants stay dissolved in the solution and, when over-\ndosed, result in a higher TSC for the treated water, as demonstrated in\nFig. 8b. It is noteworthy that the increase in the supernatant's surface\ncharge was more pronounced with the QT and M-FA coagulants due to\ntheir higher cationic charge.\nThe fact that TSC was close to zero when optimal turbidity was ob-\ntained with both tannin coagulants indicates that the primary mecha-\nnism employed by the tannin coagulants was charge neutralization.\nAlso, this was supported by the lower dosage requirement observed for\nthe QT coagulant compared to the ST coagulant due to its higher charge\ndensity. This finding was consistent with earlier studies (Fang, 2007)\nwhich presented charge neutralization as the dominant mechanism used\nby tannin-based coagulants to destabilize colloids in kaolin/river mix-\ntures at neutral pH.\nUV 254 and SUVA 254 are vital indicators used to obtain information\nabout dissolved organic compounds in water samples. As shown in\nFig. 8c, the UV254 in the treated water sample decreases as the coagulant\ndosage increases. The decrease in UV254 absorbance translates into a\nhumic reduction of 90%, 80%, 69% and 68% in treated natural water at\noptimal dosage for the PIX-322, QT, M-FA, and ST coagulants, respec-\ntively. This is consistent with earlier researchers who reported that ferric\nsalts are more effective in removal of humic substances than tannin\ncoagulants (Heiderscheidt et al., 2016). Surprisingly there was a steep\nincrease in UV 254 absorbance with the M-FA coagulant, which tran-\nscended the initial absorbance value of the kaolin/river water when\ncoagulant concentration exceeded the optimal dosage. This indicates\nthat when overdosed, the residual M-FA coagulant in the treated water\naffect strongly UV254 absorbance. The other coagulants showed a minor\nincrease in the UV 254 absorbance when optimal dosage range was\nexceeded, which suggests good removal of organic compounds despite\noverdose. The initial SUVA value (>3 1/mg-m) displayed in Fig. 8d for\nthe kaolin/river mixture is a typical value of natural waters composed of\nhydrophobic, humic and aromatic organic compounds of higher mo-\nlecular weight (Edzwald and Van Benschoten, 1990). The SUVA value\nreduction below 2 1/mg--m at optimal dosages for all coagulants implies\nthat the optimally treated river/kaolin waters will later be characterized\nby non-humic, hydrophilic and lower molecular weight organic frac-\ntions. It is noteworthy that the SUVA value of M-FA increased\n\n\n3.6.2. Industrial process waters\n\n\nTo evaluate the performance of the coagulants in industrial appli-\ncations, the tannin coagulants were used to enhance particle settlement\nin two process waters (referred to hereinafter as PW1 and PW2). It is\nimportant to restate that the verifiable difference between the PW1 and\nPW2 water samples was that PW2 possessed a higher concentration of\nsuspended solids, a higher pH value, higher COD, higher UV254 absor-\nbance and higher conductivity, as shown in Table 1.\nFig. 9 shows the dosage curves of turbidity reduction, total surface\ncharge, UV254 and COD for the treated PW1 samples. The QT coagulant\nachieved a very high turbidity reduction (>85%) with a relatively low\ndosage (10 mg/l). However, with a higher dosage, the ST coagulant was\nalso able to achieve an impressive turbidity reduction of 76% with a\ncoagulant dosage of 50 mg/l, demonstrating the effectiveness of the\ncoagulants in the effluent of concern. In contrast to the kaolin/river\nwater, the dosage curve obtained from both coagulants showed that the\nturbidity of the treated water was stable over a wide range of coagulant\ndosage, and charge reversal did not occur even at an extreme coagulant\ndosage (Fig. 9b). However, the total surface charge approached zero as\nthe coagulant dosage increased. The stabilization of the surface charge\nover a wide dosage range tends to suggest that electrostatic patch\nadsorption was the dominant mechanism for colloid destabilization and\nthat charge neutralization may have occurred to some extent. For the\nelectrostatic patch mechanism to occur, chain polymers carrying\nopposite charges are believed to be adsorbed onto the colloids' surface,\nforming a charge-like mosaic, which is then attracted by other colloids.\nEarlier studies have noted that the destabilization of colloids by tannin-\nbased coagulants is complex and might have been ensured through\nvarious mechanisms (Fang, 2007; Hameed et al., 2016). Hameed et al.\n(2016) referred to both charge neutralization and bridging as the\ndominant mechanisms for colloid destabilization in their treated\neffluent.\n\n\nThe UV254 absorbance reduction with dosage for the PW1 sample is\nshown in Fig. 9c. The UV absorbance was partially reduced by the QT\ncoagulant but not by the ST coagulant. The dosage curve of chemical\noxygen demand (COD) for the treated PW1 sample revealed that, at\noptimal dosage for turbidity reduction, the QT and ST coagulants\nsimultaneously achieved 50% COD reductions with 10 mg/1 and 50 mg/\n1 dosages, respectively (Fig. 9d). Both tannin coagulants exhibited an\nidentical COD reduction pattern at lower coagulant dosages, but after\nthe dosage exceeded 60 mg/1, the COD was observed to increase with\nthe ST coagulant while the COD continued to decrease with the QT\ncoagulant.\nTo further evaluate the performance of the tannin-based coagulants\nin industrial applications, QT and ST coagulants were tested in another\nprocess water (PW2). A very small amount of coagulant was needed to\nbring the turbidity to a low level (Fig. 10a). This can probably be\nattributed to the availability of a nucleation site for floc formation\nprovided by the high concentration of suspended solids (Amy and\nChadik, 1983), thereby reducing the coagulation burden and favouring\nthe precipitation of suspended particles (Zhou et al., 2020). Addition-\nally, the TSC of PW2 was small (\u221211.5 \u00b5eq/1), which indicated a small\ncationic demand. The lack of a visible reduction in the total surface\ncharge of the treated water sample in Fig. 10b (TSC becoming less\nnegative) further affirms that the cationic charge of modified tannin\ncoagulants did not make any significant contribution to the destabili-\nzation of the colloids. Considering the lower molecular weight charac-\nteristic of the coagulants and the water quality of the PW2 sample, it is\nreasonable to conclude that the main destabilization mechanism\nemployed was patching.\nNeither of the tannin coagulants was able to reduce the UV absor-\nbance and COD of the PW2 sample (Figs. 10c and 10d). This was\n\n\n11\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\na) 60\n\n\nb)\n\n\n0\n\n\n60\n50\n\n\nTotal surface charge (peq/l)\n\n\n40\n\n\nTurbidity (NTU)\n\n\n30-\n\n\n20\n20\n\n\n-15\n\n\nQT Coagulant\nST Coagulant\n\n\nQT Coagulant\nST Coagulant\n\n\n10-\n\n\n-20\n\n\n550\n50\n\n\n50\n\n\n100\n\n\n150\n\n\n200\n\n\n100\nDosage (mg/l)\n\n\n150\n\n\n200\n\n\n0\n\n\n0\n\n\nDosage (mg/l)\n\n\nc) 0.20\n\n\nd) 50\n\n\n0.18\n\n\n10\n40\n\n\nChemical Oxygen demand (mg/l)\n\n\n0.16-\n\n\n0.14-\n\n\nUV 254(1/cm)\n\n\n0.12-\n\n\n30\n\n\n0.10-\n\n\n20\n\n\n0.08\n\n\n0.06-\n\n\n10\n10\n\n\n0.04-\n\n\nQT Coagulant\nST Coagulant\n\n\nQT Coagulant\nST Coagulant\n\n\n0.02-\n0.00\n\n\n0\n\n\n0\n\n\n100\nDosage (mg/l)\n\n\n50\n\n\n50\n\n\n150\n\n\n200\n\n\n0\n\n\n100\n\n\n150\n\n\n200\n\n\nDosage (mg/l)\n\n\nFig. 9. Effect of coagulant dosage on a) turbidity b) total surface charge c) UV 254 absorbance and d) COD of the PW1 water using QT and ST coagulants.\n\n\na) 160\n\n\nb)\n-7\n\n\n140\n\n\nTotal surface charge (\u03bceq/l)\n\n\n-8\n\n\n120\n\n\nTurbidity (NTU)\n\n\n100\n\n\n80\n08\n\n\n1\n\n\n60-\n\n\nQT Coagulant\nST Coagulant\n\n\n40-\n\n\nQT Coagulant\nST Coagulant\n\n\n-12\n\n\n20-\n\n\n-13\n\n\n0\n\n\n120 140 160\n\n\n0\n\n\n20\n\n\n40\n\n\n60\n\n\n80\nDosage (mg/l)\n\n\n100\n\n\n0\n\n\n20\n\n\n40\n\n\n60\n\n\n80\nDosage (mg/l)\n\n\n100\n\n\n120\n\n\n140 160\n\n\nd)80\n\n\nc) 0.25\n\n\n70\n\n\nChemical oxygen demand (mg/l)\n\n\n60\n\n\n0.20-\n\n\n50\n\n\nUV 254(1/cm)\n\n\n0.15-\n\n\n40\n\n\n30\n\n\n0.10-\n\n\n20\n\n\n0.05-\n\n\nQT Coagulant\nST Coagulant\n\n\nQT Coagulant\nST Coagulant\n\n\n10\n\n\n0.00\n\n\n0\n\n\n120\n\n\n140 160\n\n\n80 100\nDosage (mg/l)\n\n\n160\n\n\n0\n\n\n20\n\n\n40\n\n\n60\n\n\n100\nDosage (mg/l)\n\n\n80\n\n\n0\n\n\n20\n\n\n40\n\n\n60\n\n\n120\n\n\n140\n\n\nFig. 10. Effect of coagulant dosage on a) turbidity b) total surface charge and c) UV254 absorbance d) COD of the PW2 using QT and ST coagulants.\n\n\n12\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\nattributed to the inability of the tannin coagulants to retain their\ncationic charge at higher pH. The pH value of PW2 was 8.6 and\ndecreased only to 8.2 with the highest used dosage, thus the pH was at a\nhigher level than with PW1 and kaolin/river water. Earlier studies have\nhighlighted that the charge possession of tannin-based coagulant is pH-\ndependent. Fang (2007) found that the charge density of a tannin-based\ncoagulant decreased from 3.07 meq/g to 0.2 meq/g with increasing pH\nvalues from 4 to 9. Another possible reason for the coagulants' inef-\nfectiveness in reducing the UV absorbance and COD was the quality of\norganic matter in the process water. Marhaba and Pipada (2000) have\ndemonstrated that coagulants are less effective in removing simple and\nwell-dissolved organic matter, and others have reported that hydrophilic\ndissolved organics are not amenable to coagulation (Amy et al., 1987).\n\n\nQT JR-PW2\n\n\nST JR-PW2\n\n\nSSPW2\n\n\n3.7. Characterization of jar test residues and suspended solids of process\nwater samples\n\n\nT\n1040\n\n\n1030\n\n\n1050\n\n\n1020\n\n\nBE(eV)\n\n\nFig. 11. Overlaying comparison of Zn2p spectra for PW2 suspended solids\n(SSPW2) and jar test residues (QT JR-PW2 and ST JR-PW2).\n\n\nXPS analysis of the surface elemental content was performed on the\njar test residues of all of the treated water samples and suspended solids\nof the original process water samples. The XPS survey for the jar test\nresidues obtained from the kaolin/river water showed that they\npossessed new peaks located at 103.2 eV and 75 eV, assigned to Si2p and\nAl2p, respectively. Al and Si were believed to have originated from the\nkaolin introduced to the river water. Further analysis was not performed\non the kaolin/river jar test residue because no additional elements were\ndetected. Besides, analysis of the river water's nutrient and elemental\ncontent showed a very low concentration of elements (Table S4).\nThe XPS analysis of surface elemental content of suspended solids in\nthe PW1 (SSPW1) sample revealed that it contained only C, O and Si.\nThe Si2p peak was located at 101.5 eV in the Si2p high-resolution XPS\nspectrum. The XPS wide spectra revealed that the PW1 jar test residues\n(JR-PW1) from both tannin coagulants contained mainly C, N, O and Si.\nThe high-resolution spectra showed that the Si2p peaks of the jar test\nresidues were located at 101.7 eV while the N1s peaks were located at\nthe same BEs as pure tannin coagulants.\nThe XPS characterization of suspended solids in the PW2 (SSPW2)\nsample showed that the particles were composed of several elements\n(Ca, C, Mg, N, O, S, Si and Zn). Zn was the element with the most\ninteresting characteristics since the high-resolution Zn2p spectrum\nshowed two-component doublets. The fitted Zn2p3/2 peaks were located\nat 1022.1 eV and 1024.7 eV and the Zn2p1/2 peaks were positioned at\n1045.2 eV and 1047.8 eV, respectively. The fact that PW2 was water\nfrom a gas scrubber obtained from a ferrochrome smelter and a high\nresolution of S2p peak at 162.4 eV was attributed to metallic sulfide\n(Fantauzzi et al., 2015) suggests the presence of ZnS. Sulphur reacts with\nzinc at above hydrothermal temperature to form ZnS (Gao et al., 2015).\nThus, it is reasonable to conclude the sulphur exists as ZnS in the sus-\npended solids. X-ray diffraction (XRD) confirmed the presence of ZnS in\nthe suspended solids and further suggested that the crystalline zinc\nphases of the SSPW2 were wurtzite (~23%), elemental Zn (~15%) and\nsphalerite (~12%), supporting the multi-component structure observed\nwith XPS. The dual-peak structures observed with XPS have been re-\nported as a characteristic of multi-component zinc (Raveendra et al.,\n2018; Wang et al., 2015). Raveendra et al. (2018) ascribed the double\ncomponent structure to Zn2+ ions in Zn-O species associated with\nwurtzite. To understand the changes induced by the coagulants, XPS\nanalysis was carried out on the PW2 jar test residues (JR-PW2), which\nconfirmed the presence of all elements earlier detected in the SSPW2\n(Fig. S5). Overlaying the Zn2p spectra for the pristine sample and jar test\nresidues (Fig. 11) showed a broadening of the JR sample peaks by\n~1.0 eV towards higher binding energies. The broadening increased the\nenergy gap between the component peaks of the JR samples by 1.7 eV,\nfrom 2.1 eV (SSPW2) to ~3.8 eV (JR-PW2). Peak fitting of the Zn2p3/2\nof the JR-PW2 samples revealed that an additional peak at 1025.9 eV\nmight be related to Zn-N interactions (Fig. S6, only spectrum of QT\nJR-PW2 shown). Yahya et al. (2002) have reported the emergence of\n\n\nsimilar peaks as a result of complexation between zinc cation and the\nnitrogen in the amine of a chitosan coagulant.\nAnalysis of the surface elemental content of the PW2 water sample\npresented in supplementary Table S5 confirmed that sulphate and zinc\nwere present in high concentrations. An additional elemental analysis\nwas performed on the treated water samples (Table S5) to evaluate the\nresidual elemental concentration after coagulation. The probe into the\nresidual elemental concentration of the treated PW2 water sample\nshowed that the two tannin coagulants reduced the sulphate concen-\ntration by 20% from 400 mg/1 to ~320 mg/l. The elemental analysis of\nthe treated water sample also revealed that the QT coagulant reduced\nzinc from an initial concentration of 3.8 mg/1 to 1.1 mg/1 (74%). At the\nsame time, the ST coagulant achieved a zinc reduction of 93% (0.26 \u03bcg/\n1). Metal sequestration was not of the utmost importance during this\nstudy, and thus, the reason for the impressive performance of the ST\ncoagulant was not investigated. However, some studies have reported\nthe ability of N-glycosidic compounds (glucosamine derivatives) pro-\nduced as a side product in the ST coagulant to provide chelating sites for\nthe complexation of various cationic metals (Allscher et al., 2008; Yano\net al., 1999; Yano and Mikata, 2002). More importantly, Yano et al.\n(1999) have demonstrated that N-glucoside ligands from D-glucosamine\ncould effectively ligate zinc atoms.\n\n\n4. Conclusions\n\n\nThis study has shown that melamine can be used as a nitrogenating\nagent during the aminomethylation of tannins. With this synthesis, low-\nto-medium charge density and low molecular weight products with low\nFA content can be obtained, depending on the tannin raw material. A\nhigher dosage FA and melamine had a detrimental effect on the stability\nof the QT coagulant, while a higher modification temperature acceler-\nated the crosslinking process in both tannin modifications. The results of\nthe XPS analysis confirmed that iminium ions were successfully\nembedded into the tannin matrix after aminomethylation. ESI and\nMALDI-TOF analyses revealed that the Mannich modification led to the\ndepolymerization of the quebracho tannin. In general, the results ob-\ntained from the jar tests showed that both types of modified coagulants\nwere effective in turbidity reduction, with QT slightly outperforming the\nST coagulant due to its relatively higher charge density. Nevertheless,\nthe ST coagulant could still be considered a potential coagulant, espe-\ncially in industrial applications, because of its stability. Coagulation\nstudies demonstrated that charge neutralization, bridging and patching\nmechanisms could be employed by the tannin-based coagulants when\napplied as a coagulant, depending on the effluent surface characteristics.\nThe results of the XPS analysis of jar test residue suggested that metals\nsequestered from the treated effluent were bonded within stable\n\n\n13\n\n\nA. Bello and T. Leivisk\u00e4\n\n\nIndustrial Crops & Products 176 (2022) 114336\n\n\norganometallic complexes.\n\n\nSpringer, Berlin, Heidelberg, pp. 341\u2013359. https://doi.org/10.1007/978-3-642-\n76093-8_22.\n\n\nEnvironment and Climate Change Canada. Draft Screening Assessment of Certain\n\n\nOrganic Flame Retardants Substance Grouping. https://www.canada.ca/en/envi\nronment-climate-change/services/evaluating-existing-substances/risk-management-\n\n\nDeclaration of Competing Interest\n\n\nscope-melamine.html.\n\n\nThe authors declare that they have no known competing financial\ninterests or personal relationships that could have appeared to influence\nthe work reported in this paper.\n\n\nFang, G., 2007. Evaluation and Performance of a Tannin-based Polymer as a Coagulant in\nWater Treatment. Doctoral dissertation. Imperial College London. https://core.ac.\nuk/download/pdf/295694.pdf.\nFantauzzi, M., Elsener, B., Atzei, D., Rigoldi, A., Rossi, A., 2015. Exploiting XPS for the\nidentification of sulfides and polysulfides. RSC Adv. 5, 75953-75963. https://doi.\norg/10.1039/c5ra14915k.\nFeng, Q., Zhao, W., Wen, S., Cao, Q., 2017. Copper sulfide species formed on malachite\nsurfaces in relation to flotation. J. Ind. Eng. Chem. 48, 125\u2013132. https://doi.org/\n10.1016/j.jiec.2016.12.029.\nFood and Agriculture Organization, General standard for contaminants and toxins in\nfood and feed (CXS 193-1995), 2019. https://www.fao.org/fao-who-codexalimentar\n\n\nAcknowledgements\n\n\nThis work was funded by University of Oulu graduate school and\nMaa- ja vesitekniikan tuki ry. The authors would also like to acknowl-\nedge the support of VTT (Technical Research Centre of Finland) and the\nSilvachimica team for providing the tannin samples.\n\n\nius/sh-proxy/en/?lnk=1&url=https%253A%252F%252Fworkspace.fao.org%252F\nsites%252Fcodex%252FStandards%252FCXS%2B193-1995%252FCXS_193e.pdf.\n\n\n(Accessed 17 October 2021).\n\n\nGao, X., Wang, J., Yu, J., Xu, H., 2015. Novel ZnO-ZnS nanowire arrays with\nheterostructures and enhanced photocatalytic properties. CrystEngComm 17,\n6328-6337. https://doi.org/10.1039/c5ce01078k.\n\n\nAppendix A. 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Jpn 75, 2097-2113. https://doi.org/10.1246/\nbcsj.75.2097.\n\n\nYano, S., Inoue, S., Yasuda, Y., Tanase, T., Mikata, Y., Kakuchi, T., Tsubomura, T.,\nYamasaki, M., Doe, M., 1999. First Synth. Charact. zinc 1851-1855.\n\n\nYao, L.D., Li, F.Y., Li, J.X., Jin, C.Q., Yu, R.C., 2005. Study of the products of melamine (C\n3N 6H 6) treated at high pressure and high temperature, Phys. Status Solidi Appl.\nMater. Sci. 202, 2679-2685. https://doi.org/10.1002/pssa.200421177.\n\n\nYeh, T.S., Lin, T.C., Chen, C.C., Wen, H.M., 2013. Analysis of free and bound\nformaldehyde in squid and squid products by gas chromatography-mass\nspectrometry. J. Food Drug Anal. 21, 190\u2013197. https://doi.org/10.1016/j.\njfda.2013.05.010.\n\n\nYubo, C., Yong, Z., Ning, Y., 2016. Synthesis and characterization of bio-based phenol-\nformaldehyde resol resins from bark autoclave extractives.. Prod. J. 66, 18-28.\nhttps://doi.org/10.13073/FPJ-D-15-00001.\nZhang, L., Tang, X., Rothman, N., Vermeulen, R., Ji, Z., Shen, M., Qiu, C., Guo, W.,\nLiu, S., Reiss, B., Freeman, L.B., Ge, Y., Hubbard, A.E., Hua, M., Blair, A., Galvan, N.,\nRuan, X., Alter, B.P., Xin, K.X., Li, S., Moore, L.E., Kim, S., Xie, Y., Hayes, R.B.,\nAzuma, M., Hauptmann, M., Xiong, J., Stewart, P., Li, L., Rappaport, S.M.,\nHuang, H., Fraumeni, J.F., Smith, M.T., Lan, Q., 2010. Occupational exposure to\nformaldehyde, hematotoxicity, and leukemia-specific chromosome changes in\ncultured myeloid progenitor cells. Cancer Epidemiol. Biomark. Prev. 19, 80\u201388.\nhttps://doi.org/10.1158/1055-9965.EPI-09-0762.\nZhao, Y.X., Gao, B.Y., Shon, H.K., Cao, B.C., Kim, J.H., 2011. Coagulation characteristics\nof titanium (Ti) salt coagulant compared with aluminum (Al) and iron (Fe) salts.\nJ. Hazard. Mater. 185, 1536\u20131542. https://doi.org/10.1016/j.jhazmat.2010.10.084.\nZhou, J., Liu, J., Zhao, Z., Peng, W., Cui, F., Liang, Z., 2020. Microcystis aeruginosa-laden\nwater treatment using peroxymonosulfate enhanced Fe(II) coagulation: Performance\nand the role of in situ formed Fe304. Chem. Eng. J. 382, 123012 https://doi.org/\n10.1016/j.cej.2019.123012.\nZhu, G., Zheng, H., Zhang, Z., Tshukudu, T., Zhang, P., Xiang, X., 2011. Characterization\nand coagulation-flocculation behavior of polymeric aluminum ferric sulfate (PAFS).\nChem. Eng. J. 178, 50\u201359. https://doi.org/10.1016/j.cej.2011.10.008.\n\n\nSincero P., A., Gregoria S., A., 2003. Physical-chemical treatment of water and\n\n\nwastewater. IWA publishing, Boca Raton, London, New York, Washington D.C.\nSon, D., re kr Cho, S., Nam, J., Lee, H., Kim, M., 2020. X-ray-based spectroscopic\ntechniques for characterization of polymer nanocomposite materials at a molecular\nlevel. Polymers 12, 1053. https://doi.org/10.3390/POLYM12051053.\nStrupat, K., 2005. Molecular weight determination of peptides and proteins by ESI and\nMALDI. Methods Enzym. 405, 1\u201336. https://doi.org/10.1016/S0076-6879(05)\n05001-9.\n\n\n15\n15\n"}, "expected_output": {"claims": [{"unit": "wt%", "value": 0.00042, "evidence": ["An estimation of the\nconcentration of FA content in a treated water sample, if 100 mg/1\ndosages of coagulants are applied and assuming all residual FA stays in\nthe treated water, shows a significantly low residual FA of 1.8 mg/1\n(0.00018 wt%) and 4.2 mg/1 (0.00042 wt%) for the QT and ST co-\nagulants modified at 70 \u00b0C, respectively."]}, {"unit": "wt%", "value": 0.00018, "evidence": ["An estimation of the\nconcentration of FA content in a treated water sample, if 100 mg/1\ndosages of coagulants are applied and assuming all residual FA stays in\nthe treated water, shows a significantly low residual FA of 1.8 mg/1\n(0.00018 wt%) and 4.2 mg/1 (0.00042 wt%) for the QT and ST co-\nagulants modified at 70 \u00b0C, respectively."]}]}, "metadata": {"product_category": "Chemical products", "request_id": "req_7f9e3c6cfdacc660"}}
{"id": "1144329ddbb36c9477d01b04", "input": {"query": "What are the amounts of woolen oil per kg of fiber during leather yarn spinning preparation and blending? I need the kg per kg of fiber.", "source_url": "https://data.epo.org/publication-server/rest/v1.0/publication-dates/20220316/patents/EP3967800NWA1/document.pdf", "document_text": "(19)\n\n\nEurop\u00e4isches\nPatentamt\n\n\nEuropean\n\n\nPatent Office\n\n\nOffice europ\u00e9en\ndes brevets\n\n\nEP 3 967 800 A1\n\n\n(11)\nEUROPEAN PATENT APPLICATION\npublished in accordance with Art. 153(4) EPC\n\n\n(12)\n\n\n(43) Date of publication:\n\n\n(51) International Patent Classification (IPC):\n\n\nD01G 13/00 (2006.01)\nD01G 15/00 (2006.01)\n\n\nD02G 3/06 (2006.01)\nD01G 11/00 (2006.01)\nD02G 1/02 (2006.01)\nD01G 9/00 (2006.01)\n\n\n16.03.2022 Bulletin 2022/11\n\n\nD01H 5/00 (2006.01)\n\n\n(21) Application number: 19927639.5\n\n\n(22) Date of filing: 17.05.2019\n\n\n(52) Cooperative Patent Classification (CPC):\nY02W 30/66\n\n\n(86) International application number:\n\n\nPCT/KR2019/005896\n\n\n(87) International publication number:\n\n\nWO 2020/226220 (12.11.2020 Gazette 2020/46)\n\n\n(84) Designated Contracting States:\n\n\n(71) Applicant: Atko Planning Inc.\n\n\nAL AT BE BG CH CY CZ DE DK EE ES FI FR GB\nGR HR HU IE IS IT LI LT LU LV MC MK MT NL NO\nPL PT RO RS SE SI SK SM TR\n\n\nPaju-si, Gyeonggi-do 10951 (KR)\n\n\n(72) Inventor: KIM, Ji Eon\n\n\nDesignated Extension States:\nBA ME\n\n\nSeoul 06990 (KR)\n\n\nDesignated Validation States:\n\n\n(74) Representative: Botti, Mario\nBotti & Ferrari S.p.A.\n\n\nKH MA MD TN\n\n\nVia Cappellini, 11\n20124 Milano (IT)\n\n\n(30) Priority: 09.05.2019 KR 20190054077\n\n\n(54) METHOD FOR PREPARING SPUN LEATHER YARN USING LEATHER FIBER, AND SPUN\nLEATHER YARN PREPARED ACCORDING TO PREPARATION METHOD\n\n\n(57) Disclosed are a method for preparing spun\nleather yarn using leather fiber, and spun leather yarn\nprepared according to the preparation method. The spun\nleather yarn prepared according to the preparation meth-\nod enables the manufacture of fabrics, knitted goods,\nand lace goods, which have the character of leather, and\nthus can be effectively used for various fashion applica-\ntions. In addition, the spun leather yarn can replace ex-\nisting leather materials, which have been used in bags,\nathletic shoes, and the like, and thus is not affected by\nmaterial supply in accordance with seasonal changes,\nand accordingly, the economic effect thereof is excellent.\nMoreover, the spun leather yarn has warmth retention\nand general flame retardancy, which are the properties\nof natural leather, and thus is of very high quality.\n\n\n[Fig.1]\n\n\nthe step of mixing a leather fiber, a common fiber and\na polymer fiber to prepare a mixed fiber;\n\n\nS100\n\n\nthe step of introducing water and woolen oil\ninto the mixed fiber and mixing the thereof;\n\n\n-S200\n\n\nthe step of introducing an anti-static agent\ninto the mixed fiber which the water and the woolen oil\nhave been introduced and mixing thereof; and\n\n\nS300\n\n\nthe step of spinning the mixed fiber into\nwhich the anti-static agent is introduced to prepare\na spun leather yarn;\n\n\nS400\n\n\nEP 3 967 800 A1\n\n\nPrinted by Jouve, 75001 PARIS (FR)\n\n\nEP 3 967 800 A1\n\n\nDescription\n\n\n[Technical Field]\n\n\n[0001] The present invention relates to a method for preparing spun leather yarn using leather fiber, and spun leather\nyarn prepared according to the preparation method.\n\n\n5\n\n\n[Background Art]\n\n\n[0002] Natural leather is a tough shell wrapping an animal body, and widely used in various fields, such as bags,\nshoes, furniture, and others, due to its excellent physical properties. Natural leather is separated from the animal body\nand is manufactured into a product through chemical and physical processes. Among various processes, a process of\ncutting the natural leather according to a desired design of the product is essential. However, the cutting process produces\na great deal of leather waste, and most of the leather waste is incinerated and buried. So, it comes to the fore as a\nserious environment issue. Therefore, it is necessary to study and develop various recycling methods of leather waste.\n[0003] So, the inventors of the present invention studied spun leather yarn utilizing leather waste, and found out it was\npossible to manufacture high-quality spun leather yarn, which can be manufactured as fabrics, knitted goods, and lace\ngoods, which have the nature of leather, can replace existing leather materials, which have been used in bags, athletic\nshoes, and the like, can be produced without being affected by seasons, and has warmth retention and general flame\nretardancy, which are the properties of natural leather. Accordingly, the inventors completed this invention.\n[0004] In connection with the above, Japanese Patent Laid-open Publication No. 11-021600 discloses a manufacturing\nmethod of bulky regenerated leather.\n\n\n10\n\n\n15\n\n\n20\n20\n\n\n[Disclosure]\n\n\n25\n\n\n[Technical Problem]\n\n\n[0005] Accordingly, the present invention has been made in an effort to solve the above-mentioned problems occurring\nin the prior arts, and it is an object of the present invention to provide a method for preparing spun leather yarn using\nleather fiber.\n\n\n60\n30\n\n\nThe technical problem to be solved by the present invention is not limited to the technical problem as mentioned\nabove, and another technical problem, which is not mentioned, could be clearly understood by those having ordinary\nskill in the art to which the present invention pertains based on the description below.\n\n\n[0006]\n\n\n35 [Technical Solution]\n\n\n[0007] To achieve the above objects, the present invention provides a method for preparing spun leather yarn using\nleather fiber including the steps of: mixing leather fiber, common fiber, and polymer fiber to prepare mixed fiber; putting\nwater and woolen oil into the mixed fiber, and mixing them; putting an anti-static agent into the mixed fiber into which\nwater and woolen oil have been put, and mixing them; and spinning the mixed fiber into which the anti-static agent is\nput to prepare spun leather yarn.\n[0008] Moreover, the leather fiber is obtained from leather waste, and the leather waste may be flaky scraps.\n[0009] Furthermore, the common fiber is one selected from groups of cellulose-based natural fiber, protein-based\nnatural fiber, mineral-based fiber, cellulose-based artificial fiber, protein-based artificial fiber, rubber-based fiber, inorganic\nfiber, and acetate-based fiber.\n\n\n40\n\n\n45\n\n\n[0010] Additionally, the polymer fiber is one selected from groups of polyamide-based fiber, polyester-based fiber,\npolyurethane-based fiber, polyurea-based fiber, polyacryl-based fiber, polyvinyl alcohol-based fiber, polyvinyl chloride-\nbased fiber, polyvinylidene chloride-based fiber, polypropylene-based fiber, polyethylene-based fiber, polystyrene-based\nfiber, and polyfluoroethylene-based fiber.\n[0011] In addition, the leather fiber, the common fiber, and the polymer fiber are mixed at the rate of 40 parts by weight\nto 80 parts by weight of common fiber, and 20 parts by weight to 60 parts by weight of polymer fiber on the basis of 100\nparts by weight of leather fiber.\n[0012] The water and the woolen oil are mixed at the rate of 5 parts by weight to 10 parts by weight of water and 2\nparts by weight to 10 parts by weight of woolen oil on the basis of 100 parts by weight of the mixed fiber.\n[0013] 1 part by weight to 5 parts by weight of the anti-static agent is put on the basis of 100 parts by weight of the\nmixed fiber.\n\n\n60\n50\n\n\n50\n55\n\n\n[0014] Each of the step of putting water and woolen oil into the mixed fiber and mixing them, and the step of putting\nan anti-static agent into the mixed fiber into which water and woolen oil have been put and mixing them further includes\n\n\n2\n\n\nEP 3 967 800 A1\n\n\na step of aging the mixed fiber for 8 hours to 15 hours after mixing.\n\n\n[0015] The step of preparing spun leather yarn includes steps of: carding the mixed fiber, into which the anti-static\nagent is put, using a carding machine to prepare sliver; and applying twist and draft to the sliver to prepare spun leather\nyarn.\n\n\n[0016] The carding machine includes a worker speed controller to control the speed of the worker.\n[0017] The sliver manufactured by the carding machine is 15 mm or more in length.\n\n\n5\n\n\n[0018] After the step of manufacturing the sliver, the step of manufacturing the spun leather yarn further includes: a\ncombing step of removing foreign matters from the sliver; a drawing step of producing a single strand of the sliver after\nthe combing step; and drawing the sliver again after the drawing step and roving the sliver.\nThe drawing step is carried out using a roller having a plurality of wire fillets, and an interval between the plurality\nof wire fillets is 0.01 mm to 1 mm.\n\n\n10\n\n\n[0019]\n\n\n[0020] The drawing step is carried out by spraying water in order to prevent static electricity.\n\n\n[0021]\n\n\nThe method for preparing spun leather yarn using leather fiber further includes a covering step after the step\nof manufacturing spun leather yarn.\n\n\n15\n\n\n[Advantageous Effects]\n\n\n[0022] The spun leather yarn prepared using leather fiber manufactured according to an embodiment of the present\ninvention enables the manufacture of fabrics, knitted goods, and lace goods, which have the nature of leather, thereby\nbeing capable of being effectively used for various fashion applications.\n[0023] In addition, the spun leather yarn prepared using leather fiber manufactured according to an embodiment of\nthe present invention can replace existing leather materials, which have been used in bags, athletic shoes, and the like,\nand thus is not affected by material supply in accordance with seasonal changes, and accordingly, the economic effect\nthereof is excellent. Moreover, the spun leather yarn has warmth retention and general flame retardancy, which are the\nproperties of natural leather, and thus is of very high quality.\n[0024] It should be understood that the effects of the present invention are not limited to the above-mentioned effects,\nand includes all effects derivable from the configuration of the present invention described in the detailed description or\nclaims of the present invention.\n\n\n20\n20\n\n\n25\n\n\n60\n30\n\n\n[Description of Drawings]\n\n\n[0025]\n\n\nFIG. 1 is a schematic diagram illustrating a method for preparing spun leather yarn using leather fiber according to\nan embodiment of the present invention.\nFIG. 2 is schematic diagram illustrating a structure of a carding machine according to an embodiment of the present\ninvention.\n\n\n35\n\n\n[Best Mode]\n\n\n40\n\n\nHereinafter, preferred examples or embodiments of the present invention will be described in detail. However,\nthe present invention can be implemented in various forms, is not limited to the preferred embodiments disclosed herein.\nExample: Manufacturing spun leather yarn using leather fiber according to an embodiment of the present in-\n45 vention\n\n\n[0026]\n\n\n[0027] Spun leather yarn was manufactured using a method of manufacturing spun leather yarn using leather fiber\naccording to an embodiment of the present invention.\nFirst, a leather crusher destroyed bonding of leather fiber after impurities of leather waste were shaken off.\nAfter that, a leather scutcher separated leather fiber from the leather waste of which leather fiber bonding was destroyed.\n[0029] Next, 40 parts by weight of rayon and 60 parts by weight of nylon on the basis of 100 parts by weight of the\nseparated leather fiber were mixed to prepare mixed fiber. After that, 8 parts by weight of water and 5 parts by weight\nof woolen oil on the basis of 100 parts by weight of the mixed fiber were put and mixed, and then, the mixed fiber was\naged for twelve hours. After that, 2 parts by weight of anti-static agent on the basis of 100 parts by weight of the mixed\nfiber was put and mixed, and then, the mixed fiber was aged for twelve hours.\nNext, the aged mixed fiber was spun to prepare spun leather yarn. In this instance, the spun leather yarn was\nmanufactured under a condition that a carding machine is adjusted to be reduced 40% in worker speed and to have\ntemperature of 30\u00b0C and humidity of 65% so as to manufacture the spun leather yarn.\n\n\n[0028]\n\n\n50\n60\n\n\n50\n55\n\n\n[0030]\n\n\n3\n\n\nEP 3 967 800 A1\n\n\nComparative example 1: Preparation of natural fiber Cotton being sold in the market was obtained.\n\n\nComparative example 2: Preparation of regenerated fiber Rayon being sold in the market was obtained.\n\n\nComparative example 3: Preparation of synthetic fiber Nylon being sold in the market was obtained.\nExperimental example: Measurement of physical properties of\n\n\n5\n\n\nspun leather\n\n\nyarn\n\n\n[0031] The physical properties of the spun leather yarn manufactured according to the Example of the present invention,\nand the physical properties of the cotton, the rayon, and the nylon prepared in the comparative examples 1 to 3 were\nmeasured, and then, the results were recorded in the following Table 1.\n\n\n10\n\n\n[Table 1]\nComparative\nExample 1\n4\n\n\nExample of the\npresent invention\n\n\nComparative\n\n\nComparative\nExample 3\n\n\n15\n\n\nExample 2\n\n\nDuctility\n(%)\n\n\n24.83\n\n\n28\n\n\nDry\ncondition\n\n\n18\n\n\nWet\ncondition\n\n\n25.5\n\n\n24\n\n\n36\n\n\n20\n20\n\n\nElastic recovery rate\n(%)\n\n\n82\n\n\n92.3\n\n\n75\n\n\n100\n\n\nVery bad\n\n\nVery bad\n\n\nAntibiotic action\n\n\nExcellent\n\n\nGood\n\n\n25\n25\n\n\n[0032] As shown in the Table 1, the spun leather yarn manufactured according to the Example of the present invention\nshowed excellent ductility and elastic recovery, and especially, showed excellent antibiotic action by showing excellent\nantibacterial rate as a result of an antibacterial test.\n\n\n60\n30\n\n\n[Mode for Invention]\n\n\n[0033] Hereinafter, the present invention will be described in more detail. However, embodiments or examples of the\npresent invention may be implemented in several different forms, and are not limited to the embodiments described\nherein. In addition, the present invention will be defined only by the scope of the appended claims.\n[0034] Additionally, the terminology used herein is for the purpose of describing particular embodiments or examples\nonly and is not intended to limit the present invention. As used herein, the singular forms are intended to include the\nplural forms as well, unless the context clearly indicates otherwise. It will be understood that the terms \"comprise\",\n\"include\", \"have\", etc. in the whole specification of the present invention mean that other components are not excluded\nunless explicitly described otherwise, and other components may in fact be included.\n\n\n35\n\n\n40\n\n\n[0035] In a first aspect of the present invention, provided is a method for preparing spun leather yarn using leather\nfiber including the steps of: mixing leather fiber, common fiber, and polymer fiber to prepare mixed fiber; putting water\nand woolen oil into the mixed fiber, and mixing them; putting an anti-static agent into the mixed fiber into which water\nand woolen oil have been put, and mixing them; and spinning the mixed fiber into which the anti-static agent is put to\nprepare spun leather yarn.\nHereinafter, referring to FIG. 1, the method for preparing spun leather yarn using leather fiber according to the\nfirst aspect of the present invention will be described in detail step by step. FIG. 1 is a schematic diagram illustrating a\nmethod for preparing spun leather yarn using leather fiber according to an embodiment of the present invention.\n[0037] First, according to an embodiment of the present invention, the method for preparing spun leather yarn using\nleather fiber includes a step (S100) of mixing leather fiber, common fiber, and polymer fiber to prepare mixed fiber.\n[0038] In the embodiment of the present invention, the leather fiber is obtained from leather waste, and the leather\nwaste may be flaky scraps. The flaky form may be a flat type having a long shaft and a short shaft. If scraps of the leather\nwaste are used, because the scraps have a predetermined length, spun leather yarn of a thread type can be manufactured\neasily. In this instance, the long shaft and the short shaft of the flaky form may be respectively 250mm to 450mm, and\nmay have the same length. In the meantime, the leather fiber includes 8 wt% to 10 wt% of water on the basis of the\ngross weight of the leather fiber. As described above, because the leather fiber contains a small quantity of water, the\nnext processes can be carried out smoothly to manufacture spun leather yarn.\n[0039] In an embodiment of the present invention, leather fiber can be obtained from the leather waste through the\n\n\n45\n\n\n[0036]\n\n\n50\n50\n\n\n55\n55\n\n\nEP 3 967 800 A1\n\n\nfollowing steps of: destroying bonding between leather fiber of leather waste using a leather crusher; and separating\nthe leather fiber from the leather waste, of which leather fiber bonding was destroyed, using a leather scutcher. In this\ninstance, the leather crusher is to destroy the bonding between the leather fiber of the leather waste by crushing the\nleather waste, and the crushing may be carried out by a well-known crushing method, but must be carried out by a\ncrushing method which is strong enough to destroy the bonding between leather fiber. Moreover, the leather scutcher\nis used to separate the leather fiber from the leather waste of which leather fiber bonding was destroyed. For instance,\nthe leather scutcher may separate the leather fiber from the leather waste using the heavy weight of the leather fiber.\n[0040] In an embodiment of the present invention, the common fiber may be one selected from groups of cellulose-\nbased natural fiber, protein-based natural fiber, mineral-based fiber, cellulose-based artificial fiber, protein-based artificial\nfiber, rubber-based fiber, inorganic fiber, and acetate-based fiber. The polymer fiber may be one selected from groups\nof polyamide-based fiber, polyester-based fiber, polyurethane-based fiber, polyurea-based fiber, polyacryl-based fiber,\npolyvinyl alcohol-based fiber, polyvinyl chloride-based fiber, polyvinylidene chloride-based fiber, polypropylene-based\nfiber, polyethylene-based fiber, polystyrene-based fiber, and polyfluoroethylene-based fiber.\n[0041] In an embodiment of the present invention, in more detail, the common fiber or the polymer fiber may be at\nleast one selected from groups of nylon, kapok, flax, ramie, hemp, jute, manila, abaca, sisal hemp, coir, sheep wool,\ngoat wool, cashmere, camel wool, alpaca, rabbit fur, vicuna wool, llama wool, cultivated silk, wild silk, asbestos, metal\nfiber, glass fiber, rock wool, slag wool, carbon fiber, rayon, lanital, soylon fiber, ardil fiber, uicara fiber, casein fiber,\nacetate, acetylated rayon fiber, chlorinated rubber, hydrochlorinated rubber, amilan, rilsan, purlon L, terylene, tetron,\ndacron, kodel, vycron, spandex, lycra, vycron, purlon U, urylon, glasslon, orlon, acrilan, creslan, exlan, vonnel, cashmilon,\nverel, dynel, kanekalon, vinylon, vinyon, teviron, Pesse, saran, krehalon, herculon, pylen, revon, hizex, Algil, Salon,\nand Teflon. However, the above kinds of the common fiber or the polymer fiber are enumeration of desirable examples,\nand all kinds of the common fiber can be used with no specific limitation without affecting the objects and effects of the\npresent invention.\n[0042] In the embodiment of the present invention, the leather fiber, the common fiber, and the polymer fiber are mixed\nat the rate of 40 parts by weight to 80 parts by weight of common fiber, and 20 parts by weight to 60 parts by weight of\npolymer fiber on the basis of 100 parts by weight of leather fiber. Preferably, the common fiber, and the polymer fiber\nare mixed at the rate of 50 parts by weight to 70 parts by weight of common fiber, and 30 parts by weight to 50 parts by\nweight of polymer fiber on the basis of 100 parts by weight of leather fiber. More preferably, the common fiber, and the\npolymer fiber are mixed at the rate of 60 parts by weight of common fiber, and 40 parts by weight of polymer fiber on\nthe basis of 100 parts by weight of leather fiber. If the mixing rate of the common fiber and the polymer fiber is higher\nthan the rate, because the content of the leather fiber is relatively low, texture of leather may get worse. Furthermore,\nIf the mixing rate of the common fiber and the polymer fiber is lower than the above rate, because the content of the\nleather fiber is relatively high, the leather fiber, the common fiber, and the polymer fiber are not mixed well, and spinning\nis not performed well in the spinning step since the content of basic filaments is too small.\n[0043] Next, in an embodiment of the present invention, the method for preparing spun leather yarn using leather fiber\nincludes a step (S200) of putting water and woolen oil into the mixed fiber, and mixing them.\n[0044] In the embodiment of the present invention, the water and the woolen oil are to increase a mixing rate of the\nleather fiber, the common fiber, and the polymer fiber. The water and the woolen oil are mixed at the rate of 5 parts by\nweight to 10 parts by weight of water and 2 parts by weight to 10 parts by weight of woolen oil on the basis of 100 parts\nby weight of the mixed fiber. Preferably, the water and the woolen oil are mixed at the rate of 7 parts by weight to 9 parts\nby weight of water and 3 parts by weight to 7 parts by weight of woolen oil on the basis of 100 parts by weight of the\nmixed fiber. More preferably, the water and the woolen oil are mixed at the rate of 8 parts by weight of water and 5 parts\nby weight of woolen oil on the basis of 100 parts by weight of the mixed fiber. If less than 5 parts by weight of water is\nput and less than 2 parts by weight of woolen oil is put on the basis of 100 parts by weight of the mixed fiber, because\nthe content of the water and the woolen oil is too low, the leather fiber, the common fiber, and the polymer fiber are not\nmixed well. In addition, if more than 10 parts by weight of water is put and more than 10 parts by weight of woolen oil is\nput on the basis of 100 parts by weight of the mixed fiber, because the content of the water and the woolen oil is too\nhigh, the spun leather yarn manufactured in the next step may not show the texture of leather.\n\n\n5\n\n\n10\n\n\n15\n\n\n20\n20\n\n\n25\n25\n\n\n30\n\n\n35\n\n\n40\n\n\n45\n\n\n[0045]\n\n\nNext, in an embodiment of the present invention, the method for preparing spun leather yarn using leather fiber\nincludes a step (S300) of putting an anti-static agent into the mixed fiber into which water and woolen oil have been put,\nand mixing them.\n\n\n50\n60\n\n\n[0046] In the embodiment of the present invention, the anti-static agent is put in order to increase a mixing rate of the\nleather fiber, the common fiber and the polymer fiber. 1 part by weight to 5 parts by weight of the anti-static agent is put\non the basis of 100 parts by weight of the mixed fiber. Preferably, 1 part by weight to 3 parts by weight of the anti-static\nagent is put on the basis of 100 parts by weight of the mixed fiber. More preferably, 2 parts by weight of the anti-static\nagent is put on the basis of 100 parts by weight of the mixed fiber. If less than 1 part by weight of water is put, because\nthe content of the anti-static agent is too low, the leather fiber, the common fiber, and the polymer fiber are not mixed\nwell. In addition, if more than 5 parts by weight of anti-static agent is put, because the content of the anti-static agent is\n\n\n55\n55\n\n\n5\n\n\nEP 3 967 800 A1\n\n\ntoo high, the spun leather yarn manufactured in the next step may not show the texture of leather.\n\n\n[0047] In an embodiment of the present invention, the step (S200) of putting water and woolen oil into the mixed fiber\nand mixing them, and the step (S300) of putting an anti-static agent into the mixed fiber into which water and woolen oil\nhave been put and mixing them may be carried out at the same time. However, if the steps are carried out in order, the\nmixing rate of the leather fiber, the common fiber, and the polymer fiber may be increased.\n[0048] In an embodiment of the present invention, each of the step (S200) of putting water and woolen oil into the\nmixed fiber and mixing them, and the step (S300) of putting an anti-static agent into the mixed fiber into which water\nand woolen oil have been put and mixing them may further include a step of aging the mixed fiber for 8 hours to 15 hours\nafter mixing. In this instance, preferably, the aging step may be carried out for 10 hours to 14 hours. More preferably,\nthe aging step may be carried out for 12 hours. That is, in each step, the mixed fiber is aged for a predetermined period\nof time after the water, the woolen oil, and the anti-static agent are put and mixed, so that the mixing rate of the leather\nfiber, the common fiber, and the polymer fiber is increased.\n[0049] In an embodiment of the present invention, the method for preparing spun leather yarn using leather fiber may\nrepeat the steps of: (S100) mixing leather fiber, common fiber, and polymer fiber to prepare mixed fiber; (S200) putting\nwater and woolen oil into the mixed fiber, and mixing them; (S300) putting an anti-static agent into the mixed fiber into\nwhich water and woolen oil have been put, and mixing them, several times. Preferably, the method for preparing spun\nleather yarn using leather fiber may repeat the steps (S100), (S200), and (S300) four times. In this instance, the method\nfor preparing spun leather yarn using leather fiber includes: a first mixing step of first mixing 40 wt% to 60 wt% of mixed\nfiber on the basis of the gross weight of the mixed fiber to be manufactured; a second mixing step of adding and mixing\n20% to 30% of the gross weight of the mixed fiber to be manufactured; a third mixing step of adding and mixing 10% to\n20% of the gross weight of the mixed fiber to be manufactured; and a fourth mixing step of adding and mixing 5% to\n10% of the gross weight of the mixed fiber to be manufactured. That is, fibers to be mixed are added and mixed at a\npredetermined ratio in several steps, so that the leather fiber can be evenly mixed with the common fiber and the polymer\nfiber. However, the repeated number is not limited to four, and may be five, six or more.\n[0050] Next, in an embodiment of the present invention, the method for preparing spun leather yarn using leather fiber\nincludes a step (S400) of spinning the mixed fiber into which the anti-static agent is put to prepare spun leather yarn.\nReferring to FIG. 2, the step (S400) of spinning the mixed fiber into which the anti-static agent is put to prepare spun\nleather yarn will be described in more detail.\n[0051] In the embodiment of the present invention, the step (S400) of preparing spun leather yarn includes steps of:\ncarding the mixed fiber, into which the anti-static agent is put, using a carding machine 101 which brushing the mixed\nfiber in the direction of grains to prepare sliver; and applying twist and draft to the sliver to prepare spun leather yarn.\n[0052] In the embodiment of the present invention, carding is a mixing step to mix the mixed fiber at uniform mixing\nrate. Through the carding step, the mixed fiber is separated into several strands, impurities or foreign matters contained\nin the mixed fiber are removed, and the mixed fiber is concentrated into a web having required weight per unit area. In\ndetail, the mixed fiber is supplied to a roller covered with wire fillets rotating at rapid speed by a lattice to be opened thin.\nThe mixed fiber which is opened thin is supplied to a drum (doffer) covered with wire fillets rotating at slow speed, so\nthat a web having a predetermined weight can be formed while the mixed fiber passes through the drum. In this instance,\nthe carding machine 101 includes a worker 110 for performing opening. In the present invention, the carding machine\n101 may further include a worker speed controller 210 for controlling speed of the worker 110. Referring to FIG. 2, the\nworker speed controller 210 may be mounted inside a cylinder illustrated in the drawing so as to control speed of the\nworker 110 according to the length of the mixed fiber supplied to the carding machine. Meanwhile, preferably, the speed\nof the worker 110 is controlled to be reduced by about 30% to 50% in comparison with the normal speed of a conventional\nworker. That is, because the speed of the worker 110 is lower than the speed of the conventional worker, the mixed\nfiber which is relatively shorter in length is uniformly transferred along the worker 110 without being scattered outside\nthe worker so as to form the web.\n\n\n5\n\n\n10\n\n\n15\n\n\n20\n20\n\n\n25\n\n\n30\n\n\n35\n\n\n40\n\n\n45\n\n\n[0053] In the embodiment of the present invention, the sliver manufactured by the carding machine is 15 mm or more\nin length, preferably, is 20 mm or more in length, and more preferably, is 31 mm or more in length. If the sliver manufactured\nby the carding machine is less than 15 mm, the mixed fiber cannot be manufactured in the form of yarn, and so, the\nspun leather yarn cannot be manufactured.\n\n\n50 [0054] In an embodiment of the present invention, after the step of manufacturing the sliver, the step (S400) of man-\nufacturing the spun leather yarn further includes: a combing step of removing foreign matters from the sliver; a drawing\nstep of producing a single strand of the sliver after the combing step; and drawing the sliver again after the drawing step\nand roving the sliver.\n\n\n[0055] The combing step is to remove foreign matters contained in the sliver, align fiber, and arrange fine hairs. In\nthis instance, the combing step is carried out on a wooden plate in the conventional art, but if the combing step is carried\non the wooden plate, the sliver or the web transferred from the carding machine may be torn. Therefore, preferably, the\ncombing step is carried out on a rubber plate in order to prevent the sliver or the web from being torn. Moreover, an\ninput amount of the sliver put into the combing step is about 50% of the input amount of sliver in the conventional method.\n\n\n50\n55\n\n\n6\n\n\nEP 3 967 800 A1\n\n\nSo, the method of manufacturing the spun leather yarn further according to the present invention can increase an output\nof spun leather yarn and can manufactured spun leather yarn having uniform thickness (uniformity) since a break of the\nsliver is prevented.\n[0056] In the embodiment of the present invention, the drawing step is to manufacture a single strand of the sliver,\nand is carried out using a roller having a plurality of wire fillets. In this instance, an interval between the plurality of wire\nfillets is 0.01 mm to 1 mm, preferably, 0.05 mm to 0.5 mm, and more preferably, 0.1 mm. Furthermore, the number of\nrevolutions of the roller having the wire fillets is 120 rpm to 130 rpm. Meanwhile, the drawing step is carried out by\nspraying water in order to prevent static electricity.\n[0057] In the embodiment of the present invention, the roving step is to rove the sliver by drawing again and rubbing\nthe sliver after the drawing step, and may includes a step of separating the sliver in the same thickness.\n[0058] In the embodiment of the present invention, the step of applying twist and draft to the sliver to prepare spun\nleather yarn is to release the sliver into the thickness of yarn by applying strength and uniformity to the sliver. That is,\nthe step of applying twist and draft to the sliver is a process to make the sliver have a desired thickness by drawing it,\nconcentrate and snarl the fiber by applying twist, and apply strength by providing friction between the fibers. In this\ninstance, the twist number of the fiber is 400 TM to 1,000 TM, and the thickness is 10 yarn count, 15 yarn count, 20 yarn\ncount, or 25 yarn count.\n[0059] In the embodiment of the present invention, the step (S400) of manufacturing spun leather yarn may be carried\nout at temperature of 20\u00b0C to 40\u00b0C and at humidity of 50% to 80%, preferably, at temperature of 25\u00b0C to 35\u00b0C and at\nhumidity of 60% to 70%, since the mixed fiber may be scattered into the air during the spinning step due to the short\nlength of the mixed fiber. If the mixed fiber is spun at the temperature and at the humidity, the spinning work can be\ncarried out without scattering of the mixed fiber so as to easily manufacture spun leather yarn.\n[0060] In an embodiment of the present invention, the method of manufacturing spun leather yarn may further include\na covering step after the step (S400) of manufacturing spun leather yarn.\n\n\n5\n\n\n10\n\n\n15\n\n\n20\n20\n\n\n[0061] In the embodiment of the present invention, the covering step is carried out to reinforce physical properties,\n25 such as tensile strength and flexible rate, and is started from a step of holding the prepared spun leather yarn on a\ncovering machine after winding of the prepared spun leather yarn is finished. Yarn used in the covering step may be 30\ndenier to 150 denier polyester. Alternatively, nylon or functional yarn, such as high stretch yarn, high shrinkage yarn, or\nbulky yarn, may be used for the yarn used in the covering step. S-shaped covering or Z-shaped covering may be used\nas a covering method, or a double covering which the S-shaped covering and the Z-shaped covering are performed at\nthe same time may be carried out. In case of the double covering, after the covering step, the covered yarn has a shape\nclose to alphabet S. The covering step may include a step of twisting the yarn additionally. In this instance, preferably,\nthe twist number is 100 TM to 250 TM.\n\n\n60\n30\n\n\n[Explanation of reference numerals]\n\n\n35\n\n\n[0062]\n\n\ncarding machine\n\n\n101:\n\n\n110:\n\n\nworker\n\n\n40\n\n\nworker speed controller\n\n\n210:\n\n\n[Industrial Applicability]\n\n\n[0063] The spun leather yarn prepared using leather fiber manufactured according to an embodiment of the present\ninvention enables the manufacture of fabrics, knitted goods, and lace goods, which have the nature of leather, thereby\nbeing capable of being effectively used for various fashion applications.\n[0064] Moreover, the spun leather yarn prepared using leather fiber manufactured according to an embodiment of the\npresent invention can replace existing leather materials, which have been used in bags, athletic shoes, and the like, and\nthus is not affected by material supply in accordance with seasonal changes, and accordingly, the economic effect\nthereof is excellent. Moreover, the spun leather yarn has warmth retention and general flame retardancy, which are the\nproperties of natural leather, and thus is of very high quality.\n[0065] It should be understood that the effects of the present invention are not limited to the above-mentioned effects,\nand includes all effects derivable from the configuration of the present invention described in the detailed description or\nclaims of the present invention.\n\n\n45\n\n\n50\n60\n\n\n55\n50\n\n\n7\n\n\nEP 3 967 800 A1\n\n\nClaims\n\n\nA method for preparing spun leather yarn using leather fiber comprising the steps of:\n\n\n1.\n\n\nmixing leather fiber, common fiber, and polymer fiber to prepare mixed fiber;\nputting water and woolen oil into the mixed fiber, and mixing them;\n\n\n5\n\n\nputting an anti-static agent into the mixed fiber into which water and woolen oil have been put, and mixing them;\nand\n\n\nspinning the mixed fiber into which the anti-static agent is put to prepare spun leather yarn.\n\n\n10\n10\n\n\nThe method according to claim 1, wherein the leather fiber is obtained from leather waste, and the leather waste\nmay be flaky scraps.\n\n\n2.\n\n\n3. The method according to claim 1, wherein the common fiber is one selected from groups of cellulose-based natural\nfiber, protein-based natural fiber, mineral-based fiber, cellulose-based artificial fiber, protein-based artificial fiber,\nrubber-based fiber, inorganic fiber, and acetate-based fiber.\n\n\n15\n\n\nThe method according to claim 1, wherein the polymer fiber is one selected from groups of polyamide-based fiber,\npolyester-based fiber, polyurethane-based fiber, polyurea-based fiber, polyacryl-based fiber, polyvinyl alcohol-based\nfiber, polyvinyl chloride-based fiber, polyvinylidene chloride-based fiber, polypropylene-based fiber, polyethylene-\nbased fiber, polystyrene-based fiber, and polyfluoroethylene-based fiber.\n\n\n4.\n\n\n20\n20\n\n\n5. The method according to claim 1, wherein the leather fiber, the common fiber, and the polymer fiber are mixed at\nthe rate of 40 parts by weight to 80 parts by weight of common fiber, and 20 parts by weight to 60 parts by weight\nof polymer fiber on the basis of 100 parts by weight of leather fiber.\n\n\n25\n25\n\n\nThe method according to claim 1, wherein the water and the woolen oil are mixed at the rate of 5 parts by weight\nto 10 parts by weight of water and 2 parts by weight to 10 parts by weight of woolen oil on the basis of 100 parts by\nweight of the mixed fiber.\n\n\n6.\n\n\n60\n30\n\n\n7. The method according to claim 1, wherein 1 part by weight to 5 parts by weight of the anti-static agent is put on the\nbasis of 100 parts by weight of the mixed fiber.\n\n\n8. The method according to claim 1, wherein each of the step of putting water and woolen oil into the mixed fiber and\nmixing them, and the step of putting an anti-static agent into the mixed fiber into which water and woolen oil have\nbeen put and mixing them further includes a step of aging the mixed fiber for 8 hours to 15 hours after mixing.\n\n\n35\n\n\nThe method according to claim 1, wherein the step of preparing spun leather yarn includes steps of:\n\n\n9.\n\n\n40\n\n\ncarding the mixed fiber, into which the anti-static agent is put, using a carding machine to prepare sliver; and\napplying twist and draft to the sliver to prepare spun leather yarn.\n\n\n10. The method according to claim 9, wherein the carding machine includes a worker speed controller to control the\nspeed of the worker.\n\n\n45\n\n\n11. The method according to claim 9, wherein the sliver manufactured by the carding machine is 15 mm or more in length.\n12. The method according to claim 9, wherein after the step of manufacturing the sliver, the step of manufacturing the\nspun leather yarn further includes:\n\n\n60\n50\n\n\na combing step of removing foreign matters from the sliver;\n\n\na drawing step of producing a single strand of the sliver after the combing step; and\ndrawing the sliver again after the drawing step and roving the sliver.\n\n\n50\n55\n\n\n13. The method according to claim 12, wherein the drawing step is carried out using a roller having a plurality of wire\nfillets, and\n\n\nwherein an interval between the plurality of wire fillets is 0.01 mm to 1 mm.\n\n\n8\n\n\nEP 3 967 800 A1\n\n\n14. The method according to claim 12, wherein the drawing step is carried out by spraying water in order to prevent\nstatic electricity.\n\n\n15. The method according to claim 1, further including:\n\n\na covering step after the step of manufacturing spun leather yarn.\n\n\n5\n\n\n10\n10\n\n\n1\n15\n\n\n20\n20\n\n\n25\n\n\n30\n30\n\n\n35\n\n\n40\n\n\n45\n\n\n50\n50\n\n\n55\n55\n\n\n9\n\n\nEP 3 967 800 A1\n\n\n[Fig.1]\n\n\nthe step of mixing a leather fiber, a common fiber and\na polymer fiber to prepare a mixed fiber;\n\n\nS100\n\n\nthe step of introducing water and woolen oil\ninto the mixed fiber and mixing the thereof;\n\n\nS200\n\n\nthe step of introducing an anti-static agent\ninto the mixed fiber which the water and the woolen oil\nhave been introduced and mixing thereof; and\n\n\nS300\n\n\nthe step of spinning the mixed fiber into\nwhich the anti-static agent is introduced to prepare\na spun leather yarn;\n\n\nS400\n\n\n[Fig.2]\n\n\n101\n\n\n110\n110\n\n\n210\n\n\n10\n10\n\n\nEP 3 967 800 A1\n\n\nINTERNATIONAL SEARCH REPORT\n\n\nInternational application No.\nPCT/KR2019/005896\n\n\n15\n\n\nCLASSIFICATION OF SUBJECT MATTER\n\n\nA.\n\n\nD02G 3/06(2006.01)i, DOIG 13/00(2006.01)i, D01G 11/00(2006.01)i, D01G 15/00(2006.01)i, D02G 1/02(2006.01)i,\nD01H 5/00(2006.01)i, D01G 9/00(2006.01)i\n\n\nAccording to International Patent Classification (IPC) or to both national classification and IPC\n\n\nB. FIELDS SEARCHED\n\n\n10\n10\n\n\nMinimum documentation searched (classification system followed by classification symbols)\n\n\nD02G 3/06; A43B 23/02; D02G 3/00; D02G 3/02; D02G 3/04; D03D 15/00; D01G 13/00; D01G 11/00; D01G 15/00; D02G 1/02;\nD01H 5/00: D01G 9/00\n\n\nDocumentation searched other than minimum documentation to the extent that such documents are included in the fields searched\nKorean utility models and applications for utility models: IPC as above\nJapanese utility models and applications for utility models: IPC as above.\n\n\n15\n\n\nElectronic data base consulted during the international search (name of data base and, where practicable, search terms used)\neKOMPASS (KIPO internal) & Keywords: leather, waste, blend\n\n\n20\n20\n\n\nC. DOCUMENTS CONSIDERED TO BE RELEVANT\n\n\nCitation of document, with indication, where appropriate, of the relevant passages\n\n\nCategory*\n\n\nRelevant to claim No.\n\n\nKR 10-2006-0133027 A (ZHANG, Liwen) 22 December 2006\n\n\n1-15\n\n\nY\n\n\nSee abstract; claims 1-3; paragraphs [0014], [0017]-[0023].\n\n\n25\n\n\nKR 10-1154602 BI (KOREA SILK RESEARCH INSTITUTE et al.) 08 June 2012\nSee claims 1, 5, 6, 9.\n\n\n1-15\n\n\nY\n\n\nJP 2008-512574 A (LIU, L.) 24 April 2008\n\n\nY\n\n\n9-14\n\n\nSee paragraphs [0005]-[0007], [0013].\n\n\n30\n30\n\n\nKR 10-2004-0027419 A (SEO, Jung Eun) 01 April 2004\n\n\nY\n\n\n15\n\n\nSee abstract; and claims 1, 2.\n\n\nKR 10-2016-0133939 A (SHOEMAX CO., LTD. et al.) 23 November 2016\nSee the entire document.\n\n\nA\n\n\n1-15\n\n\n35\n\n\n40\n\n\nFurther documents are listed in the continuation of Box C.\n\n\nSee patent family annex.\n\n\nSpecial categories of cited documents:\n\n\n*\n\n\nlater document published after the international filing date or priority\ndate and not in conflict with the application but cited to understand\nthe principle or theory underlying the invention\n\n\n\"T\"\n\n\n\"A\" document defining the general state of the art which is not considered\nto be of particular relevance\n\n\n\"E\" earlier application or patent but published on or after the international \"X\" document of particular relevance; the claimed invention cannot be\nfiling date\n\n\nconsidered novel or cannot be considered to involve an inventive\nstep when the document is taken alone.\ndocument of particular relevance; the claimed invention cannot be\nconsidered to involve an inventive step when the document is\ncombined with one or more other such documents, such combination\nbeing obvious to a person skilled in the art\n\n\n\"L\" document which may throw doubts on priority claim(s) or which is\n\n\n45\n\n\ncited to establish the publication date of another citation or other\nspecial reason (as specified)\n\n\n\"y\"\n\n\n\"O\" document referring to an oral disclosure, use, exhibition or other\n\n\nmeans\n\n\n\"p\" document published prior to the international filing date but later than \"&\" document member of the same patent family\nthe priority date claimed\n\n\nDate of the actual completion of the international search\n\n\nDate of mailing of the international search report\n\n\n50\n50\n\n\n06 FEBRUARY 2020 (06.02.2020)\n\n\n06 FEBRUARY 2020 (06.02.2020)\n\n\nAuthorized officer\n\n\nName and mailing address of the ISA/KR\n\n\nKorean Intellectual Property Office\n\n\nGovernment Complex Daejeon Building 4, 189, Cheongsa-ro, Seo-gu,\nDaejeon, 35208, Republic of Korea\n\n\nFacsimile No. +82-42-481-8578\n\n\nTelephone No.\n\n\n55\n55\n\n\nForm PCT/ISA/210 (second sheet) (January 2015)\n\n\n11\n\n\nEP 3 967 800 A1\n\n\nINTERNATIONAL SEARCH REPORT\nInformation on patent family members\n\n\nInternational application No.\n\n\nPCT/KR2019/005896\n\n\n5\n\n\nPatent family\nmember\n\n\nPatent document\n\n\nPublication\ndate\n\n\nPublication\ndate\n\n\ncited in search report\n\n\n10\n10\n\n\nCN 1263908 C\nCN 1673430 A\n\n\nKR 10-2006-0133027 A\n\n\n22/12/2006\n\n\n12/07/2006\n28/09/2005\n15/11/2007\n16/08/2007\n11/12/2012\n27/10/2005\n\n\nJP 2007-532786 A\n\n\nUS 2007-0186352 A1\n\n\nUS 8328878 B2\n\n\nWO 2005-100655 A1\n\n\n15\n\n\nKR 10-1154602 B1\n\n\n08/06/2012\n\n\nKR 10-2010-0115037 A\n\n\n27/10/2010\n\n\nJP 2008-512574 A\n\n\n24/04/2008\n\n\n20/12/2006\n02/03/2005\n13/06/2007\n27/03/2008\n16/03/2006\n\n\nCN 1291084 C\nCN 1587468 A\nEP 1795634 A1\n\n\n20\n20\n\n\nUS 2008-0072628 A1\nWO 2006-026899 A1\n\n\nNone\n\n\n01/04/2004\n\n\nKR 10-2004-0027419 A\n\n\nKR 10-2016-0133939 A\n\n\n23/11/2016\n\n\nNone\n\n\n25\n\n\n30\n30\n\n\n35\n\n\n40\n40\n\n\n45\n\n\n50\n50\n\n\n55\n55\n\n\nForm PCT/ISA/210 (patent family annex) (January 2015)\n\n\n12\n\n\nEP 3 967 800 A1\n\n\nREFERENCES CITED IN THE DESCRIPTION\n\n\nThis list of references cited by the applicant is for the reader's convenience only. It does not form part of the European\npatent document. Even though great care has been taken in compiling the references, errors or omissions cannot be\nexcluded and the EPO disclaims all liability in this regard.\n\n\nPatent documents cited in the description\n\n\nJP 11021600 A [0004]\n\n\n\u2022\n\n\n13\n"}, "expected_output": {"claims": [{"unit": "kg/kg", "value": 0.02, "evidence": ["The water and the woolen oil are mixed at the rate of 5 parts by\nweight to 10 parts by weight of water and 2 parts by weight to 10 parts by weight of woolen oil on the basis of 100 parts\nby weight of the mixed fiber."]}, {"unit": "kg/kg", "value": 0.1, "evidence": ["The water and the woolen oil are mixed at the rate of 5 parts by\nweight to 10 parts by weight of water and 2 parts by weight to 10 parts by weight of woolen oil on the basis of 100 parts\nby weight of the mixed fiber."]}]}, "metadata": {"product_category": "Textiles, footwear & apparel", "request_id": "req_236ab2da0a031922"}}
{"id": "a7f2f446bf8ba716d5bc80e9", "input": {"query": "What is the conversion ratio of fresh tomatoes to polpa (finely chopped tomatoes) in kg of fresh tomatoes per kg of final polpa product? This will help understand the yield and any water/ingredient additions.", "source_url": "https://mutti-parma.com/app/uploads/sites/24/2024/09/mutti-catalogue-2020-update-04-copy.pdf", "document_text": "SOLO POMODORO.\nMUTTI\n\n\nPARMA\n\n\nFood service\n\n\nPRODUCT\nCATALOGUE\n\n\nSun, earth, water and time.\nMutti's recipe for success in your kitchen.\n\n\n\"WE WOULD\nDESCRIBE\nOURSELVES FIRST\nAND FOREMOST AS\nTHE GUARDIANS OF\nAUTHENTIC TOMATO\nFLAVOUR.\"\n\n\n- Francesco Mutti\n\n\nOur entire product range stems from a single basic\ningredient: a flawless fruit at perfect ripeness.\nWorking with this single raw ingredient means that\nwe are tomato experts. We know all there is to know\nabout our tomatoes because we monitor the fruit\nfrom field to the factory, and from the factory right to\nyour kitchen. Fostering a constant dialogue with our\nseed selectors, farmers and transporters. Selecting\nonly superior quality with the most rigorous analysis.\nUsing traditional methods passed on through the\ngenerations with innovative production processes.\nThese steps guarantee that each of our products\nmeets the same quality standards. That's the promise\nfrom the two-lion brand since 1899.\n\n\nMutti Milestones\n\n\nA LONG\n\n\nLOVE\nSTORY\n\n\nIN ITALIAN KITCHENS\nFOR OVER A CENTURY\n\n\nSTRATTA DI PONODORA\nQUALITA SUPERIORE\n\n\nA family tradition steeped in innovation and\ngreat passion. The history of the Mutti\ncompany is bound together by the love of the\npure tomato and the need to preserve one of\nnature's greatest gifts at perfect ripeness.\nThis is a dedication to the pursuit of perfection\nwoven through four generations for more than\na hundred years.\n\n\nFRAT MUTTI\nSOM BASILICANOVA\n\n\n2\n\n\n1850\n\n\nFMUTTI\n\n\nPOMODORO\n\n\nA PIONEER\n\n\nGiovanni Mutti (1804-1894), who has been described as \"ahead of his time\", showed real innovation by applying the crop rotation technique to his farming.\nThe 'sidereal system', which is now widely used in organic farming, takes into account the influence of the sun and the moon on plant cultivation. This allowed the\nsoil to recover its nutrients while cutting down on the use of both natural and chemical fertilisers. Crop rotation is an important practice in modern farming,\nand is proof that the secret to exceptional tomatoes lies in Italy's ancestral, countryside traditions.\n\n\n1899\n\n\nTHE ADVENTURE BEGINS\n\n\nGiovanni's nephews, Marcellino Mutti (1862-1941) and Callisto (1870 - 1936), created the Fratelli Mutti company. The focus of the family's farming tradition began\nto shift towards manufacturing, but the tomatoes were still processed using artisanal methods, which gained in efficiency year after year.\nThe burgeoning tomato industry had started to take root.\n\n\n1909\n\n\nTHE VISION OF AN INDUSTRY\n\n\nUgo Mutti (1893 - 1980) was very young when he designed a process that would transform the family's farming business. He suggested to his father, Marcellino,\nthat they create a small factory to produce tomato extract. The sous-vide cooking technique had only just been fine-tuned and tomato concentrate, instead of being\npreserved in rectangular bars, could now be canned. This was a turning point for the future canned food industry, and the product was sold under the Mutti brand\nname right from the start. Known as 'conserva nera' or 'black preserve', it was the forerunner of modern tomato concentrate.\n\n\n1911\n\n\nTHE TWO-LION BRAND\n\n\nMarcellino Mutti registered his brand in the same year as an international exhibition celebrating 50 years since the unification of Italy. It shows two fighting lions\nprotecting the brand's first awards. These prizes were a testament to the passion with which he produced his tomato concentrate. In a country where illiteracy was\nstill the norm, it was essential for food companies to make their products stand out with a strong and simple image. This image had to be highly memorable and\neasy to recognise at the local shop. Marcellinoi knew the Mutti logo was a success when shoppers started requesting the concentrate from 'the two-lion brand'.\n\n\n1914\n\n\nAWARDS OF PRESTIGE\n\n\nShortly after we started producing tomatoes, the quality of our products had already become legendary, and Mutti began to receive its first prestigious awards.\nAfter the Medaglia d'Oro di 1\u00b0 Grado (first grade gold medal) at the 1911 exhibition, came the Gran Croce award, the highest order of merit in Italy. As well as being\na food product that was easy to store and sell, Mutti tomato extract became the symbol of Italian gastronomic excellence. Since that time, countless accolades have\nsingled out the quality of our products. They form the basis of the bond of trust that Mutti first established with its customers over a hundred years ago.\n\n\nDouble Concentrate\n\n\nPROVENANCE\n\n\nSOLO\n\n\nPOMODORO.\n\n\nMUTTI\n\n\nAst\n\n\nPARMA\n\n\nDOPPIO\nCONCENTRATO DI POMODORO\nDOUBLE\nCONCENTR\u00c9 DE TOMATE\n\n\nAND\n\n\nTOMAT\nNITED TOW!\n\n\nPRODUCTION\n\n\n100%\nITALIAN\n\n\nF. Mutti\n\n\nRICH & INTENSE\n\n\nunique product with a bright red colour, fresh\nfragrance and intense flavour. Adding some\nMutti Double Concentrated Tomato Paste to\ndishes is a great way to bring flavour, texture\nand colour. It's perfect for recipes that require\nlong cooking times such as bolognese sauce,\ncasseroles and stews. It can also be used to\nenhance quick recipes as a quick dash of\nseasoning for pasta or a delicious sauce to\n\n\nMutti Double Concentrated Tomato Paste is\nobtained from tomatoes from the Emilia\nRomagna and neighbouring regions such as\nPiemonte, Lombardia, Toscana and Veneto.\nThe tomato, superior in quality thanks to\nMutti's trademark rigorous quality control\nprocesses, is evaporated according to a\nhistoric, natural recipe that preserves all\nproperties of the fresh tomatoes. Only the\nheart of the tomato is extracted to create a\n\n\n6x1\n\n\nPACK\nSIZES\n\n\n6KG OF\nTOMATOES = 1KG\nOF CONCENTRATE\n\n\n2.15KG TIN\n\n\nDECISO\n\n\n28\u00b0\n\n\nBOLD, RICH\nFLAVOUR.\nDENSE TEXTURE\n\n\nMINIMUM\nBRIX\n\n\naccompany a dish.\n\n\n4\n\n\nPost-war Innovation\n\n\nNITIVE 21701\n\n\n1951\n\n\nGRAN MARCA DUE LEONI\nSHE CHORIFICENCE ATOA T\n\n\nODORO\n\n\nTHE ALLE PERSONE DELICATE\n\n\nT. the Multi\n\n\nGMODO\n\n\nMUTTI\nCALILICANOVA\n\n\nSTRAT\n\n\nCONCENTRATE IN A TUBE\n\n\nfor millions of Italian families. Once people\nhad overcome their initial surprise, they were\nwon over by its practical nature and the\nmoney they could save. The concentrate\ndidn't spoil, because it was no longer in\ncontact with the air, and could be used\nas needed. In addition, the cap in the shape\nof a thimble was useful for housewives.\nThe legendary \"thimble tube\" had arrived.\n\n\nConcentrate has come a long way since it\nwas first made in the form of 'conserva\nnera', cooked in large saucepans for many\nhours and then dried in the sun. Thanks to\nthe invention of sous-vide cooking,\nconcentrate made its appearance, followed\nby double concentrate in 1922, and from\n1938, triple concentrate. The packaging also\nsaw major changes - the large boxes used\nby greengrocers to sell concentrate to their\n\n\nconsumers were replaced with smaller packs\nthat were suitable for domestic consumption.\nThis was how Mutti concentrate first found\nits way into Italian kitchens. These days,\nthere is nothing surprising about preserving\nfood in an aluminium tube. But in 1951, this\nform of packaging was synonymous with\ncream or toothpaste: the idea of Ugo Mutti to\npack concentrate in a tube was a major\nrevolution, both for the preserve industry and\n\n\n5\n\n\nPolpa Finely Chopped Tomatoes\n\n\nPROVENANCE\n\n\nSOLO POMODORO\nMUTTI\n\n\nSOLD FORODORG\nMUTTI\n\n\nPARMA\n\n\nPARMA\n\n\nPOLPA\n\n\nPOMODORI IN FINISSIMI PEZZI\nTOMATES CONCASS\u00c9ES FINES\nFINELY CHOPPED TOMATOES\nFEINSTES TOMATENFRUCHTFLEISCH\n\n\nPOLPA\nPOMODORI IN FINISSIMI PEZZI\nTOMATES CONCASS\u00c9ES FINES\n\n\nPRODU\n\n\nPROFESSIONAL\n\n\n100%\nITALIANO\n\n\nPRODUZION\n100%\nITALIANO\n\n\nMODORO\n\n\nMutter\n\n\nMutte\n\n\nFICATI\n\n\nSECRET SINCE 1971\n\n\nadditives to balance the flavour. Containing\n\n\nMutti Polpa is made using tomatoes from the\nEmilia Romagna region, picked perfectly ripe\nand packed within hours of harvesting.\nCombine this with Mutti's patented\ncold-processing technique and the result is a\nrich, red, fragrant polpa, free of acidity\n\n\nPACK\nSIZES\n\n\nmore juice than Passata, it therefore can\n\n\n5x1\n\n\nalso be used for recipes that require a long\n\n\n5KG OF\nTOMATOES = 1KG\nOF POLPA\n\n\n5KG POUCH\n4KG TIN\n400G TIN\n\n\ncooking time or higher temperatures.\n\n\nMutti Polpa is also suitable for using\nuncooked thanks to its fresh, fragrant\nflavour and the total absence of artificial\nadditives and impurities.\n\n\nFresco\n\n\nregulator, that closely captures most the\n\n\n5.2\u00b0\n\n\nflavour of the just-harvested tomato.\n\n\nFRESH\nFLAVOUR,\nRICH TEXTURE\n\n\nMINIMUM\nBRIX\n\n\nMutti Polpa requires minimal cooking and no\n\n\n6\n\n\nLegend in a Tin\n\n\nORIO ALIMENTAR\n\n\nVINILIQUORI\n\n\n1971\n\n\nPOLPA di\nPOMODORO\n\n\nDUE LEON.\n\n\nMUTTI\nPARMA BASILICANOVA ITALIA\n\n\nTHE BIRTH OF POLPA\n\n\nFor Mutti, the 1970s began with a spectacular turn of events that was\nmasterfully orchestrated by Ugo Mutti. He gave himself an ambitious\nobjective: to make a fresh, ready-to-use product with a less watery\nconsistency than traditional peeled tomatoes. After several trials and\nwith the help of an innovative cold-processing method, the legendary\ntomato pulp was created. It is still Mutti's flagship product today, and is\nloved by millions of cooks around the world for its fresh taste and the\nquality of its tomatoes.\n\n\n7\n\n\nPassata\n\n\nPROVENANCE\n\n\nSOLO POMODORO\nMUTTI\n\n\nPOHODORO\n\n\nLa Vounts Mall\n\n\nMUTTI\n\n\nMUTTI\n\n\nPARMA\n\n\nPARMA\n\n\nPRODUCTION\n\n\nQ\nRassaila\nPar\u00e9e de tomates\nTomato puree-Passierte Tomaten\n\n\n100%\n\n\nO\nRassala\n\n\nPARMA\n\n\nITALIANI\n\n\nPassata\nTomato Pur\u00e9e\n\n\nnen PAS\n\n\nMatte\n\n\nDelicate & Velvety\n\n\nITE TUM\n\n\n100%\nITALIAND\n\n\nPROFESSIONAL\n\n\nPROFESSIONAL\n\n\nTomato puree-Passierte Tomaten\n\n\nSWEET & VELVETY\n\n\nMutti Passata is a dense, fragrant product\nmade using only tomatoes from the Emilia\nRomagna region that have passed Mutti's\nstrict quality tests. First processed to\neliminate the skin and seeds, the tomatoes\nare then pureed multiple times through\nenormous sieves. The end result is a smooth\nand velvety passata, rich and red in colour,\n\n\ndense in texture and fragrant in flavour.\nPassata is the quickest and easiest Mutti\nproduct to prepare as it only requires\nheating, thereby allowing its bright red\ncolour, velvety texture and sweet flavour to\nbe retained. Its rich texture also allows it to\nbe diluted without compromising its unique\ncharacteristics of fragrance and flavour.\n\n\nPACK\nSIZES\n2.5KG TIN\n700G BOTTLE\n400G BOTTLE\n5KG POUCH\n\n\n2x1\n\n\n2KG OF\nTOMATOES =\n1KG OF PASSATA\n\n\nDolce\n\n\n7.5\u00b0\n\n\nMINIMUM\nBRIX\n\n\nSWEET\nFLAVOUR,\nVELVET TEXTURE\n\n\n8\n\n\nNew Techniques\n\n\n1986\n\n\nNGROSSO\nROGHE\n\n\nGancia\n\n\nGascia\n\n\nPOLONIALI\n\n\nDROGHE\n\n\nMODORG\n\n\nMUTTI\n\n\nPARNA\n\n\nPassata\nTomato Pur\u00e9e\n\n\nNET\n\n\nDelicate & Velvety\n\n\nMUTTI TOMATOES ARE BOTTLED\n\n\nAs soon as technology made it possible to use glass in the food\npreservation industry, Mutti began using this material, which is not\nonly recyclable but also has good preservation properties. In a natural\nevolution from tins, Polpa was first bottled in 1986. Soon after, an\nentirely new bottled Passata range was created. The decision was\nmade by Marcello Mutti (b.1940) who took an active role in the\ncompany's development from 1965. Bottled passata became the catalyst\nfor the food preservation industry on the Italian market, and the two-lion\nbrand's quality ensured that it stood out on the supermarket shelf.\n\n\n9\n\n\nPizza Sauce\n\n\nPROVENANCE\n\n\nSOLO POHODORO\nMUTTI\n\n\nSOLO POMODORO\nMUTTI\n\n\nPARMA\n\n\nPARMA\n\n\nPIZZA SAUCE\nAROMATIZZATA\nSAUCE PIZZA AROMATIS\u00c9E\nSALSA PER PIZZA\n\n\nPIZZA SAUCE\nCLASSICA\nSAUCE TOMATE POUR PIZZA\nCLASSIC CLASSIQUE\n\n\nPROFESSIONAL\n\n\nPROFESSIONAL\n\n\nF Matte\n\n\nPRODITE\n00\n100%\nTEALBANO\n\n\nPRODUTO\n100%\nTTALIANO\n\n\nDORO\n\n\nDORO\n\n\n\u03c3\u03ba\u03bf\u03bb\n\n\nOWO\n\n\nTHE ART OF PIZZA\n\n\nMutti Pizza Sauce has no additives or\npreservatives and is produced from only fresh\ntomatoes harvested just a few hours prior to\nprocessing. It is rich flavour and dense in\ntexture and spreads uniformly over the pizza\ngiving just the right fragrance and colour to\nmake pizzas unique in aroma and taste.\nMutti Pizza Sauce is available in two\n\n\nflavours, Classica, and Aromatizzata, which\nis seasoned with traditional ingredients of\nonion, oregano and fresh basil. Due to its\ndense texture, Mutti Pizza Sauce is best used\ndiluted with water. This can be done without\nlosing the rich consistency, making it a high\nperforming product creating greater yield and\nefficiencies in commercial kitchens.\n\n\n30%\n\n\nPACK\nSIZES\n\n\nYIELD INCREASE.\nJUST ADD 1.5L\nOF WATER\n\n\n5KG POUCH\n4.1 KG TIN\n400G TIN\n\n\nDECISO\n\n\n10.5\u00b0\n\n\nRICH\nFLAVOUR.\nDENSE TEXTURE\n\n\nMINIMUM\nBRIX\n\n\n10\n\n\nPasta Sauce\n\n\nPROVENANCE\n\n\nSOLO\n\n\nPOMODORO\n\n\nMUTTI\n\n\nPARMA\n\n\nPASTA SAUCE\n\n\nQDA PRODUZION\n\n\nREADY TO USE Matti\n\n\n100%\nITALIANO\n\n\nTOMATO SAUCE\n\n\nPROFESSIONAL\n\n\nRICH\nTEXTURE WITH\nVIBRANT CHUNKS\nOF PERFECTLY RIPE\nTOMATOES\n\n\nTIME SAVER\nQuick to open\n\n\nEASY TO POUR\nEasy-pour handle\n\n\nLESS WASTE\nCrushable packaging\n\n\nFRESH & RICH\n\n\nMutti Pasta Sauce has no additives or\npreservatives and is produced with\nperfectly ripe, just harvested tomatoes and\nfresh basil. It has a rich texture and\ndelicate, fresh tomato and basil flavour -\na very balanced taste with no acidic flavours.\n\n\nIt has an intense red colour with\ndistinguishable leaves of fresh basil.\nMutti Pasta Sauce is ready to use -\nsimply stir through hot pasta. It also\nmakes the perfect base to your other\npasta sauces.\n\n\nPACK\nSIZES\n\n\n49\n\n\nTOMATOES\nIN EVERY\nPOUCH\n\n\n3KG POUCH\n\n\nFresco\n\n\n7.0\u00b0\n\n\nFRESH\nFLAVOUR,\nRICH TEXTURE\n\n\nMINIMUM\nBRIX\n\n\n11\n\n\nPackaging\n\n\nMORE\n\n\nLESS\n\n\nMODORD\n\n\nSOLO POMODORO.\nMUTTI\n\n\nMUTTI\n\n\nSOLO POMODORO\nMUTTI\n\n\nPARMA\n\n\nPOLPA\n\n\nPARMA\n\n\nPOMODORI IN FINISSIMI PEZZI\nTOMATES CONCASS\u00c9ES FINES\nFINELY CHOPPED TOMATOES\nFEINSTES TOMATENFRUCHTFLEISCH\n\n\nPARMA\n\n\nRassaila\n\n\nD\n\n\nPelati\n\n\nPROPT\n\n\nPROFESSIONAL\n\n\ngastronomia\n\n\nA PRODUDON\n100%\nITALIANO\n\n\nMUTTI\n\n\nMUTTI\n\n\nMatti\n\n\n100%\nITALIANO\n\n\nPOLPA\n\n\nPOLPA\n\n\nPROFESSIONAL\n\n\nNOVIT\u00c0 IN BUSTA\n\n\n100%\n\n\nSACLE CARE ACEAST\n\n\nPur\u00e9e de tomates\nTomato puree - Passierte Tomaten\n\n\n2 bestex 5 kg-10kg\n\n\n2\n\n\nCRUSHABLE PACKS\n\n\nTIME SAVER\nQuick to open\n\n\nLESS WASTE\nCrushable packaging\n\n\nEASY TO POUR\nEasy-pour handle\n\n\nA true design innovation, our Stand Up Pouch range helps maximise\nefficiencies in commercial kitchens. Made from a highly durable material\nwith a built-in handle for ease of use, the aseptic pouch allows the\nproduct to maintain its freshness once packed as well as traditional tin\nformats do. It's also a pack that is much more environmentally friendly,\nand waste is drastically reduced with the crushable design. Removing the\nneed to open tins also conserves precious food service preparation time.\n\n\n12\n\n\nInnovations\n\n\nTASTE\nWASTE\n\n\nSOLO POMODOR\nMUTTI\n\n\nSOLO POMODORO.\nMUTTI\n\n\nBOLD PONODONG\nMUTTI\n\n\nPARMA\n\n\nPIZZASAUCE\nCLASSICA\nSAUCE TOMATE POUR PIZZA\nCLASSIC CLASSIQUE\n\n\nPARMA\n\n\nPARMA\n\n\nPASTA SAUCE\n\n\nPIZZASAUCE\n\n\nAROMATIZZATA\nSAUCE PIZZA AROMATIS\u00c9E\nSALSA PER PIZZA\n\n\nQDA PRODUZION\n\n\nREADY TO USE Matti\nTOMATO SAUCE\n\n\n100\n\n\n100%\nITALIANOS\n\n\nMUTTI\n\n\nTI\n\n\nPROFESSIONAL\n\n\nMUTTI\n\n\nMall 100%\n\n\nLASAUCE PIZZASAUCE\n\n\nTRALIANO\n\n\nCLASSICA\n\n\n100%\n\n\nRICH\nTEXTURE WITH\nVIBRANT CHUNKS\nOF PERFECTLY RIPE\nTOMATOES!\n\n\ntex5 kg-10bg\n\n\n2 botex5bg-g\n\n\nPRODUCTS AND SIZES\n\n\nPOLPA FINELY\nCHOPPED\nTOMATOES\n\n\nPELATI\nPEELED\nTOMATOES\n\n\nPIZZA SAUCE PIZZA SAUCE\nCLASSICA\n\n\nPASSATA\n\n\nPASTA\nSAUCE\n\n\nTOMATO\n\n\nAROMATIZZATA\n\n\nPUREE\n\n\n3 x 3kg\nPOUCHES\n\n\n2 x 5kg\nPOUCHES\n\n\n2 x 5kg\nPOUCHES\n\n\n4 x 2.3kg\nPOUCHES\n\n\n3 x 3kg\nPOUCHES\n\n\n2 x 5kg\nPOUCHES\n\n\nNo additives or\npreservatives, only\nfresh tomatoes\nflavoured with\nnatural ingredients.\n\n\nPerfectly ripe, just\nharvested tomatoes,\ncold crushed for a\ntrue fresh tomato\ntaste.\n\n\nProduced from just\nharvested, perfectly\nripe tomatoes, rich in\nflavour and dense in\n\n\nOur 100% Italian\nPeeled Tomatoes\nsimply pureed to a\nvelvety smooth\n\n\nA dense, fragrant and\nvelvety puree of just\nharvested tomatoes,\nwith the skins and\nseeds removed.\n\n\nReady to use pasta\nsauce made with\nperfectly ripe\ntomatoes. Simply\nheat and serve.\n\n\ntexture.\n\n\nsauce.\n\n\n13\n\n\nCherry Tomatoes\n\n\nPROVENANCE\n\n\nLa famiglia Ma\nUna passin\n\n\nSOLO\n\n\nPOMODORO\n\n\nMUTTI\n\n\ne scelli\nnude maturi\nfer letto l'anno\npre ace\n\n\ndalla sel\n\n\nche an\n\n\nPARMA\n\n\nginn\n\n\nPRODUZION\n\n\nCERTIFI\nITALIANIS\n\n\nMatti\n\n\n100%\n\n\nPROFESSIONAL\n\n\nCiliegini\n\n\nPOMODORINI\n\n\nCiliegini\n\n\nC\n\n\nleal\n\n\nSMALL & SUCCULENT\n\n\nRenowned for being extremely sweet with an\nintense, fragrant flavour, Mutti Cherry Tomatoes\nare ripened under the warm sun of southern\nItaly and are carefully selected to offer the\nfreshness of just-harvested fruit all year round.\n\n\nThey are soft and succulent yet maintain their\nform well. As the skin of Mutti Cherry Tomatoes\nhas a thin texture, they are ideal to make\nquick, delicious sauces or even to use with\nfine fish or meat.\n\n\nPACK\nSIZE\n\n\n150\n\n\nFRUITS\nPER TIN\nMINIMUM\n\n\n2.5KG TIN\n400G TIN\n\n\n6.0\u00b0 Dolce\n\n\nMINIMUM\nBRIX\n\n\nSWEET\nFLAVOUR\n\n\n14\n\n\nPeeled Tomatoes\n\n\nPROVENANCE\n\n\nSOLO POMODORO.\nMUTTI\n\n\nSOLO POMODORO.\nMUTTI\n\n\nLa fami\n\n\ng\u00e9rateur\nSeekat\n\n\nUna\n\n\n3 Tage\n\n\ndall\n\n\nPa\n\n\nPARMA\n\n\nTyto-take\n\n\nPARMA\nPelati\n\n\nPelati\n\n\nta\n\n\nL\n\n\ngastronomia\n\n\nFMatter\n\n\ngastronomia\n\n\nPRODUZIO\n100%\nITALIANO\n\n\nPilali\n\n\nPROFESSIONAL\n\n\nPele Pilate Pilate Pilate Pi\n\n\nQE PRODUZIONE\n\n\nFMatt\nPROFESSIONAL\n\n\nNOVIT\u00c0: IN BUSTA\n\n\n100%\nITALIANIS\nPelo Pilat\n\n\nMESGEVOLE\n\n\nFACE BA APRIRE FACILE DASMALTRE\n\n\nAlale P\n\n\nFLESHY & FIRM\n\n\nGenerously fleshy and firm, Mutti Selezione\nGastronomia Peeled Tomatoes have their\nskins removed and are immersed in a velvety\npassata. Processing and packaging all occur\nwithin 24 hours of harvesting to preserve the\nauthentic taste of freshly picked tomatoes.\n\n\nWith a higher than average Brix, the product\nis noticeably dense and rich in texture upon\nopening and holds its flavour, richness and\ncolour for longer cooking times. The density\nalso allows for a greater yield with lower than\naverage amounts required per portion.\n\n\nFRUITS\nPER\nPACK\nTIN - 25\nPOUCH - 23\n\n\nPACK\nSIZE\n\n\n2.5KG TIN\n2.3KG POUCH\n\n\nDECISO\n\n\n5.7\u00b0\n\n\nMINIMUM\nBRIX\n\n\nBOLD, RICH\nFLAVOUR\n\n\n15\n\n\nPeeled San Marzano Tomatoes\n\n\nPROVENANCE\n\n\nD'ORIGINE\n\n\nAZIONE\n\n\nENOMIN\n\n\nTETTA\n\n\nF. Muttin\n\n\nS.MARZANO\n\n\nPOMODORO\n\n\nche ne affed\n\n\nSOLO POMODORO\nMUTTI\n\n\nDELL'AGRO SA\n\n\nD.O.P.\n\n\nerin\n\n\nPARMA\n\n\nVOCERINO\n\n\nSARNES\n\n\n2500g\n2650m\n\n\nRACCOLTA 2015\n\n\nDELICIOUSLY RICH\n\n\nMutti San Marzano Tomatoes are grown in\n\n\nare a vibrant red colour and have a dense\nand even texture. They are preserved in a\nvelvety and delicious passata and are\nprocessed very quickly in order to retain the\nfreshness and taste of freshly picked fruit.\n\n\nAgro Sarnese Nocerino, between Naples\nand Salerno. They benefit from Protected\nDesignation of Origin status (PDO), which is\na testament to their unique character and\nhigh quality. These delicate, fleshy tomatoes\n\n\nPACK\nSIZES\n\n\n25\n\n\nFRUITS\nPER TIN\nMINIMUM\n\n\n2.5KG TIN\n400G TIN\n\n\n85%\n\n\n5.5\u00b0\n\n\nUNBROKEN\n\u03a4\u039f\u039c\u0391\u03a4\u039f\nFRUIT\n\n\nMINIMUM\nBRIX\n\n\n16\n\n\nOur Ingredients\n\n\nD\n\n\n50\n\n\nWATER\nTHE SOURCE OF\nEXCEPTIONAL TASTE\n\n\nEARTH\n\n\nWHERE OUR PLANTS\n\n\nTAKE ROOT\n\n\nHumans have interacted with the earth since\n\n\nWater gives our tomatoes their wealth of\nflavour and helps us take care of the fruit,\neven after harvest. When the tomatoes arrive\nat the factory they are sorted and washed\nwith gentle water jets so as not to damage\nthem. Water is a precious resource, and we\ndon't like wasting it. Season after season, we\nlimit our consumption and streamline its use.\n\n\nthe dawn of time. It is a relationship built\nupon respect. We also respect our farmers\nand know each of them by name, because\nthey are members of our family. We are\n\n\nhonest and open with them and they know\n\n\nthat passion and devotion are the hallmarks\nof Mutti suppliers.\n\n\nMUTTI'S\nRECIPE FOR\nSUCCESS\nIN YOUR\nKITCHEN\n\n\nSUN\n\n\nTIME\n\n\nPATIENCE TO WAIT FOR\nTHE PERFECT HARVEST\n\n\nTHE WARMTH THAT\nBRINGS SWEETNESS\n\n\nSun and rain are our main allies in bringing\nour tomatoes to the perfect ripeness. There is\nnothing better than a hot summer on the\nplains for tomatoes that are full of flavour.\nThe fruit is ready to eat once it is bright red\n\n\nChoosing the right time to pick tomatoes means\nfinding the perfect balance: neither too early nor\n\n\ntoo late. In this art form, perfection and\npunctuality go hand in hand. Our farmers\nmonitor their tomato fields with the utmost care\nand attention, looking out for signs of optimum\nripeness. The quality of the harvest, which is\nrewarded with a warm handshake, is the result\nof our constant commitment from field to fork.\n\n\nin colour. This is when its nutritional benefits\nare the most highly concentrated. So the sun\n\n\nis Mutti's greatest ally.\n\n\n17\n\n\nThe Five Key Stages\n\n\nIf you need one more reason to choose Mutti, look no further than its exceptional taste.\nThis stems from our responsible approach to nature and our desire to foster healthy living for all.\nEach stage is designed with a single purpose: to retain all of the tomato's natural goodness.\n\n\nMaintaining a constant dialogue with our farmers means that\nwe can choose the varieties that are best suited to each plot of\nland, and produce the finest tomatoes.\n\n\nPlanting\n\n\nOur goal is to know all there is to know about the tomatoes.\nWe share all of our experience with our farmers and guide them\nthroughout the development of their crop.\n\n\nFarming\n\n\nKnowing how to choose the right time to harvest is the key to\nachieving the best flavours. During the tomato harvest itself,\nour factories are open around the clock, because ripe tomatoes\nsimply cannot wait. The fruit that we process at Mutti\nheadquarters are all grown within an average radius of just\n100 kilometres.\n\n\nHarvest\n\n\nThere is no greater satisfaction than receiving a batch of Mutti\ntomatoes ready for processing. It means that they have passed\nour stringent quality control tests, that they have reached\noptimal ripeness and that in just a few hours, they will\nbecome the products that we use every day in our kitchens.\n\n\nSelection\n\n\nWe limit the time between the tomatoes' arrival at the factory and\nthe point at which they are then processed. Using innovative\ntechnology to sort our tomatoes with the greatest care, production\ntime is cut to a minimum and heat treatments mean that the\ntomatoes are handled more gently.\n\n\nProduction\n\n\nThe Golden Tomato\n\n\nA QUEST FOR PERFECTION\n\n\n\"THIS AWARD HELPS TO STRENGTHEN THE MUTUAL TRUST\nBETWEEN OURSELVES AND OUR FARMERS AND REFLECTS THE\nCARE WE TAKE TO SATISFY THOSE WHO CHOOSE US EACH DAY.\"\n-Francesco Mutti\n\n\nWe do everything we can to bring you the best\n\n\nquality products. That's why every year Mutti\npresents the Pomodorino d'oro (Golden Tomato)\nquality award to the best producer. Year on\nyear this award allows us to constantly improve\nthe quality we deliver. In the end, it's a win-win\nsituation: Mutti, our farmers and the customer\nall benefit from this award.\n\n\nMUTTI\n\n\n19\n\n\nRespect for our Land\n\n\nWATER FOOTPRINT\n\n\nSustainable water use is important to us. In 2010 we partnered with WFF Italy and became the\nfirst Italian company to calculate our water footprint - which is the amount of water consumed\nacross the entire production process - and take concrete action to reduce it.\n\n\nAMOUNT REQUIRED TO MAKE\n\n\n500,000,000\n\n\n625,000,000\n\n\nPASTA DISHES\n\n\nLITRES OF WATER*\n\n\nINSIEME A MUTTI\nPER RIDURRE\n\n\nL'IMPATTO\nAMBIENTALE\n\n\nWWF\n\n\nENERGY FOOTPRINT\n\n\nAs well as water resource consumption, we have been working hand in hand with WFF Italy to help farmers\nfind sustainable ways of reducing CO\u2082 emissions. We've invested in technology, training and technical\nhelp for farmers and producers, which help us to contribute to the development of our land.\n\n\nCO\n\n\n2\n\n\n56,385\n\n\n12,400\n\n\nTONNES OF CO2 IN THE AIR**\n\n\nMILAN-LONDON FLIGHTS\n\n\n*Litres of water saved between 2012 and 2014.\n(Data compiled by WFF Italy and the Euro-Mediterranean Centre for Climate Change, CMCC).\n\n\n** CO\u2082 emissions prevented between 2010 and 2014, in comparison with the 2009 performance level.\n(Data compiled by WFF Italy and the energy efficiency group Officinae Verdi).\n\n\nDefining Sustainability\n\n\n\u201cIT'S A\nPARTNERSHIP\nTHAT\nSTRENGTHENS\nTHE BONDS\nBETWEEN\nEVERYONE\nINVOLVED.\"\n\n\nThe company encourages constant dialogue between\ninstitutions - the research centres and universities\ninvolved alongside WFF Italy - farmers, canning\nfactories and packaging companies. The aim is to\ncreate a solid foundation of shared values to\noptimise each partner's contribution and translate\nthe principle of respect for the land into tangible\nactions. This is how we define sustainability: a set of\nchoices that give meaning to the word \"taste\".\nIt is not simply a question of economic sustainability,\nwithout which no company can survive, but also of\nenvironmental and social sustainability. When you\nbuy a Mutti product, you are not only choosing a high\nquality ingredient, you are making a much more\nmeaningful gesture.\n\n\nProduct Overview\n\n\nPRODUCT\nDESCRIPTION\n\n\nPALLET CARTON\nQUANTITY\n\n\nPRODUCT\nSIZE\n\n\nPRODUCT\nCODE\n\n\nLAYER CARTON\nQUANTITY\n\n\nMUTTI\n\n\nPolpa Finely Chopped Tomatoes\n\n\n12 x 400g\n\n\n160\n\n\nMPOLPA400(12)\n\n\n16\n\n\nPOLPA\n\n\nHUTTI\n\n\n160\n10\n\n\nMutti Pizza Sauce\n'Aromatizzata' Herbs & Spices\n\n\nMPSAU400(12)\n\n\n12 x 400g\n\n\nPZZASAUCE\n\n\n16\n\n\nMUTTI\n\n\n160\n10\n\n\nCherry\n\n\nCherry Tomatoes\n\n\n12 x 400g\n\n\nMCT400(12)\n\n\n16\n\n\n20\n20\n\n\n12 x 400g\n\n\nMPASS400(12)\n\n\nPassata\n\n\n100\n\n\nMUTTI\n\n\n80\n50\n\n\nPassata\n\n\nMPASS700(12)\n\n\n12 x 700g\n\n\n16\n\n\nMUTTI\n\n\n64\nHO\n\n\n6 x A9 (2.5kg)\n\n\nMutti San Marzano Peeled Tomatoes\n\n\nMPEELSMA9(6)\n\n\n8\n\n\nMUTTI\n\n\nMUTTI\n\n\n112\n12\n\n\nMPASSA9(3)\n\n\nMutti Passata (Tomato Puree)\n\n\n3x A9 (2.5kg)\n\n\n16\n\n\nMUTTI\nPelati\n\n\nMutti Pelati (Peeled) Tomatoes\n\n\n6 x A9 (2.5kg)\n\n\nMPEELA9\n\n\n8\n\n\n64\n\n\nMUTTI\n\n\nPIZZA SAUCE\n\n\n80\n60\n\n\nMPSAUCLASA12(3)\n\n\n3x A12 (4.1kg)\n\n\nMutti Pizza Sauce Classica\n\n\n16\n\n\nMUTTI\n\n\nPIZZASAUCE\n\n\n80\n50\n\n\n3x A12 (4.1kg)\n\n\nMPSAUAROMA12(3)\n\n\nMutti Pizza Sauce Aromatizzata\n\n\n16\n\n\n22\n\n\nPRODUCT\nSIZE\n\n\nPRODUCT\nCODE\n\n\nPRODUCT\nDESCRIPTION\n\n\nLAYER CARTON\nQUANTITY\n\n\nPALLET CARTON\nQUANTITY\n\n\nMUTTI\n\n\nMutti Polpa\nFinely Chopped Tomatoes\n\n\n3x A12 (4.05kg)\n\n\nMPOLPA12(3)\n\n\n16\n\n\n80\n\n\nPOLPA\n\n\nMUTTI\n\n\nMutti Tomato Paste\n\n\n112\n12\n\n\n3 x A10 (2.15kg)\n\n\n16\n\n\nMTOMPA10\n\n\nMUTTI\n\n\n6 x A9 (2.5kg)\n\n\nMCTA9(6)\n\n\n8\n\n\nMutti Cherry Tomatoes 6xA9\n\n\n64\n\n\nCiliegini\n\n\nMUTTI\n\n\nMutti Pizza Sauce\nAromatizzata Bag in Box\n\n\nPIZZASRICE\n\n\n75\n75\n\n\n25\n\n\nMPSAUAROMA(2X5KG)\n\n\n2 x 5kg (BIB)\n\n\nMUTTI\n\n\n25\n25\n\n\nMutti Pizza Sauce\nClassica Bag in Box\n\n\nMEERSAUCE\n\n\nMPSAU(2X5KG)\n\n\n2 x 5kg (BIB)\n\n\n75\n\n\nHUTTI\nPASTA SAUCE\n\n\nMPASTASAU(3X3KG)\n\n\nMutti Pasta Sauce\nBag in Box\n\n\n16\n\n\n3 x 3kg (BIB)\n\n\n64\n\n\nMUTTI\n\n\nMutti Polpa Finely Chopped Tomatoes\nBag in Box\n\n\nPOLPA\n\n\nMPOLP10KG\n\n\n2 x 5kg (BIB)\n\n\n25\n\n\n75\n\n\nMUTTI\n\n\n48\n42\n\n\nPolak\n\n\nMutti Peeled Tomatoes\nBag in Box\n\n\n4 x 2.3kg (BIB)\n\n\nMPEEL4X2.3KG\n\n\n8\n\n\nMUTTI\n\n\n64\nTO\n\n\nMutti Passata (Tomato Puree)\nBag in Box\n\n\n16\n\n\n3 x 3kg (BIB)\n\n\nMPASS(3X3KG)\n\n\n23\n\n\nCertificate of Guarantee\n\n\nA FOCUS ON QUALITY\n\n\nQuality, transparency and reliability are the values\nwhich underpin everything that we do.\n\n\nIn 1999, Mutti was the first company to obtain the Integrated Production\nCertificate according to the UNI 11233 standard. This certification of the\nagricultural production process provides for very strict controls from cultivation\nthrough to harvest, further reinforcing our 100% dedication to quality\nfor every single product that bears our name.\n\n\nIn 2001, based on the ISO 22005 standard, Mutti obtained certification on the\ntraceability of the supply chain, from the seed to the finished product, which\nallows us to certify the total absence of GMOs and the use of only\n100% Italian grown tomatoes.\n\n\nE\n\n\nTRACEABILITY TO\nGUARANTEE ITALIAN\nORIGIN - ISO 22005\nCERTIFICATION\n\n\nISO 14001\nCERTIFIED\n\n\nFOOD SAFETY\nBRC CERTIFIED\n\n\nQUALITY\nMANAGEMENT\nSYSTEM ISO 9001\nCERTIFICATION\n\n\nFOOD SAFETY\nIFS CERTIFIED\n\n\nINTEGRATED\nPRODUCTION UNI\n11233 CERTIFIED\n\n\nOCCUPATIONAL\nHEALTH AND SAFETY\nMANAGEMENT\nSYSTEM\n\n\nQUALITY\nCONTROLLED REGION\nEMILIA ROMAGNA QC\nCERTIFICATE\n\n\n24\n\n\nSOLO POMODORO.\nMUTTI\n\n\nPARMA\n\n\nCONTACT US\n\n\nAustralian Distributor - Sandhurst Fine Foods Coy\n1800 500 362 | Info@sandhurstfinefoods.com.au\n\n\nwww.muttiparma.com.au | info@muttiparma.com.au\nMUTTI AUSTRALIA, SUITE 7, 87-103 EPSOM ROAD, ROSEBERY, NSW 2018\n1800 940 607\n"}, "expected_output": {"claims": [{"unit": "kg fresh tomatoes per kg polpa", "value": 5, "evidence": ["5KG OF\nTOMATOES = 1KG\nOF POLPA\n"]}]}, "metadata": {"product_category": "Food & beverages", "request_id": "req_4ccdef38340fe4fe"}}
{"id": "822758a235fa68220d5294b7", "input": {"query": "What is the hydrogen supply ratio (alpha) defined as the ratio of hydrogen supplied to stoichiometric requirements in continuous catalytic hydrogenation reactors? Extract numerical values or ranges for typical hydrogen excess ratios used in practice.", "source_url": "https://ris.utwente.nl/ws/files/6485238/Westerterp88development.pdf", "document_text": "Chemical Engineering Science, Vol. 43, No. 8, pp. 2229\u20132236, 1988.\nPrinted in Great Britain.\n0009-2509/88 $3.00+0.00\nPergamon Press plc\nDEVELOPMENT OF CATALYTIC HYDROGENATION REACTORS FOR THE FINE CHEMICALS INDUSTRY\nK.R. Westerterp, K.B. van Gelder, H.J. Janssen and M.H. Oyevaar\nChemical Reaction Engineering Laboratories\nChemical Engineering Department\nTwente University\nP.O. Box 217, 7500 AE Enschede\nThe Netherlands\nAbstract\nA survey is given of the problems to be solved before catalytic hydrogenation reactors can be\napplied in a multiproduct plant in which selectivity problems are experienced. Some results are\nreported on work done on the reaction kinetics of two multistep model reactions and on\nmathematical modelling and experimental verification of the models. Since hydrogenation reactions\nare often very exothermic, cooling by solvent evaporation has been applied where appropriate.\nSufficient information has been collected and correlated to enable operation of multiproduct\ncatalytic reactors of the slurry or packed bubble column type; interdependence of operating\nvariables is so complex that a mathematical model is indispensable.\n1. Introduction\nIn our laboratories cooperation with industry is intensive. It is our job to study the basic\nfundamentals for new operations or processes, which will be further elaborated by our partners and\neventually developped on a larger scale into new processes and syntheses. This approach places us\nin a position to publish on our results. Since 1984 we are a.o. engaged in developping catalytic\nhydrogenation processes for the fine chemicals industry. In this undertaking we have a partner who\nhas a production program which includes around twenty different hydrogenation reactions of a\nvariety of raw materials. These hydrogenations are carried out according to the classical B\u00e9champs\nprocess using Fe-powder and a HCL solution. As a consequence large amounts of iron hydroxides and\nneutralization salts have to be disposed of, a very costly affair. So it was obvious to look into\npossibilities for catalytic hydrogenation.\nIn catalytic hydrogenation a number of problems can be anticipated. It is too costly to build say\n20 different small hydrogenation units if these reactions have to be carried out under different\noperating conditions. As a consequence one or two larger units have to be developped for\nmultiproduct use. Hydrogenation reactions often are highly exothermic - in our case adiabatic\ntemperature rises of 200-500 K are common - so methods have to be devised to cope with high heat\nproduction rates. In case of producing an intermediate product in a set of consecutive reactions\nor of combined reactions selectivity is a problem and operating conditions must be kept under\naccurate control. The number of catalysts to be used preferably should be small because of\nhandling and recovery costs. It would be best to have one single, universal catalyst.\nFinancial losses should be kept to a minimum. For raw materials this requires short and quick\nstart-up and stop procedures, for hydrogen it requires low excess of hydrogen or recirculation of\nnon converted material. For the catalyst it demands a good recovery system and in case of\ndeactivation simple reactivation procedures. Therefore, a number of problems has to be solved\nbefore catalytic hydrogenations can be introduced on a large scale in a fine chemicals plant.\n2. Development objectives\nIn order to solve the problems at hand we decided to opt for continuous operation of the reactor\nand to use an evaporating solvent to remove the reaction heat in the reactor. Miniplants have been\nbuilt for catalytic hydrogenations in a slurry reactor, in a packed bubble column and in a\ncocurrent trickle flow column reactor. In addition, two separate miniplants have been build for\nthe study of gas-liquid interfacial areas in several different reactor types, one for pressures in\nthe range of 0.1-2 MPa and the other one for 0.1-10 MPa. To test the performance of the reactors\nexperimentally two model hydrogenation reactions have been selected and the reaction rates of the\nindividual reaction steps have been determined. Mathematical models have been developed to\ndescribe the different reactor types and their behaviour. The models have been tested\nexperimentally. Interfacial areas are determined by a chemical method using systems of amines and\nCO2-gas. In this communication we report briefly on some of the results we have obtained. Some\nfifteen papers are being prepared to report more extensively on our results.\n3. Model reactions\nHaving considered many possible complex reactions we decided upon the catalytic hydrogenation of\n2,4-dinitrotoluene (DNT) and of 2,4,6-trinitrotoluene (TNT) in methanol and as our model\n2229\n2230\nK. R. WESTERTERP et al.\nG12\nreactions, and supported palladium as the catalyst. The heat of reaction of the reduction of a\nsingle nitrogroup is around 600 kJ/mole; at reasonable concentrations of TNT adiabatic temperature\nrises of 300 to 500 K occur. On theoretical grounds we can expect that the hydrogenation of DNT\nproceeds in two parallel paths, each of which consist of two consecutive reactions when\namino-nitro compounds are formed as the sole intermediates. The hydrogenation of TNT involves an\neven more complicated scheme.\nA large amount of literature has been published in recent years concerning the experimental\ninvestigation and the modelling of the catalytic hydrogenation of several nitrosubstituted\naromatic compounds. Most of the literature refers to the hydrogenation of nitrobenzene. In\ngeneral, aromatic nitrocompounds are converted to the amines quantitatively. Commonly Ni and\nsupported Pd or Pt are used as catalysts. In general the reaction rates reported are zero order in\nthe nitrocompound and between zero and first order in hydrogen. They can be described with\nLangmuir-Hinselwood rate equations. Burge et al. (1980) reported that the hydrogenation of\nnitrobenzene, using Raney nickel as a catalyst, did involve the formation of azoxy- and of\nazobenzene as stable intermediates. The formation of azoxybenzene as an intermediate has also been\nreported by Yao and Emmet (1962). Burge concluded that the reaction mechanism followed a reaction\npathway similar to that proposed by Haber (1898) for the electrochemical reduction of\nnitrobenzene. Collins et al. (1982) reported that the reaction pathway of the reduction of\nnitrobenzene also obeyed Haber's scheme when nickel boride was used as a catalyst. The maximum\namounts of azoxy- and of azobenzene which builded up during the reaction were found to be\nnegligable. Bird and Thompson (1980) stated that the reduction of any nitrogroup on an aromatic\nring, using supported palladium as a catalyst, will follow the same pathway as chemical reduction,\nbut that it is very difficult to terminate the reaction at one of its intermediates. Yucelen\n(1984) studied the kinetics of the hydrogenation of 2,6-dinitrotoluene using Pt/C and Pt/Al.\ncatalysts. The reaction proceeded according to a simple scheme of two consecutive reaction A \u2192 B \u2192\nC, in which B represents the only intermediate 2-amino-6-nitrotoluene and where C is the\nend-product 2,6-diaminotoluene.\n11203 as\nAlthough some studies have been published with respect to the catalytic hydrogenation of 2,4-DNT,\npracticable models of this reaction and of the hydrogenation of TNT have not yet been developed.\nAcres and Cooper (1972) and Bird and Thompson (1980) reported about the accurrence of two\nintermediates in half-hydrogenated samples of 2,4-DNT: 2-amino-4-nitrotoluene and 4-amino-2-\nnitrotoluene, respectively. Unfortunately they did not give any further information, neither on\nthe change of the concentrations of the substrate and the products as a function of the course of\nthe reaction, nor on the analytical technique used. Pawlowski and Kricsfalussy (1981) found a\nsingle step mechanism for the conversion of 2,4-DNT to 2,4-DAT. Acres and Cooper (1972) have found\n2-amino-4-nitrotoluene (2-A-4-NT) and 4-amino-2-nitrotoluene (4-A-2-NT) as stable intermediates in\nhalf hydrogenated samples of DNT.\nThe kinetics of the hydrogenations of DNT and TNT have been investigated in a fully automated\nbatch slurry reactor, which was operated isothermally and under constant pressure. As a catalyst\n5% Pd on active carbon was used. The consumption of hydrogen during reaction was calculated from\nthe drop in a hydrogen pressure. A set of preliminary experiments has been performed at several\ntemperatures, agitation rates, catalyst and substrate concentrations to warrant the exclusion of\nany mass transfer effects that possibly mask the real reaction rates. For both DNT and TNT sets of\nhydrogenation experiments have been carried out at different hydrogen pressures to determine the\neffect of the temperature and hydrogen pressure on the overall conversion rates. Routine analysis\nwere performed with a gas chromatograph which is equipped with a flame ionization detector and\nfitted with 6 ft., 1/8 inch column packed with Tenax-GC, 60-80 mesh. So far we have concentrated\nmainly on quantitative analyses of reaction mixtures of DNT. The concentrations of the relevant\nreaction compounds as a function of time have been determined at reactor temperatures of 308, 323\nand 357 K and at a constant hydrogen pressure of 2 MPa. In the samples DNT, 2-A-4-NT, 4-A-2-NT and\nDAT have been identified by means of comparison with the retention times of the pure compounds.\nTwo other products have been identified as 4-hydroxylamino-2-nitro-toluene (2-N-4-HAT) and 2-\namino-4-hydroxylamino-toluene (2-A-4-HAT) by means of GC-MS analysis. Note that the hydroxylamines\nare formed at the para-position only and that an ortho-nitrogroup is converted directly to an\naminogroup. The concentrations of 2-A-4-HAT were found to be so low that the pressure of this\ncompound has been neglected.\nA plot of the hydrogen conversion as a function of the reaction time during a typical hydrogena-\ntion of DNT at 308 K is presented in Fig. 1. In Fig. 2 the concentrations of the major compounds\nare plotted versus the hydrogen conversion. In Fig. 2 also the reaction pathway can be observed.\nAt first DNT is converted in two parallel reactions into 2-A-4-NT and 2N-4HAT. The consecutive\nreactions do not start before almost all of the DNT has been consumed. This is known as molecular\nqueueing and can be explained in terms of a competitive adsorption of the reacting species on the\ncatalyst surface. Once all of the DNT has been consumed the 2-N-4-HAT is converted into the amine\n4-A-2-NT which subsequently reacts to DAT, while in the parallel pathway simultaneously the\n2-A-4-NT is converted to DAT. In Fig. 1 a sharp decrease is shown in the overall reaction rate to\na conversion level of 0.4.\nAs a consequence of the strong molecular queueing in the initial stage\nthe reaction exhibits zero order behaviour with respect to DNT. The observed decrease of the\noverall reaction rate was caused by rapid deactivation of the catalyst in an induction period.\nThe kinetic rate data of the hydrogenation of DNT and TNT have been modelled using Langmuir-\nHinshelwood rate equations. The rate of conversion of species i into species j is given by rij\nccat Kij Hi; where OH and i are the relative concentrations on the catalyst surface of hydrogen\nand of species i respectively. The experimental results indicate separate active sites for\nhydrogen and for the aromatic molecules on the catalyst surface. Assuming dissociative adsorption\nof hydrogen on the catalyst O\u0434 and O\u2081 are given by:\n\u2713(\u043aHPH2/RT)\n=\n\u043e\u043d +(PRT)\nand\nQi Xi\n0\ni\n(KACAO \u00af +QjXj\nG12\nDevelopment of catalytic hydrogenation reactors\n2231\nKA/C\nis\nwhere \u043a is the product of the solubility constant of hydrogen in methanol and the true adsorption\nconstant, Q\u2081 is the relative adsorption strength of species i with respect to DNT or TNT,\nthe adsorption constant, CAO is the initial concentration of DNT or TNT and further Xi\nThe model for the hydrogenation of DNT is based on the reaction pathway presented in Fig.\nwas found that the overall hydrogen consumption rate of the hydrogenation of TNT could be\ndescribed accurately over the whole temperature range of the experiments by a lumped model of\nthree consecutive reactions TNT DNHAT DANT\u2192TAT. Unfortunately we lacked sufficient quantitative\ndata of the liquid phase concentrations of the reaction components to set up a realistic extended\nmodel for the hydrogenation of TNT.\n1.00\n1.00\n4HA-2NT \u2014 4A-2NT\nDNT\nDAT\n2A-4NT\n0.80\n0.80\n0.60\nhydrogen\nconversion\n0.40\n0.20\nhydrogen pressure: 1.90 MPa\nreactor temperature: 308 K\ncatalyst concentration: 1.203 kg/m\u00b3\ninitial DNT concentration: 0.220 kmole/m\n0.60\nCiCo\n0.40\n0.20\n0.00\n0.00\n0\n1000\n2000\n3000\n4000\n5000 6000\n0.00\n0.20\n0.40\n0.60\n0.80\n1.00\nreaction time\nhydrogen\nconversion\nFigure 1:\nHydrogen conversion vs. reaction time for a typical hydrogenation\nof DNT in a mechanically agitated three-phase batch reactor.\nFigure 2:\nRelative concentration C/C vs. hydrogen consumption for a typical\nhydrogenation experiment of DNT (reaction conditions see figure 1).\nSymbols: DNT, 4HA-2NT, \u2666 4A-2NT, 2A-4NT, \u25a1 DAT.\nFor the parameter estimation of both models we used the RKPES computer program developed by Klaus\nand Rippin (Klaus and Rippin, 1979) which applies the maximum likelihood approach. The parameters\nhave been estimated for each experimental temperature separately. The temperature dependent\nrelations for the parameters have been derived from the single temperature results. The relations\nare summarized in Table 1. The data for DNT comprised the data of the hydrogen conversion as a\nfunction of the reaction time of three different experiments (100 to 120 points) in addition to th\ndata of the concentrations of the reaction components as a function of the reaction time (20 to 25\nsets of 5 concentrations) for each temperature. For TNT only data of the hydrogen conversion as a\nfunction of the reaction time have been used. The t-values of the estimated rate parameters range\nup from around 10 to 60 whereas the t-values of the adsorption parameters are poor. This can be\nexplained by the large difference in the relative adsorption strength of the reacting\ncomponents on the catalyst: when a component i is strongly adsorbed and nearly covers the whole\ncatalyst surface the relative surface concentration i 1 and has thus become insensitive to the\nexact\nvalue of the adsorption constant.\nTable 1: Parameter values in the rate equations for the hydrogenation of DNT and TNT.\n==\nsubstrate and\nreaction constants\nadsorption parameters\nreaction pathway\npreexponential\nfactor\n[mole/kg cat s]\nactivation\ntemperature\n(K)\npreexponential\nfactor\nactivation\ntemperature\n(K)\nDNT\n12\n*23\n1.31 10\n5.14 1013\n10\n7500\nK1\n10000\n2.\nk34\n2.60 10\u00ba\n7400\n*15\n2.13 1011\n8800\n\u17e9\u17e9\u17e9\n0.15\n130\n[m\u00b3/mole]\n-3000\nvapour\nline\ncondensor\n[-]\n2500\n7.11\n[-]\n1100\ngas out\ncondensate returned\nQ4\n0.011\n(-]\n-320\nk\u22644\n7.17 1013\n10200\nQs\n0.4\n[-]\n660\nreactant feed\ncatalyst feed\nhydrogen feed\nstirred\ntank\nreactor\nproduct out\nTNT\n*12\n3.3 1011\n8800\nK\u2081\n-10\n3.8 10\n1 2 3 4\n*23\nK34\n1.2 1010\n[m/mole] -9100\n8100\nQ2\n1600\n(-)\n1.8 10.10\n8100\nQ3 1.8 10\n(-)\n3500\n7200\nFigure 3:\nOutline of a continuous stirred three-phase tank\nreactor with an evaporating solvent.\nFor both systems: K\u2081\u2013 9.14 10\u00b98\n\u00b0exp (10360/T) [m\u00b3/mole]\n4. Mathematical modelling and experimental confirmation\n4.1 Continuous stirred three-phase_slurry_reactor\nA series of catalytic hydrogenation experiments has been carried out in a mini-installation with a\nthree-phase slurry reactor using an evaporating solvent to control the temperature. In Fig. 3 the\nbasic line-up of the installation is presented. A solution of the reactant, a concentrated slurry\nof the catalyst and hydrogen gas are fed to the reactor separately. We assume that the liquid and\nthe gas phase in the reactor are well mixed and that the gas phase is saturated with solvent\nvapour. The reaction products and the catalyst leave the reactor in the product stream. The excess\nof non-converted hydrogen gas leaves the reactor via the vapour line to a condensor and carries\naway the evaporated solvent, thus withdrawing heat from the reactor. The solvent is condensed and\nreturned to the reactor: the excess gas leaves the system via a back-pressure control\nvalve.\nAt constant reactor pressure and constant hydrogen excess the heat withdrawn from the reactor by\nsolvent evaporation will rise strongly with increasing temperature because of the exponential\n2232\nK. R. WESTERTERP et al.\nG12\nincrease of the rate of evaporation. The maximum attainable reactor temperature, the boiling point\nof the reaction mixture, can never be reached, because the rate of evaporation would become\ninfinitely large. The rate of heat production is also affected by the temperature because the\npartial pressure of hydrogen is lowered by the increase of the solvent vapour pressure. At a\ncertain temperature the rate reducing influence of a lower partial pressure of hydrogen will\nbecome larger than the rate enhancing effect of a higher temperature. Consequently at increasing\nreactor temperatures the rate of reaction passes through a maximum and then decreases. At the\nboiling point the reaction stops completely because the hydrogen cannot dissolve anymore into the\nboiling liquid (Westerterp and Crombeen (1983).\nA mathematical model of the slurry reactor has been developed which describes the hydrogen\nconversion, the rate of heat production (HPR), the total rate of heat withdrawal (HWR) and the\nliquid phase composition as a function of the reactor temperature.\nThe material balances of the reaction components are solved simultaneously using the kinetic model\nof the reaction system. Thus the overall conversion rate of hydrogen gas and the HPR can be\ndetermined. Using the lumped reaction model, for the hydrogenations of TNT we write:\nv,1\u00b0C\n$x,1 (CAO CA\nand CD\n=\nCAO\n(CA+\n+\nVR, 1 AB\n\u0441\u0432 \u00b0C)\n==\n*v,1\u00baB VR,1 (rAB rBC)\nVR,1 (BC r CD)\n=\n\u0441 catki H1 is the rate\nWhere A, B, C and D represent TNT, DNHAT, DANT and TAT respectively rij\nof the reaction of i into j. The material balance for the hydrogen is given by:\n*H\u2082out = \"H\u2082in *v,1 H2A0H2\n0 v,1 H\u2082 Ao represents the amount of moles of hydrogen required for complete conversion of the TNT\nin the feed to TAT, H\u2082 is the hydrogen conversion. * H2in is related to the reactant feed by\n*H\u2082in* *v,1\u00aa\u00b9H\u2082\u02daAo' in which a is defined as the amount of moles of H\u2082 fed to the reactor over the\namount required for complete conversion. Now we can write:\n*H\u2082out = *v,1\u00b0H2\u00baAo(a\n-\n5H2\n>\nThe HPR is calculated on the assumption that the heat liberated by the reaction is directly\nproportional to the rate of hydrogen conversion in the reactor:\nHPR =\nv,1\u00baAoH2 (HR)\nFor the HWR three contributions have been distinguished:\n1) the heat absorbed by the cold feed streams of the reactant solution, the catalyst slurry and\nthe hydrogen gas:\nv, feed (pCp) 1 (T-Tfeed)\n+\n*v, cat (pCp) 1 (T-Tcat)\n+\n\u0441\n,1\u00b0VH2Ao (MCp) H2 (T-TH2)\nHWR ef\n2) the heat absorbed by evaporation of the solvent. We assume that the gas and the liquid in the\nreactor are at equilibrium. The vapour stream that leaves the reactor has the same composition\nas the gas-phase in the reactor and consists of hydrogen, methanol and reaction water. It is\nrelated to H\u2082out by:\nHWR\n=\nev\nvap\n*H\u2082out \u0440 (pgxs+ Ph*w)\nvapW\nThe denominator of this equation represents the hydrogen pressure in the reactor.\n3) the heat absorbed by the condensate returned to the reactor. We assume complete condensation of\nthe vapour in the condensor, so we write:\nHWR\n=\ncond *H\u2082out\nPSXs (MCD)s+ PWW (MCD)W (T-T,\n\u0420 (poxs+ pwxw)\ncond)\nIn Fig. 4 the results of model calculations of the hydrogenation of TNT are presented for a\ntypical set of operating conditions. In Fig. 4a the HPR and HWR and in Fig. 4b the concentrations\n1,5\n1,0\nRelative\n1,0\nHPR and HWR\n|\n0,5\nHPR\n1\n0,8\nTNT\nDN-HAT\nDA-NT\nG/C,TNT\n0,6\n0,4\nTAT\n0,2\n0,0\n30\n0,0\n50\n70\nreactor temperature\n90\n110\n30\n50\n70\n90\n110\nreactor temperature\n\u0441\nFigure 4a:\nResults of calculations with the reactor model.\nDimensionless heat production rate (HPR) and heat\nwithdrawal rate (HWR) as a function of the reactor\ntemperature at constant residence time of the liquid.\nConstant catalyst food and gas feed temperature of 20\u00b0C.\nFeed temperature: a: 20\u00b0C, b: 30\u00b0C, c: 50\u00baC, &: 70\u00baC.\nFigure 4b:\nResults of calculations with the reactor model.\nDimensionless concentrations G /CTNT,o in the\nreactor product as a function of the reactor temperature\nat the same operating operating conditions as in\nfigure 4b case b. The dashed lines indicate the lower\nand upper stable operating point, respectivily.\nG12\nDevelopment of catalytic hydrogenation reactors\n2233\nof the components in the reactor product are plotted versus the reactor temperature at constant\nresidence time of the liquid in the reactor. The HPR and HWR have been normalized by dividing them\nby the maximum HPR at complete conversion of the TNT. The concentrations have been normalized by\ndivision by the feed concentration CAO. Because the calculations have been based on the lumped\nkinetic model of TNT the concentrations plotted in Fig. 4b do not represent the true product\ndistribution in the reactor but only give an impression of how the concentrations of the reaction\ncomponents change as a function of the reactor temperature.\nBecause the kinetic model does describe the hydrogen consumption rate very well the heat effects\ncalculated with the reactor model are realistic. Curves a, b, c and d in fig. 4a represent the HWR\nat four different feed temperatures. The reactor works at a stationary operating point when the\nstability criteria HWR-HPR and SHWR/8T > SHPR/ST are fullfilled. In Fig. 4a multiplicity occurs at\nthe feed temperatures of the a and b curves. In Fig. 4b the dashed lines indicate the lower and\nupper stable operating point of case b. The relative concentrations in the reactor outlet stream a\nthe lower stationary point are 0.42 TNT, 0.55 DNHAT and around 0.03 DANT and in the upper\nstationary point are 0.04 TNT, 0.10 DNHAT, 0.51 DANT and 0.35 TAT. Multiplicity can be avoided by\nusing a slightly higher feed then used for the calculations of case b. The most important\nparameters which control the behaviour of the reactor are the total pressure in the reactor, the\nresidence time of the liquid phase in the reactor and the temperature of the reactant feed.\nIn total of 22 hydrogenation of TNT experiments have been carried out in the mini-installation. By\nvarying the operation conditions a temperature range of around 35 to 100\u00b0C has been covered. In\nFig. 5 the experimental results are compared with the results of model calculations with respect\nto the hydrogen conversion and the temperature at the operating point. The agreement is\nsatisfactory, especially when realising that small errors in the calculation of the HPR or HWR may\nresult in a drastic change of the predicted steady state temperature.\n1.00\n0.80\n0.60\nI\n\"\nhydrogen\nconversion\n0.40\nI\nDIO\n0.20\nmodel calculations\nexperimental results\n0.00\n30\n50\n70\n90\nreactor temperature\nC\n110\nFigure 5: A comparison of experimentally determined hydrogen conversions\nand reactor temperatures and those predicted by our model.\n4.2 The three-phase_packed_bed_bubble_column_reactor\nA model comprising plug flow and axial dispersion of mass and heat is used to describe the\nbehaviour of the packed bed reactor. The derived model equations are based on the following\nassumptions: 1. the reactor is adiabatic and isobaric, 2. the gas phase is in equilibrium with the\nliquid phase when it enters the catalyst bed, 3. the gas and the liquid feed temperature are\nequal, 4. there are no mass transfer limitations, 5. the temperature in all three phases is\nuniform in a plane perpendicular to the direction of flow, 6. there are no concentration profiles\nin the radial direction, 7. the reactants are not volatile, 8. there is no heat conduction in the\naxial direction, 9. the gas phase is ideal, 10. the contribution of the reactants to the specific\nheat and the molar weight of the liquid phase can be neglected, 11. the available catalyst area is\nfully wetted and utilised, 12. a mass transfer rate equation is used to describe the evaporation\nof the solvent. The partial pressure of the solvent at the gas-liquid interphase is equal to the\nvapour pressure of the solvent.\nThe gas-liquid equilibrium at the entrance of the catalyst bed is achieved by assuming that the\ngas and the liquid feed flow through an adiabatic mixer in which this equilibrium is reached\nbefore entering the catalyst bed. Since some evaporation will occur in the mixer, the temperature\nat the entrance of the catalyst bed will be lower then the feed temperature.\nThe dimensionless equations based on the above assumptions are:\n1\n* for the reactants\n-\ndz\ndx.\n-\ndo\n+ 8\n* for hydrogen in the gas phase YH dz\n* for the solvent in the gas phase\n* for the liquid phase az +\n* enthalpy balance # d\u00b2\ni\nY\n\u13d3\u13be\nSt' P\nG\n+ Da\n= 0\nn R prod, i\na H\ndz\nDa n\n\u03b1\nRprod,\n0\nd Y\nsol\n+\nH 1\nsol\nSt P (\ndz\nsol)\n\u03a8\u0398\n0\n0\n\u0424\u043e\nof\n(S01 \u2717sol) \u00bd\n-\nY\n+40\n+\n+\nPen az\nCPGMGG)\n= \u043e\n* from the definition of mole fractions X sol\nThe boundary conditions for these equations are:\nde\n-Da\nn \u03a3(R, ATad,j)\nSt' PATev,sol\n=\n\u03a3 X j\nand\n\u0445\u0434\u043e\n1\nde\n* at Z = 0\n(CPGMG\u00aeGB\n+\n\u00b0L) (e\n- 0 +)\n+\nZ=O\nZ=O\nPen\ndz\n= \u00b0\n=\n1\n2234\nK. R. WESTERTERP et al.\n-pem\nX\nand\ni, Z-O-\n==== 1 -\nY sol\n$L\n=\n\u03a6\nL, Z=0\u00b0\nLO\nG\nG,Z=O\nd X.\n| z-o-\n-\nY sol\n=\nY\nsol,\nPem\nZ=O\u00af\n~1,Z=0+\nYH\nde\n* at Z=1\ndx i\n= 0\nG12\nThe conditions at Z=0\u00af follow from the balances around the mixer. The dimensionless groups and\nvariables are defined in the notation. The dimensionless temperature 0 is defined in accordance to\nthe definition used by Westerterp and Crombeen (1983).\nThe hydrogenation of TNT to TAT was chosen as a model reaction in our cocurrent upflow packed bed\nreactor. The kinetic rate equations presented in section 3 were used in the model calculations.\nThe catalyst used in the reactor is a shell catalyst of 0.08wt% Pd on 4.2*4.2 mm porous\ncilindrical pellets. It was assumed that the carrier does not influence the kinetic parameters. An\nallowance is made however, for a difference in the activity of the catalysts. The term Ccat is\nreplaced by n no Pp. in which pp is the packing density, a scaling factor and In the parameter\ndescribing the catalyst activity.\n\u05ea\nTwo important operating parameters are the hydrogen supply ratio a and the reactor pressure. The\nhydrogen supply ratio a is defined as the ratio of the number of moles of H\u2082 fed to the reactor to\nthe number of moles of reactant fed to the reactor. If n\u2081 is defined as the number of moles of the\nH2 needed for complete conversion of all reactant into the desired product i, we can define a\nsupply ratio a\u2081 as: di a/n. With this definition a value of a; 1 means that the exact amount\nof H2 is fed to the reactor necessary to convert all reactant into product i. For industrial\npurposes it can be advantageous to keep a\u2081, if i is the desired product, close to 1 to avoid\nhydrogen losses or high recycle ratios for unconverted hydrogen. On the other hand, if the value\nof\n=\nis chosen close to 1 the hydrogen partial pressure will drop to very low values because\npractically all hydrogen is consumed. Consequently the concentration of hydrogen at the catalyst\nsurface will decrease, resulting in lower reaction rates and the need for larger reactors.\nFurthermore, the value of a determines how much hydrogen is present and thus how much solvent\nevaporates. Therefore, a also has a strong influence on the amount of heat that is absorbed by\nevaporation and thus on the temperature level in the reactor.\nThe reactor pressure determines the boiling point of the solvent and thus the maximum temperature\nthat can be reached in the reactor.\nTypical results from a model run are given in Fig. 6a-d. Some important effects can be noted from\nthese figures:\nbecause more then one mole of solvent evaporates for each mole of H\u2082 consumed, the gas flow rate\nincreases.\n- the mole fraction of H2 drops to low values near the end of the catalyst bed due to the\nevaporation of the solvent.\nThe influence of CTAT and the reactor pressure on the fraction of the reaction heat that is\nabsorbed and removed by evaporation of the solvent is given in Fig. 7.\nA comparison between the temperature and conversion profiles measured during an experimental run\nand calculated with the model are given in Fig. 8 a and 8b. The parameters that were varied to\nobtain a best fit the model and the experiment are Pe\nFrom this figure we\nconclude that the between well describe the experimentally determined temperature and unversion\nprofiles. Since the kinetic data available at the present time do not allow calculation of the\ncomposition of the liquid phase, no conclusions can be drawn concerning the selectivity to the\nreaction intermediates.\n1.0\n\u2191\n0.8\n0.6\n5\n[-]\n0.4\n0.2\n3\nvapour\n2\nGa\n[-]\n1\nLiquid\n6c\n0.0\n0\n0.0\n0.2\n0.4\n0.6\n0.6\n1.0\n0.0\n0.2\ndimensionless length (-)\n0.4\ndimensionless length [-] -\n0.6\n0.8\n1.0\n110\n100\nT\n90\n[\u00b0C]\n80\n1.0\n0.8\n0.6\n\u0423\u043d\nob\n[-]\n0.4\n0.2\n6d\n0.0\n70\n0.0\n0.2\n0.4\n0.6\n0.8\n1.0\n0.0\n0.2\n0.4\n0.6\n0.8\n1.0\ndimensionless length [-]\ndimensionless length [-]\nFigure 6: Calculated profile for the packed bed reactor model.\nModel parameters: P10, Per-15, P-500 kPa, OTAT-2.4, food temperature = 50 \u00b0C, adiabatic temperature rise-170 \u00b0C, Da-5.\na. conversion, b. temperature, c. dimensionless flows, and d. mole fraction of H\u2082 in the gas phase\nG12\n[-]\n100\nP-0.2 MP\n90\nP-0.4 MPa.\n80\n|\n70\nP=0.6 MPa\n60\n\u00b0\n1\n2\n3\n$\n\u03b1 [-]\nDevelopment of catalytic hydrogenation reactors\n110\nCalculated\n\u25a0 Experimental\n100+\nT\n(\u2103)\n90\n1.0\n0.8\n0.6\nH}\n0.4\n2235\n0.2\nCalculated\nExperimental\nBO\n0.0\n0.0\n0.2\n0.4\n0.6\n0.8\n1.0\n0.0\n0.2\n0.4\n0.6\n0.8\n1.0\nZ [-]\nZ (-)\nFigure 7: Percentage of the reaction heat (at 100% conversion)\nthat is used for evaporation of the solvent. Model parameters are:\nP-10, Po-1.5, feed temperature - 55 \u00b0C and Du - 3.5.\nFigure 8: Comparison between experimental data and calculated profiles. Model parameters are:\n-2\n111\n\u25a0 5.0, Pe -0.9 and St' - 760.\nh\nTNT,0\nDa = 6.95, \u03b1 TAT = 2.4, y = 5.38, n = 0.365, Pe\nExperimental data: \u2013 1.06\u00b710 \u00b4 m/hr, \u2666 = 1.08 Nm/hr, T-50 \u00b0C, x, -0.00873\nadiabatic temperature rise - 170\u00b0C, P - 490 kPa and reactor length - 1 m.\na. temperature profile, b. conversion profile.\n5.\nDiscussion and conclusions\nIn a bird's eye view we presented some results of our program for the development for catalytic\nhydrogenation reactors for multi-product plants. We focussed on the kinetics of our model\nreactions, the modelling of our reactors and on the experimental verification of our models.\nWe showed both theoretically and experimentally - that for every exothermic reactions heat\ncan easily be achieved by evaporation of a solvent. It is now possible to install the\ncooling area outside of the reactor. Temperature control can be easily achieved by adjusting the\nreactor pressure.\nremoval\nA disadvantage of the evaporation of the solvent is the subsequent decrease of the partial\npressure of hydrogen. This may result in a decrease of the rate of reaction. The decrease in the\npartial pressure of hydrogen can be compensated for by increasing the reactor pressure, by using a\nlarger excess of hydrogen or by employing larger interfacial areas.\nThe interdependence of all the operating variables of the reactors is so complex, that a catalytic\nhydrogenation reactor in a multi-product plant can not be run without the use of a mathematical\nmodel that describes its behaviour. The good agreement observed between our models and the\nexperimental data gives us the confidence that our models will be adequate for this purpose.\nThe most difficult aspect is the acquisition of sufficiently reliable kinetic data. For fine\nchemicals the tremendous effort required to obtain these data with present methods is prohibitive.\nNew methods have to be developed to this end. This will be a subject of further study, especially\nfor the hydrogenation of TNT.\nAcknowledgement The investigations were partly supported by the Dutch Technology Foundation STW\nand partly by Andeno BV, Venlo, the Netherlands.\nReferences\n1. Acres, G.J.K. and Cooper, B.J., 1972, J. Appl. Chem. Biotechnol., 22, 769-785.\nBird, A.J. and Thompson, D.T., 1980, Academic Press, New York.\n2.\n3.\nBurge, H.D., Collins, D.J. and Burtron, H.D., 1980, Ind. Engng. Chem.\n389-391.\nProd. Res. Dev., 19,\n4.\n5.\n6.\n7.\n8.\n9.\nCollins, D.J., Smith, A.D. and Burtron, B.H., 1982, Ind. Engng. Chem. Prod. Res. Dev., 21\n279-281.\nHaber, F., 1898, Z. Elektrochem., 22, 506.\nKan, Kin-Mun and P.F. Greenfield, 1983, A. I.Ch.E.J. 29, 123-132.\nKlaus, R.A. and Rippin, D.W.T., 1979, Comp. & Chem. Engng., 3, 105, see also Klaus, R.A.,\n1981, Thesis E.T.H. Z\u00fcrich.\nMichelsen M.L. and K. \u00d8stergaard, 1970, Chem. Eng. Sci. 25, 583-592.\nPawlowski, J. and Kricsfalussy, Z., 1981, Chem.-Ing.-Techn., 53, 652-654.\n10. Yao, H.C. and Emmet, P.H., 1962, J. Am. Chem. Soc., 84, 1086-1091.\n11. Y\u00fccelen, F., 1984, Thesis E.T.H. Z\u00fcrich.\n12. Westerterp K.R. and Crombeen P.R.J.J., 1983, Chem. Eng. Sci., 1331-1340.\n13. Westerterp K.R., van Swaaij W.P.M. and Beenackers A.A.C.M., 1984, John Wiley & Sons,\nChichester, UK.\nNotation\n2A-4NT\n2A-4HAT\n2N-4HAT\n4A-2NT\nag\n\u0412\u043e\nC\nC\ncat\n\u0441\u0440\n2-Amino-4-Nitrotoluene\n2-Amino-4-Hydroxylaminotoluene\n2-Nitro-4-Hydroxylaminotoluene\n4-Amino-2-Nitrotoluene\nspecific gas/liquid interfacial area\nBodenstein number\nconcentration, [mol m\u00b3]\nconcentration of catalyst, [kScatm\u00af\u00b3]\nspecific heat\ndimensionless specific heat of the gas\naxial dispersion coefficient\n[1/m]\n[kJ/kg K]\nphase defined as Eypi/p\n[m/s]\n2236\nK. R. WESTERTERP et al.\nG12\nDa\nDANT\nDNHAT\nDNT\nDamk\u00f6hler number defined as no Pb k\u2081,S L/OLO XA,O\ndiamino nitro tolue DAT\ndinitro hydroxylamine toluene\n2,4-dinitrotoluene\n2,4-diaminotoluene\nHPR\nHeat Production Rate, W\nHWR\nHeat Withdrawal Rate, [w]\nk\nk t\nK\nL\nM\nMG\nni\nP\nP1\nPR\nP\nmass transfer coefficient, [ms]\nreaction rate constant, [mole s 1kg cat \u00af1]\nadsorption constant, (m\u00b3 mol]\nreactor length\nmolar mass\n[m]\n[kg/kmol]\ndimensionless molar mass of the gas phase defined as Ey, M./MH2.\nnumber of moles of hydrogen needed for complete conversion of 1 mole of reactant to\ncomponent i\ndimensionless pressure defined as P/Pg\nvapour pressure of pure species i,\ntotal pressure\nStandard Pressure\n[Pa]\n[N/m\u00b2]\nPeclet number for heat dispersion\n\u0398\nr\n1\nR.\ndimensionless reaction rate: ri/pbk 1, s.\nprod,i\nPeclet number for mass dispersion\nrelative adsorption strength,\nrate of conversion, [mole s\u00a8\u00b9kg,\ndimensionless production rate of component i\nmodified Stanton number (also called NTU)\ngat 1]\nT\nT\u00c5T\nTNT\nVr,1\ntemperature\n[K]\nstandard temperature\n2,4,6-triaminotoluene\n2,4,6-trinitrotoluene\nliquid volume in the reactor, [m\u00b3]\nliquid mole fraction of species i\ndimensionless mole fraction of component i: \u00d7\nmole fraction of component i in the gas phase\ndimensionless axial coordinate = h/L\nA,O\n\u03b1 i\nB\n\u03a5\n\u0394\u0397\nev\nAHri\n\u0394\u0397\nAT.\nAT\n\u20ac G\nvap\nad,i\nev,i\n5H2\"\n5\nhydrogen supply ratio: ratio of the number of moles of hydrogen fed to the reactor per unit\nof time to the number of moles of TNT fed to the reactor per unit of time\nhydrogen supply ratio based on component 1\nratio of heat capacities of the gas and the liquid phase in the feed\nmolar ratio of methanol to hydrogen in the feed\nenthalpy of evaporation\nreaction enthalpy\nheat of evaporation, (kJ mole\u00af1]\n[kJ/mole]\n[kJ/mole of component i converted]\ndimensioness adiabatic temperature rise for reaction i\ndimensionless heat effect of evaporation for component i\nhold up of the gas phase\n[m\u00b3gas/m\u00b3reactor]\nhydrogen conversion, defined as the amount of hydrogen consumed by the reaction over the\namount required for complete conversion\ncatalyst activity\nscaling factor from the Pd on Carbon slurry catalyst to the Pd on Alumina catalyst used in\nthe packed bed reactor: 1.28 * 10\u00b0\nrelative surface concentration\neffective axial thermal conductivity\nstoichiometric coefficient of hydrogen of the overall reaction\nn\n\u041f\u043e\n\u04e8\n0\ndimensionless temperature: -(E\u2081/RT_) (1\nT/T\n[W/m\u00b2K]\n\u03bbeff\n[kg/m\u00b3]\n\u03a1\ndensity\n[kgcat/m\u00b3reactor]\nPb\nspecific catalyst mass\n\u03a6\nfeed rate of hydrogen g\u00e5s, (mole s\u00af1]\nmolar flow rate per unit of cross sectional\ndimensionless flow: $100\narea\nof the reactor [kmol/m\u00b2s]\n8\nH\u2082in\n*H\u2082out\nflow rate of the hydrogen off-gas, [mole s\u00af\u00b9]\ntotal liquid feed rate to the reactor, [m\u00b3s-1]\nsub and super-scripts\nax\nin axial direction\nsol\nb\nrefering to catalyst bed\nG\nsolvent\ngas\ncat\ncatalyst, catalyst feed stream\nH\ncf\ncold feed streams\nL\nhydrogen\nliquid\ncond\ncondensate returned\n\u043e\nev\nevaporation of the solvent\nS\nfeed\nreagent feed stream\nS\nfeed conditions\nrefering to standard conditions\nsolvent, methanol\n1\nliquid\nW\nreaction water\nO\ninitial\n"}, "expected_output": {"claims": [{"unit": "-", "value": 2.4, "evidence": ["Two important operating parameters are the hydrogen supply ratio a and the reactor pressure. The hydrogen supply ratio a is defined as the ratio of the number of moles of H\u2082 fed to the reactor to the number of moles of reactant fed to the reactor.", "For industrial purposes it can be advantageous to keep a\u2081, if i is the desired product, close to 1 to avoid hydrogen losses or high recycle ratios for unconverted hydrogen.", "Da = 6.95, \u03b1 TAT = 2.4, y = 5.38, n = 0.365, Pe", "\u03b1 [-]", "Two important operating parameters are the hydrogen supply ratio a and the reactor pressure. The hydrogen supply ratio a is defined as the ratio of the number of moles of H\u2082 fed to the reactor to the number of moles of reactant fed to the reactor. If n\u2081 is defined as the number of moles of the H2 needed for complete conversion of all reactant into the desired product i, we can define a supply ratio a\u2081 as: di a/n. With this definition a value of a; 1 means that the exact amount of H2 is fed to the reactor necessary to convert all reactant into product i. For industrial purposes it can be advantageous to keep a\u2081, if i is the desired product, close to 1 to avoid hydrogen losses or high recycle ratios for unconverted hydrogen.", "Model parameters: P10, Per-15, P-500 kPa, OTAT-2.4, food temperature = 50 \u00b0C, adiabatic temperature rise-170 \u00b0C, Da-5.", "hydrogen supply ratio: ratio of the number of moles of hydrogen fed to the reactor per unit of time to the number of moles of TNT fed to the reactor per unit of time"]}]}, "metadata": {"product_category": "Chemical products", "request_id": "req_21fcbf13dee52ada"}}