2 marzo 2009 << la metallurgia italiana

Transcript

Memorie >>
Selezione materiali
CRITERI DI SCELTA DEL MATERIALE
PER L’ALLEGGERIMENTO DI VETTURE
SPORTIVE AD ALTE PRESTAZIONI
P. Veronesi, A. Pivetti, A. Baldini, M. Loiacono, G. Poli
Per le vetture di lusso ad alte prestazioni, il primo elemento di competitività è dato dalle prestazioni dinamiche
della vettura, ragion per cui il fattore peso sta assumendo nel tempo una rilevanza crescente. L’introduzione di normative sempre più severe in ambito di emissioni, strettamente correlate al consumo della
vettura ha inoltre indotto i progettisti a spostare la propria attenzione non più sulla potenza pura ma sul
rapporto potenza/peso. In questo ambito, è stato studiato l’alleggerimento di una vettura sportiva ad
alte prestazioni, in particolare di una Lamborghini Murciélago, andando a proporre nuovi materiali per
la realizzazione di particolari del gruppo sospensivo. Ottimizzando la scelta del materiale, è possibile ridurre
il peso, rispetto alla soluzione attuale, del 30-35% relativamente alla molla sospensione anteriore, del 50-70%
relativamente alla barra antirollio posteriore, nonché alleggerire il braccio anteriore inferiore dal 3 al
30%, dipendentemente dallo stato tensionale del componente, impiegando opportunamente acciai basso legati,
leghe di alluminio e leghe di titanio.
PAROLE CHIAVE: acciaio, alluminio e leghe, mat. compositi, titanio e leghe, selezione materiali
INTRODUZIONE
Negli ultimi 20 anni la progettazione e realizzazione di autoveicoli ha subito notevoli cambiamenti per poter venire incontro a nuove e differenti esigenze da parte degli utenti.
Innanzitutto sono cambiati gli standard per quanto riguarda
confort e abitabilità: lo spazio a disposizione dei passeggeri
è aumentato ad ogni nuova generazione di veicoli e con esso le
dimensioni delle vetture.
Parallelamente a questo, la richiesta di vetture sempre più sicure per gli occupanti ma anche per i pedoni ha reso necessario
un aumento delle dimensioni medie delle vetture e di conseguenza un aumento di peso delle stesse [1].
Negli ultimi anni ci si è però resi conto che questo trend
non poteva proseguire poiché il miglioramento tecnologico
di propulsori e combustibili, per quanto notevole, non sarebbe
stato in grado di compensare l’aumento di peso delle vetture,
soprattutto considerando le richieste sempre maggiori del mercato in termini di prestazioni dinamiche (migliore acceleraP. Veronesi, G. Poli
Dipartimento di Ingegneria dei Materiali e dell’Ambiente,
Via Vignolese 905, 4110 Modena - Italy
A. Baldini, M. Loiacono
Dipartimento di Ingegneria Meccanica,
Via Vignolese 905, 4110 Modena - Italy
A. Pivetti
Lamborghini, Via Modena 12, S. Agata Bolognese - Bologna - Italy
zione, ripresa, maneggevolezza, minori consumi/emissioni). Il
peso di una vettura infatti può influenzare numerosi parametri
della progettazione oltre che le prestazioni del veicolo stesso:
- veicoli più pesanti richiedono maggiori potenze per ottenere prestazioni analoghe a mezzi più leggeri; questo
comporta solitamente un maggiore consumo di carburante
e quindi maggiori emissioni di CO2.
- Un peso maggiore significa anche maggiori inerzie e quindi
minor prontezza di risposta ai comandi e minor piacere di guida.
- In caso di incidente una maggiore massa implica una maggiore energia cinetica da dissipare e quindi richiede delle prestazioni di resistenza strutturale maggiori da parte del veicolo
In questo ambito è opportuno inoltre ricordare i nuovi limiti
di emissioni di CO2 (130 g/km) imposti dalla Comunità Europea che entreranno in vigore nel 2012: per riuscire a rientrare
in tali limiti sarà fondamentale lavorare su una riduzione dei
pesi delle vetture, perché la sola efficienza dei motori non sarà
assolutamente sufficiente. I problemi maggiori li incontreranno
sicuramente le vetture di dimensione medio-grande, nonché
quelle ad alte prestazioni, che sono l’oggetto di studio del
presente lavoro.
Nel settore delle vetture di lusso ad alte prestazioni il primo
elemento di competitività è dato dalle prestazioni dinamiche
della vettura; in questo settore, il fattore peso sta assumendo nel
tempo una rilevanza sempre maggiore. L’introduzione di normative sempre più severe in ambito di emissioni, strettamente
correlate al consumo della vettura ha infatti indotto i progettisti
a spostare la propria attenzione non più sulla potenza pura
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s
Fig. 1
Gruppo sospensivo: braccio anteriore inferiore.
suspension system: front lower arm.
ma soprattutto sul rapporto potenza/peso. Ciò può essere
conseguito cercando di ridurre il peso della vettura il più possibile. L’alleggerimento della vettura non può prescindere dalla
scelta dei materiali e dei processi, che però deve essere
condotta in maniera oculata e su basi razionali. Troppo
spesso, infatti, nella realizzazione di un nuovo componente
ci si basa solo sulle conoscenze pregresse: ciò significa che si
tende spesso a utilizzare quei materiali che per esperienza ed
uso comune sono stati utilizzati in passato. Questo approccio,
però, limita notevolmente l’inventiva e non porta quasi mai a
scelte e soluzioni innovative.
Nel presente lavoro si è scelto di adottare una procedura
basata su un processo di scelta il più possibile oggettivo e
di ampio respiro [2] costituito da quattro fasi consecutive
di traduzione dei requisiti di progetto, screening, classificazione e ricerca informazioni di supporto sui materiali
candidati per l’applicazione di interesse. Verranno presentate tre applicazioni del processo di selezione dei materiali
a componenti differenti dell’autovettura, precisamente una
Lamborghini Murciélago, focalizzando l’attenzione su parti
del gruppo sospensivo, quali braccio anteriore inferiore, molla anteriore e barra antirollio posteriore.
ANALISI ED ALLEGGERIMENTO BRACCIO ANTERIORE
INFERIORE
Il componente, rappresentato in Fig. 1, è fissato da un lato al
telaio mediante due boccole e dall’altro al portamozzo. Circa
a metà lunghezza del braccio è presente una vite prigioniera
sulla quale viene fissata la barra antirollio.
Considerando le diverse sollecitazioni a cui il braccio è sottoposto in seguito agli scuotimenti del portamozzo e alle forze
laterali in curva, si può in prima approssimazione considerare il componente come un tirante o puntone, ovvero soggetto unicamente a sforzo normale. In realtà la presenza dell’attacco alla barra antirollio e le forze longitudinali agenti sul
veicolo generano un momento flettente sul braccio, che quindi
può comportarsi in certe situazioni anche come una trave.
Attualmente, il componente è realizzato impiegando due materiali: le boccole di fissaggio al telaio ed al portamozzo sono in
acciaio al carbonio, con allungamento a rottura del 22%; i tubi
trafilati e saldati sono realizzati sempre con un acciaio al
carbonio, ma avente caratteristiche meccaniche leggermente
inferiori
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Traduzione dei requisiti di progetto
Per il componente in esame, il vincolo tecnico più stringente è
legato alla sua rigidezza in esercizio.
Il braccio, infatti, deve svolgere una funzione di guida per il portamozzo, influenzando quindi sia il comportamento cinematico
che dinamico della sospensione: è quindi importante che la
rigidezza del componente sia elevata, per limitare al massimo gli
scostamenti dal comportamento ideale della sospensione. Deformazioni eccessive potrebbero incidere in maniera negativa
sul comportamento dinamico della vettura. Dal punto di vista
della resistenza meccanica, ovviamente il braccio deve poter reggere tutti i carichi a cui è sottoposto, senza subire deformazioni
plastiche e riducendo al minimo masse e ingombri. Dato il tipo
di impiego, il materiale utilizzato per il componente dovrà anche
essere facilmente saldabile e resistente a corrosione.
Riassumendo, in accordo con il metodo esposto in [2], la traduzione dei requisiti di progetto è la seguente:
- Funzione: tirante/puntone con funzione di guida per il mozzo
ruota
- Vincoli: rigidezza elevata, allungamento a rottura >15%,
sufficiente resistenza strutturale, buona resistenza a fatica e a
corrosione
- Obiettivo: minimizzare la massa del componente
- Variabile libera: materiale da utilizzare, forma geometrica delle
sezioni
Screening e classificazione: calcolo indici di prestazione e
individuazione materiali alternativi
Considerando lo stato tensionale del componente, due indici di prestazione del materiale da utilizzare per valutare le
prestazioni del componente, e da massimizzare, sono quelli
relativi a rigidezza e resistenza per un tirante:
(1)
(2)
dove E= modulo di Young, ρ= densità, σf= resistenza del materiale, nel senso di un suo collasso strutturale, che nel caso di
metalli si fa corrispondere alla resistenza allo snervamento.
Qualora il componente si trovi a lavorare a flessione, sono
da considerare altri due indici del materiale relativi a rigidezza
e resistenza riferite ad una trave, ovvero:
(3)
(4)
Lo screening dei materiali utilizzabili per l’applicazione può
essere condotto per via grafica, utilizzando appositi software,
come il CES-Cambridge Engineering Selector (Grantadesign). Su
due assi ortogonali vengono quindi rappresentati gli indici di
prestazione dei materiali candidati, eventualmente rapportati
agli indici calcolati per il materiale in uso. In quest’ultimo
caso, il materiale attualmente in uso si troverà alle coordinate
(1,1). Per convenzione, si preferisce condurre la scelta minimizzando gli indici di prestazione, per cui sarà sufficiente considerare l’inverso degli indici espressi nelle equazioni 1-4.
In Fig. 2 sono rappresentati i grafici di selezione relativi al componente in oggetto, considerato come tirante rigido di massa minima (a) o come trave rigida di massa minima (b), rapportato alla
soluzione attualmente in uso. Data l’inversione degli indici,
i materiali con prestazioni migliori della soluzione attuale giaceranno nel terzo quadrante.
Tutti i materiali posti nel terzo quadrante rappresentano soluzio-
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a
s
Fig. 3
b
Costo per unità di massa dei materiali – database CES-Edupack 2005.
Suspension system: front lower arm.
diagrammi di scelta per tirante rigido di massa
minima e per trave rigida di massa minima, con indici
dei materiali rapportati al materiale attualmente in uso
per il componente.
Selection diagrams for a stiff tie having minimum mass
and for stiff beam having minimum mass; materials
indexes are normalized to the properties of the material
currently used for the component.
Materiale
AISI 94B30 (G94301), temprato e rinvenuto a 205°C
AISI 4042 (DIN 42MnMo7), temprato e rinv. a 205°C
AISI 4340 (UNI 40NiCrMo7), normalizzato
Al-Lega da getto S520.0
Al-2026 T3
Al-2024 T3
Al-6082 T4 (anche forgiato)
Al-7020 T5
Al-7033 T6
Informazioni di supporto e scelta del materiale
Le informazioni ricavate dagli indici di prestazione hanno
M1/M10
1.02
1.00
1.10
1.06
1.00
0.99
0.99
0.96
0.96
M2/M20
5.50
3.59
2.60
1.92
2.77
2.50
1.53
2.15
3.07
M3/M30
1.00
1.00
1.00
1.79
1.68
1.67
1.69
1.64
1.67
M4/M40
2.3
2.3
1.9
2.2
2.8
2.6
1.9
2.3
3.0
s
s
Fig. 2
ni alternative in grado di conferire migliori prestazioni al componente, relativamente all’obiettivo individuato e ai vincoli agenti.
Un fattore inizialmente non considerato, ma che può giocare un
ruolo determinante nella scelta del materiale, è il costo. In Fig. 3
è rappresentato il diagramma relativo al costo per unità di massa
di differenti tipologie di materiali, con indicate le barre relative
ad acciai, leghe di alluminio e titanio.
Dall’analisi dei grafici di Fig. 2 e 3, si può notare come dal punto
di vista della rigidezza specifica (M1), si ha una famiglia di materiali in grado di avere prestazioni superiori all’acciaio: le leghe di
titanio; tuttavia i miglioramenti sono possibili ma in misura
ridotta (pochi punti percentuali) e a costi nettamente maggiori. Non sembra quindi opportuno proseguire verso una
scelta del genere, considerando i limitati vantaggi in rapporti
ai costi. Dal punto di vista della resistenza specifica (M2),
invece, si possono ottenere miglioramenti considerevoli.
Relativamente alle prestazioni di rigidezza flessionale, i grafici
mostrano che vi sono soluzioni in grado di fornire buoni miglioramenti: in particolare le leghe di alluminio sembrano
offrire le migliori soluzioni.
La rosa di candidati per l’applicazione, si restringe pertanto ai
materiali riportati in Tab. 1, con i relativi indici di prestazione
riferiti al materiale attualmente in uso (apice “0”).
Tab. 1
indici
di prestazione
dei materiali
candidati.
Materials
indexes for
candidate
materials.
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permesso di individuare una rosa di candidati, tra cui scegliere il
materiale migliore. Si ricorda che le informazioni presenti nei database sono comunque generiche, a bassa precisione, soprattutto
relativamente a parametri resistenziali quali resistenza a corrosione e fatica.
Per quanto riguarda gli acciai, le proprietà di resistenza a corrosione e a fatica sono ben note e più che sufficienti, considerando
anche quanto già noto per applicazioni precedenti. Analogamente,
le leghe di alluminio, eventualmente anodizzate, offrono buone
proprietà di resistenza nei tipici ambienti stradali. Per quanto
riguarda la resistenza a fatica le leghe di alluminio individuate presentano buone prestazioni, soprattutto le serie 2000 e 7000,
sebbene sia noto che il diagramma di Wohler per tali leghe non
raggiunga mai condizioni di stazionarietà al crescere del numero di
cicli di fatica, al contrario di quanto avviene per le leghe ferrose.
Pertanto, le possibili soluzioni ottimali sono fondamentalmente
due a seconda dei vincoli prestazionali/di progetto:
- se la sollecitazione principale agente sul componente è quella
a trazione, mentre quella a flessione è trascurabile, allora la
soluzione migliore è ancora data dagli acciai ad alta resistenza
(acciai basso legati o acciai a basso contenuto di carbonio);
- se invece la sollecitazione principale è quella di flessione, le leghe di alluminio rappresentano sicuramente la scelta
migliore; in particolare se il vincolo progettuale più stringente è
quello di rigidezza le leghe della serie 6000 rappresentano la scelta
più opportuna; se è la resistenza a limitare le prestazioni allora la
scelta migliore è data dalle leghe delle serie 2000 e 7000. Come già
sottolineato è opportuno verificare sempre le prestazioni di
resistenza a corrosione e a fatica della lega selezionata.
ANALISI ED ALLEGGERIMENTO MOLLA SOSPENSIONE
ANTERIORE
La molla cilindrica della sospensione anteriore è un componente
sottoposto ad elevate sollecitazioni statiche ma soprattutto dinamiche (ovvero a fatica) ed ha come scopo principale quello di
immagazzinare e restituire la maggior parte possibile dell’energia causata dagli scuotimenti della sospensione. In fase di progettazione del cinematismo della sospensione e di conseguenza del comportamento dinamico del veicolo, uno dei parametri
di progetto fondamentali da stabilire è dato dalla rigidezza della
molla. Parallelamente, un altro aspetto da tenere in considerazione
riguarda il calcolo e la verifica degli ingombri massimi.
Analogamente al procedimento seguito per il caso precedente, la
traduzione dei requisiti di progetto porta alle seguenti condizioni:
- Funzione: immagazzinare e restituire energia con le minori dissipazioni possibili
- Vincoli: resistenza ai carichi in esercizio; resistenza a fatica; resistenza alla corrosione; rigidezza prestabilita; coefficiente di perdita inferiore a 0.01; ingombro compatibile con spazi disponibili
- Obiettivo: minimizzare la massa del componente
- Variabile libera: materiale da utilizzare
Screening e classificazione: calcolo indici di prestazione e
individuazione materiali alternativi
L’indice di prestazione per un materiale per molle efficienti e leggere è noto da letteratura [3] ed è
pari a :
(5)
Risulta utile esprimere l’indice dell’equazione (5) in forma logaritmica, ovvero:
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s
Fig. 4
Diagramma di scelta per molla efficiente e leggera.
Material selection chart for efficient and light spring.
(6)
che rende possibile visualizzare i materiali candidati su un grafico
di scelta con assi la resistenza specifica e il modulo elastico specifico.
Il materiale attualmente utilizzato è un acciaio per molle ad alta
resistenza basso legato, contenente Silicio e Cromo; il materiale
rispetta le specifiche della normativa tedesca DIN 17223/2. Il
trattamento termico di tempra e rinvenimento è prescritto su questo tipo di acciai per aumentarne le proprietà di resistenza a rottura. Il trattamento superficiale di pallinatura viene invece impiegato
per aumentare le proprietà di resistenza a fatica del componente.
Il relativo diagramma di scelta è riportato in Fig. 4, insieme alla
retta di selezione, che altro non che la rappresentazione dell’equazione (6). Spostando l’intercetta all’origine verso il basso corrisponde ad aumentare l’indice M1; pertanto, traslando la retta si
selezioneranno materiali via via più prestazionali.
La rosa decisamente ampia di materiali apparentemente utilizzabili non tiene conto dell’ingombro: materiali con basso E, come i
materiali polimerici, infatti, richiedono che si debba realizzare
la molla con sezioni di filo molto grandi, impedendo quindi
al componente di essere installato nell’ammortizzatore esistente.
Per quanto riguarda i materiali compositi con rinforzo a fibra lunga, è evidente la difficoltà realizzativa di un componente come
la molla cilindrica. Essi sono tuttavia ottimi candidati per la
realizzazione di elementi elastici efficienti e leggeri, sebbene con
altre geometrie [4].
Considerando anche il fattore costo (omesso nel presente lavoro),
l’elenco di materiali candidati si restringe a:
- Leghe di titanio: Ti4.5Fe6.8Mo1.5Al (TIMETAL® LCB);
Ti3Al8V6Cr4Zr4Mo (“Ti 38644”); Ti15V3 Cr3Sn3Al; Ti6Al2Sn2Zr2Mo; Ti6Al6V2Sn (6-6-2); Ti6Al2Sn2Zr2Mo; Ti6Al4V
- Acciai: AISI 5160 (DIN 65MnCr4), temprato e rinvenuto; AISI 9255
(UNI 55Si8), temprato e rinvenuto
Informazioni di supporto e dimensionamento del
componente
Individuati i migliori materiali per ogni famiglia è possibile effettuare un calcolo di primo dimensionamento e di conseguenza
una stima della possibile riduzione di peso. La molla in esame lavora unicamente in compressione: è sottoposta ad una forza statica
costante legata al peso della vettura; oltre a tale forza vi sono
altre sollecitazioni dinamiche legate allo schiacciamento della
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molla in seguito a variazioni della quota z del contatto ruotasuolo. In ogni caso, per qualsiasi scuotimento della sospensione
la molla rimane sempre sollecitata a compressione. Il filo della molla è avvolto a spirale attorno ad un cilindro fittizio di diametro costante, denominato diametro di avvolgimento. Come già
accennato, il componente deve rispettare determinati ingombri
per consentirne il montaggio: in particolare un parametro che non
può essere modificato è il diametro interno della molla. Tale quota
è infatti limitata dalla necessità di montare all’interno della molla
l’ammortizzatore. Un’altra specifica tecnica che non deve essere
modificata è la rigidezza; tale parametro infatti influenza il comportamento dinamico della vettura e viene quindi stabilito in fase
progettuale per ottenere le prestazioni dinamiche volute. Anche
questo parametro verrà quindi mantenuto costante. Per quanto
riguarda la lunghezza della molla, si cercherà di mantenerla il più
possibile simile a quella di origine, in modo da ridurre al minimo eventuali modifiche su componenti necessari al fissaggio
della molla, come per esempio i piattelli; ovviamente esistono
limitazioni legate all’ingombro longitudinale ma una variazione
di qualche millimetro può essere concessa. Va osservato che si
tratta di una molla con passo e diametro medio di avvolgimento costanti; la sua caratteristica forza-spostamento sarà quindi
lineare. Il dimensionamento ha portato ai valori riportati in
Tab. 2, espressi in termini di variazione percentuale per ragioni
di riservatezza.
ANALISI ED ALLEGGERIMENTO BARRA ANTIROLLIO
La barra antirollio svolge un compito di ausilio al sistema sospensivo solamente in situazioni in cui si ha trasferimento di carico.
Gli estremi della barra sono fissati ai braccio superiore del
sistema sospensivo: in tal modo quando si verifica un trasferimento di carico tra le due ruote dell’assale, gli scuotimenti verticali di
verso opposto tra le due ruote generano un momento torcente sulla
barra, che quindi viene sollecitata a torsione. Il componente è mostrato in Fig. 5.
Un parametro fondamentale che sintetizza il comportamento
della barra stessa è la sua rigidezza torsionale: più la barra è
rigida maggiore sarà la resistenza che essa opporrà e di conseguenza minori saranno gli scuotimenti del sistema sospensivo
dell’assale.
Nel dimensionare il componente quindi l’aspetto principale da
s
Fig. 5
Barra antirollio.
Anti-roll bar.
tenere i considerazione sarà la rigidezza torsionale che si vorrà
ottenere; come si vedrà in seguito tale proprietà dipende sia da
quote dimensionali (diametro della barra, lunghezza del braccio di
torsione) che dalle proprietà del materiale.
Oltre alle funzioni esposte al paragrafo precedente, un parametro importante per la selezione del materiale è la massima
temperatura di esercizio: la barra infatti è situata in prossimità del
motore e dell’impianto di scarico della vettura, e si troverà quindi
a lavorare ad una temperatura leggermente superiore a quella ambiente. Come limite inferiore è stata imposta una temperatura di
lavoro di 70° C.
Un altro aspetto da tenere in considerazione è il comportamento
del materiale in caso di frattura: la caratteristica desiderata, ovviamente, è che in caso di cedimenti il materiale non ceda di schianto
ma resista il più possibile alla propagazione della cricca. Il parametro solitamente utilizzato per definire il comportamento di un
materiale in caso di cedimento è la tenacità a frattura. La
traduzione dei requisiti di progetto è la seguente:
- Funzione: barra di torsione
- Vincoli: rigidezza torsionale, resistenza a rottura, resistenza a fatica e ad agenti atmosferici, vincoli geometrici di ingombro; temperatura di impiego superiore a 70°C; tenacità a frattura superiore a
15 MPa m1/2.
- Obiettivo: ottenere rigidezza di progetto con minimo peso
- Variabile libera: materiale da utilizzare, dimensione della sezione
resistente
Screening e classificazione: calcolo indici di prestazione e individuazione materiali alternativi
L’indice di prestazione per la rigidezza torsionale e per la resistenza torsionale sono i seguenti [5]:
(7)
(8)
s
Tab. 2
Dimensionamento della molla in lega di titanio e
relativa riduzione di peso.
Dimensioning of titanium alloy spring and consequent
weight reduction.
Il relativo diagramma di scelta, nella zona di interesse, è riportato
in Fig. 6.
In base alle considerazioni effettuate in precedenza relativamente all’utilizzo di componenti in materiale composito, la scelta più
ragionevole sembra quella delle leghe di alluminio da trattamento
termico e/o deformazione plastica. In base all’analisi del grafico le
leghe migliori sono quelle delle serie 5000 e 2000. In particolare
le leghe migliori sono quelle della serie 5000, che presentano
caratteristiche di resistenza sufficienti e una densità leggermente
inferiore rispetto alle altre serie, massimizzando così l’indice di
prestazione. Una scelta opportuna potrebbero essere le leghe 5052
e 5086. Tuttavia, semplificando la geometria del componente,
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Considerando il rapporto prestazioni/costo, l’utilizzo della lega
di alluminio rimane sicuramente la scelta migliore: la riduzione di
peso ottenibile è molto vicina a quella della fibra di carbonio, ma a
costi decisamente inferiori.
CONCLUSIONI
Per effettuare lo studio di alleggerimento è stato adottato un metodo basato su criteri oggettivi [2].
Grazie a questo metodo, implementato tramite software, è stato
possibile compiere un’analisi dettagliata delle possibili soluzioni
alternative in termini di scelta dei materiali. La bontà e l’affidabilità del metodo è stata verificata anche tramite un dimensionamento dei componenti, che ha mostrato come una scelta oculata
del materiale possa portare a significative riduzioni di peso.
Riassumendo i risultati ottenuti per i due componenti dimensionati sono i seguenti:
s
Fig. 6
????????.
???????.
anche una barra di torsione in materiale composito come CFRP
a matrice epossidica.
Informazioni di supporto e dimensionamento del
componente
Sulla base dei materiali individuati, il dimensionamento del
componente porta alle variazioni esposte in Tab. 3.
Diametro esterno
Spessore
Rigidezza
Carico statico max
Coeff. di sicurezza
Coeff. di sicurezza a fatica
Riduzione di peso
Al-5052
CFRP
+39.3%
+46.4%
invariato
invariato
-0.8%
+2,31 %
41.6 MPa 37.1 MPa
3.2
4
3.2
3.2
49.8%
69.7%
s
Tab. 3
Dimensionamento barra antirollio e riduzione di peso.
Dimensioning of the anti-roll bar and weight reduction.
Riduzione di Indice di prestazione
peso stimata
utilizzato
Molla sospensione
30-35%
σf2/ρE
anteriore
Barra antirollio
50-70%
G1/2/ρ
posteriore
Per quanto riguarda il braccio anteriore inferiore non è immediato poter fare stime di peso affidabili: molto dipende infatti
da quale sollecitazione prevalga e quale sia il vincolo progettuale
più stringente. Comunque analizzando tutti e quattro gli indici di
prestazione si può ipotizzare un range di diminuzione del peso
che va dal 5 al 30 %, a seconda di quale sia la prestazione
da massimizzare.
BIBLIOGRAFIA
[1] Dati tratti da: European aluminium association: www.eaa.net
[2] M. F., ASHBY, La scelta dei materiali nella progettazione industriale, CEA, (2007).
[3] W.L. JOHNSON. Mater. Sci. Forum 225–227 (1996), p. 47.
[4] W.J. YU and H.C. KIM, Compos Struct 9 (1998), p. 279–300.
[5] M.F. ASHBY and D. CEBON, Case studies in materials
selection, Granta Design Ltd, Cambridge, UK, 1995
ABSTRACT
MATERIALS SELECTION CRITERIA FOR THE WEIGHT
REDUCTION OF HIGH PERFORMANCE SPORTS CAR
Keywords: material selection, steel, composite materials, titanium
and alloys
For high-performance luxury cars, the first element of competitiveness is given
by the dynamic performances. Thus, weight reduction is becoming increasingly
important, also in the framework of the introduction of more and more restrictive regulations on emissions. Designers are currently focusing on maximizing
the power/weight ratio, and the proper material selection becomes mandatory.
This paper describes the weight reduction achievable on the suspension system
of a Lamborghini Murcielago, showing that the use of high strength low alloy
steels as well as titanium and aluminum alloys allows to reduce weight of
selected components of more than 30%.
Three different components of the suspension system have been analyzed: the
6
front lower arm (fig. 1), the front suspension spring (Table II) and the anti-roll
bar (fig. 5). According to the method proposed by M.F. Ashby [2], the design
requirements have been translated into a series of constraints, objectives and
free variables, leading to the definition of material indexes (equations 1-8) to be
used to rank the materials in order to identify the most suitable ones. Ranking
was accomplished using the CES (Cambridge Engineering Selector, Grantadesign) software, which allows to create selection charts like the ones shown in
Fig. 2, 3, 4, 6. Those diagrams help the designer to select graphically the candidate materials for the application, which, for the studied cases, resulted to be:
- Low alloy steels or aluminum alloys (Table II) for the front lower arm
- Titanium or spring steel(fig. 1), for the front suspension spring
- Aluminum alloy for the anti-roll bar
Dimensioning of the components on the basis of the selected materials is presented as well, demonstrating that a proper materials selection procedure allows to significantly decrease the weight of the studied components with respect to the current solutions.
Metalli leggeri
Memorie >>
INNOVATIVE TECHNOLOGIES IN
MOULD RELEASE AGENTS
G.Natesh, A. Colori
The drive for improved fuel efficiencies in the automobile industry has led to continuing growth in aluminium
die casting as manufacturers strive to reduce the weight of automobiles by replacing steel with light metal components. Larger and more complex parts are being cast and this has set new challenges to die casters in their
quest for improved quality and productivity. The paper examines the impact of these trends on die lubrication
and discusses an innovative lubrication technology that has evolved to satisfy these requirements.
KEYWORDS: : mould release agent, die-casting, high temperature, automotive, die lubricant, solder protection,
leidenfrost effect
High Pressure die-casting is a very popular process for making
complex mechanical parts out of light metals like aluminium
and magnesium alloys. It is capable of rapidly producing parts
with high dimensional accuracy. High pressure die-casting
grew along with the growth of the automobile industry, where the demands of assembly line manufacture spurred the
demand for a quick reliable way to make components. With
the growth of JIT manufacturing, the automobile industry still
continues to be the dominant user of high-pressure die cast
parts. Other end uses for die casting include recreational vehicles, power tools, electrical machinery, electronic components
and house-ware. The rapid growth of the world economy has
spurred a demand for all of these products and the European
die cast industry is gearing up to meet this demand.
The rising cost of fuel and increasingly stringent environment
and fuel performance regulations are forcing the auto industry to seek novel ways to achieve these goals. Weight reduction of vehicles is a key step to reducing fuel consumption, so
the industry is actively looking at replacing steel components
with aluminium and magnesium castings. With constant innovation in aluminium alloys and casting technology, improved
strength and other properties are being engineered, that allows
bigger and more complex parts to be die cast. Engine blocks,
instrument panels and complete door frames are just some of
the examples of aluminium components now being produced
by die-casting. This has led to a trend towards bigger die-cast
machines and larger shot weights.
The complexity of these large parts makes it difficult to design
internal cooling to adequately cool all parts of the die uniformly. A natural consequence of this is that die surface temperatures have increased. Previously, the die surface temperatures
before spray used to range between 250°C to 350°C. With the
estimenti di allumina, corrosi
one, leghe di alluminio, LEIS, EIS, ENA
mina, corrosi
one, leghe di alluminio, LEI
mina, corrosi
one, leghe di alluminio, LEI
large components, the maximum temperature can be as high
as 400°C while the cooler portions of the die may be as low as
200°C.
This leads to the development of localized hot spots which, in
turn, create solder problems. This places a greater dependence
on the die lubricant to provide cooling for the die surface. Yet
the higher temperatures encountered before the spray make
this difficult to do because of the Leidenfrost effect. This requires greater quantities of die lubricant to be sprayed, which
increase cycle times and costs.
The Leidenfrost phenomenon is well known to die casters.
When water is sprayed on to a hot surface, which is at a temperature well above the boiling point of water, it is unable to
make contact with the metal surface. Instead, the drops of water float on a cushion of water vapor and thus are unable to
wet the surface (Fig. 1). Die lubricant active materials are therefore, unable to be laid down on the die surface. The highest
temperature at which water, or a water based die lubricant,
can contact the metal surface is known as the Leidenfrost temperature.
Our research focused on two separate approaches to develop
high performance die lubricants. The first was to try and in-
s
Fig. 1
Schematic rendering of the Leidenfrost phenomenon.
Rappresentazione grafica del fenomeno di Leidenfrost.
1
Metalli leggeri
s
Fig. 2
Schematic representation of operating window
for die lubricants.
Rappresentazione schematica della finestra operativa per
i lubrificanti dello stampo.
crease the Leidenfrost temperature. By allowing the die lubricant to wet the surface at a higher temperature, film formation begins earlier, allowing shorter spray durations. The
second approach was to develop materials that would form a
film rapidly on the die surface at elevated temperatures. The
cooling curve of water shows the rate of cooling approaches
a maximum at a temperature known as the Nukiyama point.
(Fig. 2) [1]. By increasing the Leidenfrost temperature, we are
able to increase the operating window to form a film. However, if the film formation was slow at the higher temperatures,
any advantage gained by raising the Leidenfrost temperature would be lost. Many different factors affect the Leidenfrost
temperature. Mechanical factors like the distance and angle of
sprays, the size of the droplet and impact pressure all affect the
wetting temperature. Last, but not the least, constituents in the
spray can affect Leidenfrost temperature. [2]
Studies with water showed that the presence of dissolved salts
caused an increase in Leidenfrost point. Fig. 3 shows cooling
profiles with different fluid compositions. We see that DM water had a wetting temperature of about 315°C under experimental conditions. Soft water wets at about 320°C while hard
<< Memorie
water has a Leidenfrost point of about 340°C. Conventional
die lubricants made with soft water show similar results. Note
that these temperatures are higher than what is seen in the
field due to differing operating conditions.
We decided to screen a variety of inorganic and organic compounds and our efforts were able to develop materials that
produced significant increases in the Leidenfrost point. The
comparable values under the same conditions were in excess
of 370°C.
The benefit to a die caster is clearly seen. If a die is initially at
a high temperature and needs to be cooled to about 250°C to
get complete solidification, using these new materials could
reduce the spray time by 20% to 30%. This directly translates
into an increase in productivity, compared to conventional lubricants.
Cooling is only one of the functions of the die lube. The primary function is release, and this can only take place if an
adequate amount of the lubricant film is formed over the die
surface. Lubricant components do not always adhere to very
hot surfaces. Materials were screened for high temperature
film formation by weighing the amount of film formed by
spraying a dilute spray containing a fixed quantity of actives
onto a pre-weighed stainless steel plate maintained at specific
temperatures. Stainless steel was used as experiments showed
that H13 tool steel became oxidized and created a large source
of error in the observations. The use of stainless steel minimized this error.
In an actual die, there is wide variation in the temperature of
the die surface. When die lube is sprayed on the surface, film
can be formed rapidly at the cooler parts of the die, but may
not form very well at the hotter regions. Our initial tests tried
to maintain a single plate with a standard temperature gradient similar to what may be encountered on a die. We were
unable to get repeatable temperature gradients so we decide
to measure the die adhesion at specific temperatures. We then
compared the performance of different materials at the range
of temperatures commonly encountered on the die.
Fig. 4 shows the Hot Die Adhesion Index, which is ratio of the
weight of film formed at 350°C and 250°C. This index is a measure of the uniformity of the film formed on a tool operating
at different temperatures. A value of 100% means exactly the
same amount of film is formed at all temperatures. In actual
practice this is very difficult to achieve. The results show that
the new materials were at least 2 to 3 times more efficient than
conventional die lubricants. These results were validated, by
carrying out field trials on full-sized machines at customer sites. Field performance corresponded well with our experimental results.
s
Fig. 3
s
Fig. 4
Die cooling curve.
Curva di raffreddamento dello stampo.
Hot Die Adhesion Index.
Indice di adesione a caldo dello stampo.
2
Metalli leggeri
Memorie >>
a
s
Fig. 6
Thermal image of die before spray.
Immagine termica dello stampo prima della spruzzatura.
b
s
Fig. 7
Thermal image of die after spray (old product).
Immagine termica dello stampo dopo la spruzzatura (vecchia tecnologia).
s
Fig. 5
a) Old product (after 8 hours); b) New Product
(after 8 hours).
a) Vecchia tecnologia (dopo 8 ore); b) Nuova tecnologia
(dopo 8 ore).
The first case is from a North American die caster making
engine blocks with steel cylinder inserts on a 3500 T Ube®
machine with a total cycle time of less than 120 seconds. They
were getting good solder protection, with low overspray and
in-cavity buildup with a conventional die lubricant. When
they started casting a new engine design, they noticed solder
formation near the water jacket area on the part. This required them to do die polishing for about 30 minutes once every
8 hours. Running at richer concentrations did not help, giving buildup that also needed to be polished. Increasing the
spray duration also had a major impact on productivity.
We took thermal images of the die before and after spray to
monitor temperature profiles and spray distribution. The temperature ranged from 450°F to 750°F (232°C to 399°C) before
spray on the ejector die. We also observed that the previous
product was not covering the problem area adequately particularly at the high temperature zones. The new Safety-Lube®
product could wet the hot surface earlier providing better co-
verage and could rapidly form an adequate lubricating film
at the high temperatures seen on the die. Fig. 5 shows the
dramatic reduction in solder. The improved performance
from the Safety-Lube® product eliminated the need to polish
every shift and reduced the cleaning time by 50%.
The second example is from a European die caster making
automotive components, who were extremely concerned
about the long cycle times needed to make a particular casting. They also had problems with porosity, soldering and
in-cavity build-up which led to poor yields and productivity.
Investigation of the problem revealed a clear pattern. Fig. 6
shows that the die temperatures before spray ranged from
417°C to 230°C across the face of the die.
The incumbent product was unable to form adequate protective film against solder, therefore a long spray time was needed. However, this caused some areas of the die to be cooled
excessively, leading to in-cavity build and porosity. Reducing
the spray time gave solder, and in both cases downtime was
needed to do die polishing. As can be seen in Fig. 7, the typical die temperatures after spray with the conventional product ranged from 250°C to 160°C.
From this analysis, it was clear that we needed to form good
3
Metalli leggeri
<< Memorie
or build-up was seen, reducing the downtime for die maintenance from 3 hrs per day to 1 hour every 2 days. Because
of the shorter cycle time, the daily yield was also increased
by 15%.
These products have now been launched commercially and
acceptance of this technology has been outstanding. We have
a range of products designed for a wide range of castings.
By carrying out a comprehensive survey of the system and
customer-specific issues, we are able to ensure that the most
suitable products can be recommended.
The die casting industry is rapidly changing to meet the new
demands of their customers, New alloys are being developed
to meet the ever- increasing demands of a high strength to
weight ratio. We continue to develop new technology so we
can maintain our leading position as partners to the die cast
industry,
s
Fig. 8
Thermal image of die after spray (new product).
Immagine termica dello stampo dopo la spruzzatura
(nuova tecnologia).
solder protection at the high temperature areas quickly, so
that excessive cooling of the die did not take place. A trial
was run with the new high temperature product. Based on
the product capabilities we reduced the spray time by 27%.
The images in Fig. 8 show the die temperatures after spray
ranged from 300°C to 215°C. Most importantly, no soldering
REFERENCES
1] “Simultaneous Measurements of droplet characteristics
and surface thermal behaviour to study spray cooling with
pulsed sprays” by Humberto M Loureiro, Miguel R A Panao,
Antonio Luis L Moreira of Mechanical Engg Dept Instituto
Superior Tecnico at Lisbon Portugal.
2] “The effect of die lubricant spray on the thermal balance of
dies” by Dr James L. Graff, Chem-Trend & Dr Lothar H. Kallien, presented at the NADCA conference in Cleveland,USA
Oct 18-21, 1993
ABSTRACT
TECNOLOGIE INNOVATIVE RELATIVE A AGENTI DI
DISTACCO DALLO STAMPO
Parole chiave: metalli leggeri, pressocolata, lubrificazione
La spinta verso il miglioramento dell’efficienza energetica nel settore
automobilistico ha portato alla continua crescita dell’impiego della
pressofusione dell’ alluminio, in quanto i produttori si sforzano di
4
ridurre il peso delle autovetture sostituendo parti in acciaio con componenti in metallo leggero. Con questo processo vengono prodotte
parti sempre più grandi e più complesse e ciò ha posto nuove sfide per
gli specialisti di pressocolata nella loro ricerca di migliorare qualità e
produttività. Il presente studio esamina l’impatto di queste richieste
sulla lubrificazione degli stampi e presenta una tecnologia di lubrificazione innovativa che è stata messa a punto per soddisfare nuovi
requisiti.
Memorie >>
Trattamenti superficiali
MODERN THERMAL ELECTRON BEAM
PROCESSES – RESEARCH RESULTS
AND INDUSTRIAL APPLICATION
R. Zenker
Thermal electron beam (EB) technologies are becoming more and more attractive especially
because they are ecologically friendly and energy saving on the one hand and highly precise, excellently
controllable and highly productive on the other hand.
Using three-dimensional energy transfer fields, the interaction conditions between the EB and the surface
of the material, the conditions of the heat conduction in the material, the geometry of the part, and the load
conditions of the component can be taken into account. High flexibility, precision, and reproducibility are
typical characteristics of EB technologies and facilities. High productivity is achieved by new technological
solutions like simultaneous interaction of the EB in several processing areas (spots) or by carrying out several
processes simultaneously in modern EB facilities and systems (such as multi-chamber, lock-type and other
concepts). The influence of beam parameters and energy transfer conditions on the microstructure of the
materials and its properties will be discussed for different EB technologies. Information on ideal treatment
conditions will be given. The paper deals with the current development state regarding beam deflection
techniques, technological processes and some facility concepts, and with the state of industrial application.
KEYWORDS: electron beam processes, surface treatment, combined technologies, welding, engraving, profiling,
beam deflection techniques, materials (structure-property relation), applications
INTRODUCTION
By using the advantages of EB, modern EB technologies differ
from other technologies in their advantageous characteristics
(Tab. 1).
These characteristics are typical for all EB technologies, i.e.
welding, surface treatment, surface ablation or perforation,
which are carried out as one-spot techniques. If a multi-spot
technique or multi-process technology is applied, the effects of
these characteristics are much more effectively.
The present paper will exemplarily demonstrate the state of
the art of development and application of EB technologies.
MULTI-TOOL ELECTRON BEAM
Beam deflection techniques
The development of the two-dimensional high-frequency
Rolf Zenker
TU Bergakademie Freiberg, IWT,
Zenker Consult, Mittweida, Germany
Paper presented at the European Conference „Innovation in heat
treatment for industrial competitiveness”, Verona, 7-9 May,
organised by AIM
beam deflection technique was the beginning of a new area
of thermal electron beam (EB) technologies. The high speed
scanning (HSS) technique has been available since 1986 [1]-[3].
In 2000 a high frequency 3D beam deflection technique was
created with new possibilities for load and couture specific EB
technologies, not only for surface treatment [4]-[10] but also
for welding [7][10]-[13] and engraving [14]-[15]. These beam
deflection techniques are based on the fact that the EB can act
simultaneously in several spots [10]-[13]. In this case the same
task is realised in every spot.
A defined and exact positioning of the (mostly oscillation)
Electron beam (EB)
excellent formability and deflect ability
good beam profile
high efficiency
large penetration depth
high beam stability
EB technologies
high productivity
excellent flexibility
good process safety
high reproducibility
ecologically friendly
s
Tab. 1
Characteristics of EB and EB technologies (behind
others) .
Principali caratteristiche del fascio elettronico e delle tecnologie EB.
la metallurgia italiana >> aprile 2009
1
<< Memorie
EB is the prerequisite for the successful realisation of multispot techniques (Fig. 1). The “spot size”, i.e. the interaction
area of EB depends on the process and is very small in case
of engraving (some microns in diameter) and reaches a dimension up to ~ 100 mm2 for surface treatment. The spot
size in welding is greater than for engraving but smaller
than for surface treatment.
In case of the multi-spot technique, the same task is realised
in every spot. It can be differed between two basic beam
deflection techniques. During beam interaction
- the beam or/and the component is/are moving relatively
one against to the other (Fig. 1).
- the beam and the component are fixed in their positions,
so called “flash technique” (Fig. 2).
s
Fig. 1
MULTI PROCESS TECHNOLOGIES
Beam deflection techniques with movement of
EB and/or component [16].
Tecniche di deflessione del fascio di elettroni con movimento del fascio stesso e/o del componente [16].
Single treatment often does not fulfil all the demands made
on the properties of a component which makes additional
treatment necessary. In other cases the material requires
special additional treatment (re-heating, subsequent heating) in order to avoid undesired effects (distortion, cracking).
Making comprehensive use of the advantages of EB in this
connection means that several EB processes are implemented simultaneously. Now, there are further new and excellent multi-process technologies available, i.e. a combination of different processes in one production run is possible
and there is also the possibility of an online process control
[9][10][14]-[15].
At present there are two and three-process technologies
available (Fig. 3).
The advantages of such technologies are:
- high productivity
- new possibilities for influencing material structure and
properties (new very short local thermocycles).
s
Fig. 2
EB SURFACE TREATMENT
GD-OES depth profile of a nickel-boron coating
on Steel after 1 hour heat treatment at 400°C under
neutral atmosphere.
Profilo GD-OES nello spessore di un rivestimento nichelboro su acciaio dopo 1 ora di trattamento termico a
400°C in atmosfera inerte.
s
Fig. 3
EB multi-process technologies (examples).
Esempi di tecnologie EB multiprocesso.
2
aprile 2009 << la metallurgia italiana
Surface hardening and low tempering
(controlling component)
EB hardening (EBH) is the most successful EB surface technology used in industry. Lately, more and more multi-process technologies were applied [9][10][14][15].
Fig. 4 exemplarily shows the EBH and tempering of the
contour of a controlling camshaft in one process cycle [10].
Two energy transfer fields (hardening and tempering field,
Fig. 4b) interact simultaneously with the rotating cam.
Because of the stronger load conditions the depth at the
radius must be larger (0.65…0.75 mm) than at the flanks
(0.55…0.65 mm).This is programmable without any problems.
Another technically and economically very attractive EBH
technology is applied in case of a calotte carrier [9][10] (Fig.
4a). The surface contour is programmed as a rotation-symmetric energy transfer field with a surface contour congruent energy distribution (Fig. 4b). The resulting hardening
profile is characterised by constant EBH thickness that is
independent of the incidence angle of the EB (Fig. 4b).
The energy transfer is realised by flash technique. During
the interaction of the EB (≤ 1.0 s) the component is fitted to
the beam before crossing the α/γ transformation temperature (processing time ≤ 0.2 s). This technology guarantees a
high productivity (up to 3.500 parts per hour).
Memorie >>
a
b
c
s
Fig. 4
EB hardening of a power train component (two-process EBH technology) - a) Controlling component; b) Twoprocess technology; c) EB hardening depths.
Indurimento superficiale mediante EB di un componente di un sistema di potenza (tecnologia EBH).
a
b
c
s
Fig. 5
EB Hardening of a spherical surface (flash technique) - a) Calotte carrier; b) Energy transfer field; c) Process
thermocycle.
Indurimento superficiale mediante EBH di una superficie sferica (tecnica flash).
Combined surface treatment (tools, automotive components)
Combination EBH/Nitriding
With regard to complex load conditions for most tools and
components, especially close to the surface, the properties attainable by single treatments (mechanical, thermal, thermochemical and coating technologies), in particular, often are
insufficient. Therefore, combined processes (duplex or hybrid
process) (Fig. 6a) came into the focus of examinations and
meanwhile of industrial application [17]-[20].
In case of the sequence of the combined surface treatment
combination EBH+nitriding (N) or nitrocarburising (NC) the
level of the processing temperature of N or NC in relation to
the tempering temperature of the bulk material determines the
success of this treatment combination.
The better the tempering stability of the steel the smaller is the
hardness reduction in the previously produced EBH layer as a
result of the subsequent nitriding process.
It is true that a subsequent EBH after N (NC) transforms the
compound layer partially (wider seam of pores), but the hardness of the diffusion layer is higher than after EBH or N (NC)
[21]. In case of the component shown in Fig. 6b hardness rises
by ~ 200HV0.3 (Fig. 6c).
It has been shown that in the case of optimised process parameters the advantages of this combined treatment complement
each other and the disadvantages of the single processes cancel each other out at least partially [20][21].
Combination of EBH and HC
Hard coatings based on titanium, aluminium or chromium
carbides are successfully applied as hard wear resistant layers
for tools and components. These hard but also brittle coatings
are often unable to bring their excellent properties fully to bear
on relatively soft base materials. Therefore the base materials
are usually subjected to additional heat treatment before or after hard coating [22]-[24]. It is possible to limit the heat treatment to the highest loaded areas and up to the depth where a
martensitic transformation is necessary. The thermal loading
of the overall component is minimised.
With regard to a subsequent heat treatment of hard coated
steels it allows for prevention of undesirable changes of composition, structure and properties of the hard coating [25][26].
A very short interaction time and the process-related vacuum
support these effects. Moreover, the electron beam hardening
technology is well known to cause small changes in size and
shape which means that distortion is also reduced in that way
also. A combined EBH+HC is successful only if the treating
temperature of the hard coating process is lower than the tempering temperature of the bulk material [26].
3
a
b
<< Memorie
c
s
Fig. 6
Combination of EBH with nitriding (N) - a) Technological variants of combined surface treatment; b) Component (detail) steel 54CrV4 (N + EBH); c) Hardness profiles after combined treatment N + EBH (steel 100Cr6).
Applicazione combinata della tecnologia EHB e della nitrurazione (N).
a
b
c
s
Fig. 7
Combination of EBH and hard coating (HC) - a) Surface layers after HC and HC + EBH (tool steel); b) Surface
hardness and critical load of TiAlN coatings on C45 (H+T) (HC + subsequent EBH); c) Surface profiles after TICN +
EBH (100Cr6).
Applicazione combinata della tecnologia EHB e del rivestimento duro.
a
b
c
s
Fig. 8
EB three-spot welding - a) Gear wheel; b) Three-spot welding; c) Detail of welding detail.
Saldatura EB a tre fasci.
A subsequent EB heat treatment after HC has no significant
influences on the visual appearance and the structure of the
coatings (Fig. 7a). The achievable surface hardness (Fig. 7c)
and the hardening depth profile depend on the chemical
composition and the pre-heat-treated state of the base material and on the beam hardening conditions. A martensitic
4
layer beneath the hard coating layer is produced, resulting
in a significant improvement of the bulk material’s load support for the hard coatings. Therefore high surface hardness
and high critical loads measured by scratch tests are obtained
(Fig. 7b). Additionally, the properties gradient is improved
distinctly.
Memorie >>
Regarding the application of the HC+EBH combination the
surface deformation due to the martensitic transformation
must be taken into consideration (Fig. 7c).
EB WELDING
Welding of steel (powertrain components)
A typical powertrain welding unit is shown in Fig. 8a). The
conventional procedure used for the EB or laser welding of
gear components is a tack welding in a first step and then
the joining of the two welding partners by single-pool welding in a second step.
By contrast, in multi-pool welding the components are fixed
simultaneously at the time of the first interaction of the EB
with the material at several points (Fig. 8b).
It follows in the same processing step a simultaneous movement of several melting pools along one and the same welding seam (Fig. 8b) up to an overlapping zone in the area
which has already been welded. The number of welding
pools depends on the size and geometry of the part and on
the deflection width of the EB. In comparison to the above
mentioned two-step technology the welding time is reduced
up to one third and also the distortion is minimized because
of the lower heating of the parts [7][10][13].
A speciality of multi-pool EB welding is that the welding
seam is not perpendicular to the surface (Fig. 8b). The incidence angle of the EB and, consequently, the seam depend
on the welding diameter of the welding circle and the distance between beam source and component (Fig. 8c) [10]
[13].
One important fact for a successful application of multi-pool
welding is that the program must be optimized in relation to
the jump frequency of the beam from one melting pool to the
next and the beam oscillation for an open vapour capillary.
Welding of al alloys (cylinder liner ensembles)
The production of engine blocks as hybrid casting is state of
the art. In the automotive industry cast iron cylinder liners
are usually applied but there are also cylinder liners made of
spray-formed Al materials. The cast engine block either consists of different Al or Mg alloys.
One of the technical difficulties is the precise positioning of
each cylinder liner in the mould, one after the other. A technologically more smarter and more profitable technology is the
a
b
positioning of the liners as so called “liner ensemble” (Fig. 9a).
In this case several (2…6) cylinder liners are assembled by EB
welding and positioned in the mould as an ensemble in one
step. The technical expenditure is much lower [10][15].
Electron beam welding for that application is realised with
two-spot techniques using a welding spot and a smoothing
spot in one run (Fig. 9c, d).
It has to be taken into account that water jackets are integrated
between the cylinder liners which must stay open after welding. Because of the high penetration depth of the EB, the welding takes place only from one side across the water channel up
to a depth of 45 mm without closing it (Fig. 9c, d). The diameter of the hole must be ≥ 3.0 mm, but this is normal standard
design. The fact that the liners are welded from one side contributes to a very economical production [15].
The application of the two-spot (pool) technique is necessary
because most spray formed alloys cannot be welded easily and
have a very rough welding bead. The task of the second spot
is to smooth the bead.
EB SURFACE ABLATING
Engraving (shaft for force fit with tube)
EB engraving follows the well-known method of producing
lateral surface patterns to improve the sliding conditions [27]
[28], produce reservoirs for colour particles [29], texture the
cold rolls to improve sheet quality [30] and - in this present
case - to increase the friction in force fits [13].
The principle of these different processes is the same (Fig. 10).
At first, the EB with a small diameter remelts the surface in
a small pool (Fig. 10a). Then a vapour capillary is produced
because of the high beam energy. The vaporised material and
some of the liquid material squirt out of the capillary and a
molten shell is formed around it (Fig. 10b). Depending on the
material and the beam parameters dimples and/or protrusions develop (Fig. 10c).
By applying the EB multi-spot technique, many protrusions
can be generated simultaneously around dimples, as spot lines
(up to 200 spots per line, Fig. 11a) or as patterns (on a plane
surface up to 3.500 spots) during less than 0.15 ms [31].
Large protrusions (Fig. 11b, c) are desired and necessary for
force fits of shaft/tube assemblies. The dimples (Fig. 11c) are
necessary because they prepare the material for the protrusions.
c
d
s
Fig. 9
Two-spot welding of cylinder liner ensembles - a) Cylinder liner ensemble; b) Single cylinder with water channels; c) Welding seam (schematic); d) Welding seam.
Saldatura a due fasci di un complesso di elementi cilindrici allineati.
5
a
b
<< Memorie
c
s
Fig. 10
Principle of engraving.
Schema del processo di incisione.
Profiling (bearing inserts)
In this special application (hybrid casting) used to improve
the contact between insert and casting the profiles are much
deeper (0.1…0.5 mm) than in the above mentioned applications and have a strong melting seam at the side or around the
depression [15]. The EB acts on the surface of the component
with a high energy density (105…106 Ws/cm²), causing the
material to liquefy and/or vaporise locally within an extremely short time without having a significant thermal effect on the
surrounding area.
This EB technology is now developed for bearing shell inserts
(Fig. 12a) and cylinder liners made of Al materials [15].
It is of importance for the result of the process whether an alloy is present without (pure metals or eutectic alloy) or with a
large solid/liquid melting range (hypo-/hypereutectic alloy).
In the latter cases (Fig. 12b, c) a relatively wide re-melted border forms on the edge of the profile (fusion shell).
Depending on the casting (Al or Mg alloy) there are differences
in the effect of profiling. As a basic principle, a distinction has
to be made between:
- mechanical interlocking (Tm, Insert > Tm, Casting) and
- metallurgical connection in combination with mechanical interlocking (Tm, Insert ~ Tm, Casting)
As it is known, further factors such as the casting method and
conditions, temperature control, the position in the mould, etc.
affect the result with regard to the form fit and connection between insert and casting.
If the insert is made of an Al alloy as well as the casting the contact between insert and casting is generally good (mechanical
interlocking and metallurgical connection, Fig. 12c). Ultrasonic
measurements support these results. In areas where there is no
contact the ultrasonic waves are reflected at the interface. In
case of metallurgical connection there is no or a weak signal
(Fig. 12e).
To assess the bond strength of castings with profiled inserts,
samples were analysed using static tensile tests. The analyses
showed that a bond strength which was 3…10 times higher
than it was for non-profiled inserts (1…2.5 kN) was achieved
by EB profiling [15] (basis: force up to the point where the insert/casting connection is separated). This may be attributed
in particular to the fact that the breaking of samples did not
occur at the interface between insert and casting, as it was the
case in non-profiled reference samples, but predominantly in
the casting.
6
s
Fig. 11
Multi-spot engraving of shaft for force fits with
tubes.
Incisione a più fasci di un albero per accoppiamento
forzato con tubi.
The method of EB profiling is a promising alternative to mechanical profiling or coating. The EB method is much faster
than the other known technologies of surface “roughening”.
SUMMARY AND CONCLUSIONS
- The electron beam can be used effectively in a wide range of
surface, joining and removing technologies in all branches of
the metalworking industry.
- The availability of new beam guidance techniques is an essential prerequisite for a more and more perfect adaptation of
beam handling for technological applications.
- Further progress in EB technologies comes into effect by
multi-spot beam deflection techniques and/or multi-process
technologies.
- Load, material and component specific technological solutions for EB surface treatment, welding and surface ablation
open up new fields for the highly intelligent, flexible and productive EB.
- In the very near future, further new and unconventional design and technical solutions will be available for industrial application.
REFERENCES
1] SCHILLER, S.; PANZER, S.: Thermal surface modification by HF-deflected electron beams. In: Proceedings of the
Conference on the Laser VS the Electron Beam in Welding,
Cutting and Surface Treatment: State of the Art, Reno, 1985,
part 2, pp. 16-32
2] SCHILLER, S.; PANZER, S.: Härten von Oberflächenbahnen mit Elektronenstrahlen. In: HTM 2(1987), 5, pp. 293-
Memorie >>
a
b
d
s
Fig. 12
Multi track profiling of bearing insert.
Profilo a più tracce di un particolare di cuscinetto.
300
3] PANZER, S.; MÜLLER, M.: Härten von Oberflächen mit
Elektronenstrahlen. In: HTM 43(1980), 2, pp. 103-111
4] ZENKER, R., Electron beam surface treatment: industrial
application and prospects. In: Surface Engineering 12(1996),
4, pp. 296-297
5] ZENKER, R.; WAGNER, E.; FURCHHEIM, B.: Electron
beam – a modern energy source for surface treatment. In:
6th International Seminar of IFHT: Advanced Heat Treatment Techniques Towards the 21st Century: 15.-18.10.1997,
Kyongju, 1997
6] ZENKER, R.; FRENKLER, N.; PTASZEK, T.: Electron
beam surface treatment of Al, Mg, and Ti alloys. In: Proceedings of the 7th International Seminar of IFHT: Heat treatment and surface engineering of light alloys: 15.-17.9.1999,
Budapest, 1999, pp. 18-21
7] ZENKER, R.: Electron beam surface treatment and multipool welding – state of the art. In: EBEAM 2002, International Conference on High-Power Electron Beam Technology:
27.-29.10.2002, Hilton Head Island, 2002, pp. 12-1–12-5
8] ZENKER, R.: Structure and properties of electron beam
surface treatment. In: Advanced Engineering Materials
6(2004), 7, pp. 581-588
9] ZENKER, R.: Elektronenstrahl-Randschichtbehandlung,
Innovative Technologie für höchste industrielle Ansprüche.
Monographie, pro-beam AG & Co. KGaA, 2003
10] ZENKER, R.: Elektronenstrahlbearbeitung für Powertrainkomponenten. In: Kooperationsforum Metalle im Automobilbau, Innovationsforum in Be- und Verarbeitung,
29.11.2005, Hof, 2005
11] MATTAUSCH, G.; MORGNER, H.; DAENHARDT, J.;
ET. AL.: Survey of electron beam technologies at FEP. In:
Proceedings / EBEAM 2002: International Conference on
c
e
High-Power Electron Beam Technology, Hilton Head Island, 27.-29.10.2002, pp. 11/1-11/11
12] LOEWER, T.: Analysis, visualisation and accurate description of an electron beam for high repeatability of industrial production processes. In: Proceedings of the 7th
International Conference on Electron Beam Technologies,
Varna, 1.-6.6.2003, pp. 45-50
13] ZENKER, R.; BUCHWALDER, A.; FRENKLER, N.;
THIEMER, S.: Moderne Elektronenstrahltechnologien zum
Fügen und zur Randschichtbehandlung. In: Vakuum in der
Praxis, 17(2005), 2, pp. 66-72
14] ZENKER, R.; BUCHWALDER, A.; SPIES, H.-J.: New
electron beam technologies for surface treatment. In: Proceedings of the 7th International Conference on Electron
Beam Technologies: 1.-6.6.2003, Varna, 2003, pp. 202-209
15] ZENKER, R.; KRUG, P.; BUCHWALDER, A.; DICKMANN, T.; FRENKLER, N.; THIEMER, S.: Elektronenstrahlschweißen und –profilieren von sprühkompaktierten
Zylinderlaufbuchsen aus Al-Si-Werkstoffen. In: Zylinderlaufbahn, Kolben, Pleuel – Innovative Systeme im Vergleich,
Tagung Böblingen, 7.-8.03.2006, VDI-Verlag GmbH: Düsseldorf, 2006, VDI-Berichte 1906, pp. 259-274
16] BUCHWALDER, A.: Beitrag zur Flüssigphasen-Randschichtbehandlung von Bauteilen aus Aluminiumwerkstoffen mittels Elektronenstrahl. Dissertation TU Bergakademie
Freiberg, 2007
17] ZENKER, R.: Kombinierte thermochemisch-thermische
Wärmebehandlung. Neue Hütte 28(1983), 10, pp. 379-385
18] SPIES, H.-J.: Erhöhung des Verschleißschutzes von Eisenwerkstoffen durch die Duplex-Randschichttechnik.
Stahl und Eisen 117(1997), 6, pp. 45-62
19] KEßLER, O., HOFFMANN, F., MAYR, P.: Combinations
of coating and heat treating processes: establishing a system
for combined processes and examples. Surf. Coat. Technol.,
108-109(1998), pp. 211-216
20] ZENKER, R., SPIES, H.-J.: 15 Jahre industrielle Anwendung der Elektronenstrahl Randschicht-behandlung. 57. Härtereikolloquium, Wiesbaden, Germany, Oct 10-12, 2001
7
21] SPIES, H.-J., ZENKER, R., BERNHARD, K.: DuplexRandschichtbehandlung von metallischen Werkstoffen mit
Elektronenstrahltechnologien. Härtereitechn. Mitt. 53(1998),
4, pp. 222-227
22] SPIES, H.-J.; HOECK, K.; BROSZEIT, E. MATTHES, B.;
HERR, W.: PVD hard coatings on prenitrided low alloy
steel. Surf. Coat. Technol. 60(1993), pp. 441-445
23] HÖCK, K.; SPIES, H.-J.; LARISCH, B.; LEONHARDT,
G.; BUECKEN, B.: Wear resistance of prenitrided hardcoated steels for tools and machine components. In: Surface
Coatings and Technology 88(1996), pp. 44-49
24] KESSLER, O.: Combination of coating and heat treatment
processes. Surf. Coat. Technol. 201(2006), pp.4046-4051
25] SPIES, H.-J.; FRIEDRICH, S.; BUCHWALDER, A.: Elektronenstrahlbehandlung von PVD-Hartstoffschichten.
Mat.wiss und Werkstofftechn. 34(2003), 1, pp. 128-135
26] ZENKER, R.; SACHER, G.; BUCHWALDER, A.; LIEBICH, J.; REITER, A.; HÄßLER, R.: Hybrid Technology Hard
Coating – Electron Beam Surface Hardening. Surf. Coat.
Technol. 202(2007), pp. 804-808
27] ABELN, T.; KLINK, U.: Laserstrukturieren zur Verbesserung der tribologischen Eigenschaften von Oberflächen. In:
Konferenz Stuttgarter Lasertage 2001, pp. 61-64
<< Memorie
28] ABELN, T.: Reibungsminderung durch Laseroberflächenstrukturierung im Motorenbau. In: Zylinderlaufbahn, Hochleistungskolben, Pleuel - Innovative Systeme im
Vergleich, VDI-Berichte 1906, VDI Verlag Düsseldorf 2006,
pp. 203 - 217
29] DE MARE, C.; SCHEERS, J.; LAMBERT, F.; VERMEULEN, M.; DE GRAEF, L.; GADEYNE, Y.: Development of
the SIBETEX sheet having excellent drawability and paint
appearence. In : Revue de Metallurgie, 94(1997), 6, pp. 827836
30] EMERY, C. J., SNAITH, B.: Developments in the texturing of sheet metal surfaces. In: Proc. of 8th Intern. Conf.
Sheet Metal (2000), pp. 337-342
31] ZENKER, R.; BUCHWALDER, A.; FRENKLER, N.;
THIEMER, S.: Electron beam surface shaping/profiling. In:
8th Int. Conf. on EB Technologies, Varna: 5.-10.06.2006
ACKNOWLEDGEMENT
The author thanks his partners, cooperating in the field of
EB technologies, especially pro-beam Neukirchen, PEAK
Velbert and his team at TU Bergakademie Freiberg, Institute
of Materials Engineering.
ABSTRACT
MODERNI PROCESSI TERMICI CON FASCIO
ELETTRONICO - RISULTATI DELLA RICERCA E
APPLICAZIONI INDUSTRIALI
Parole chiave:
Le tecnologie che sfruttano le caratteristiche termiche dei fasci di elettroni
(EB) stanno diventando sempre più attraenti, in particolare perché da un
lato sono ecocompatibili e consentono un risparmio energetico, dall’altro
danno risultati molto precisi, eccellentemente controllabili e altamente
produttivi.
Quando si utilizzano campi tridimensionali di trasferimento di energia,
vanno prese in considerazione le condizioni di interazione fra il fascio
elettronico e la superficie del materiale, le caratteristiche di conduzione
8
del calore del materiale, la geometria del pezzo e le condizioni di carico del
componente. Elevata flessibilità, precisione e riproducibilità sono caratteristiche tipiche delle tecnologie e delle attrezzature EB. L’alta produttività
si è ottenuta grazie a nuove soluzioni tecnologiche, come l’interazione simultanea del fascio elettronico con diverse aree di lavoro (spots) o mediante la realizzazione di diversi processi contemporanei nei moderni impianti
EB (come il multi-camera, il “lock-type” o altre soluzioni).
Nel presente lavoro si discute l’influenza dei parametri del fascio e delle
condizioni di trasferimento di energia sulla microstruttura dei materiali
e sulle loro proprietà, per le diverse tecnologie EB. Vengono anche fornite
informazioni sulle condizioni ideali di trattamento.
Il documento descrive l’attuale stato di sviluppo delle
tecniche di deflessione del fascio, dei processi tecnologici e di alcune soluzioni impiantistiche, oltre a riportare l’attuale stato di applicabilità industriale.
Alluminio e leghe
Memorie >>
SQUEEZE CAST AUTOMOTIVE
APPLICATIONS AND DESIGN
CONSIDERATIONS
Z. Brown, C. Barnes, J. Bigelow , P. Dodd
With an increasing emphasis on vehicle weight reduction, the demand for lighter weight automotive components continues to increase. Squeeze casting is an established shape-casting process that is capable of producing lightweight, high integrity automotive components that can be used for structural applications.
In recent years the squeeze casting process has been used with various aluminum alloys to produce
near-net shape components requiring high strength, ductility, pressure tightness or high wear resistance
[1]. Squeeze casting has proven to be an economical casting process for high volume applications and
offers design and materials engineers an alternative to conventional casting processes such as gravity
permanent mold (GPM), low pressure die casting (LPDC), sand cast aluminum/ iron, and conventional
high pressure die casting (HPDC).
This paper describes Contech’s squeeze casting technology (P2000TM) and provides examples of high
volume automotive components manufactured at Contech. This paper also includes product design
considerations, an overview of process simulation techniques, a comparison of mechanical properties, and case
studies for select automotive components.
KEYWORDS: squeeze casting, aluminum, automotive applications, die casting, safety critical
INTRODUCTION
Conventional HPDC is a well-established process for the
manufacturing of a wide variety of aluminum automotive
components such as engine blocks, pump housings, oil
pans, and transmission components. Conventional HPDC
has many advantages including near-net shape capability,
low manufacturing cost, and excellent dimensional accuracy and repeatability.
Achievable casting performance is limited however, due to defects that emerge during the casting process such as gas and
shrink porosity, laminations, and inclusions. In addition,
HPDC components are not considered heat treatable, which
further limits achievable performance.
For applications that require higher component integrity
(high strength and ductility, reduced porosity, uniform
microstructure, and ability to heat treat), alternative castZach Brown, Chuck Barnes, Joe Bigelow
Contech U.S. LLC
Paul Dodd
Contech UK LLC
Paper presented at the International Conference “High Tech
DieCasting”, Montichiari, 9-10 April 2008, organised by AIM
ing processes such as squeeze casting should be considered.
Squeeze casting is an established process that builds upon conventional HPDC practices and is used to manufacture various automotive components that require high strength and
ductility, as well as applications that require high pressure tightness or wear resistance. Examples include steering column components, steering knuckles, control arms,
suspension links, pump housings, and various powertain
components [1]. The squeeze casting process is capable
of producing components with dimensional accuracy and
near-net shape capability that is equal to conventional HPDC.
Unlike HPDC however, the squeeze casting process is capable
of producing higher integrity components. As a result,
design engineers are able to further optimize current aluminum designs or substitute aluminum in place of heavy materials such as steel and cast iron.
SQUEEZE CASTING TECHNOLOGY (P2000TM)
Squeeze casting can be divided into two categories; “direct”
and “indirect”. Direct squeeze casting, often termed “liquidmetal forging”, consists of pouring metal into a lower die contained within a hydraulic press. The upper die closes over the
lower die and high pressure is applied throughout the entire
solidification process. In contrast, indirect squeeze casting consists of pouring molten metal into the cold chamber of a die
1
Alluminio e leghe
s
Fig. 1
Schematic of P2000TM casting machine.
Schema della macchina di colata P2000TM.
casting machine, ejecting the metal into the cavity at relatively
slow shot speeds, and applying pressure through the shot system during solidification [2].
Contech’s proprietary P2000TM process is considered an indirect squeeze casting process. Fig. 1 shows the schematic of a
typical casting machine. The vertical cold chamber is designed
to tilt back prior to pouring the metal into the cold chamber.
The metal is poured down the sidewall of the cold chamber to
minimize turbulence, thereby minimizing porosity and the
formation of oxide skins.
This is done outside of the casting cycle during the spraying phase so overall cycle time is minimized. The metal is
slowly forced into the preheated die cavity, and pressure is
applied throughout solidification. The slow injection speed
reduces turbulence resulting in minimal air entrapment. The
continuous application of pressure helps minimize shrink
porosity and creates rapid heat transfer at the mold/metal interface resulting in a fine microstructure (small dendrite arm
spacing (DAS) and fibrous silicon morphology). The reduced
amount of shrink and gas porosity, fine microstructure, and
ability to heat treat are factors responsible for the improved
part integrity [3].
The proprietary CONTECH P2000TM squeeze casting process
has been in high-volume production for over 25 years and
has been continuously refined throughout this timeframe.
As a result, the P2000TM casting process takes into account all
factors that can influence the quality of the casting including
die cooling systems, gating and venting configurations,
casting process parameters, die lube selection and application, alloy selection, metal handling, heat treatment, and
secondary operations. If all of these factors are considered during the design and product development phase, components
can be optimized to not only meet functional requirements
but also manufacturing requirements.
DESIGN CONSIDERATIONS
With ongoing emphasis on weight reduction, designers are
challenged with developing components that meet weight
and cost targets, while meeting functional and manufacturing requirements.
Examples of functional requirements include strength,
durability, stiffness, hardness/wear resistance, surface ap-
2
<< Memorie
s
Fig. 2
Design process for aluminum shock mount.
Configurazione del processo per la produzione di un
supporto anti-urto in alluminio.
pearance, and packaging. Manufacturing requirements include castability, dimensional capability, cycle time optimization, tooling reliability, machining stock minimization, and
overall casting quality. Determining the proper balance between these factors can be challenging. It is recommend
therefore that design engineers collaborate closely with casting engineers as early as possible during the development of a
new product.
Design Process
Creating a fully optimized casting design requires multiple
design iterations and analysis techniques. Fig. 2 shows an
example of the design process that was used to convert a
steel stamped shock mount assembly to a single aluminum
squeeze casting. Solid modeling software was used to develop the initial casting models. Finite element analysis (FEA)
was used to optimize the component geometry and ensure
all strength, durability, and stiffness requirements were met.
Process simulation tools were used to ensure manufacturing requirements were met and to indentify potential casting flaws
(i.e. porosity, flow related defects, etc.). The final design
was validated through component testing of the prototype
castings. The aluminum shock mount weighed approximately 30% less than the steel design. The number of individual
stamped and welded components was reduced from seven to
one.
By using both FEA and process simulation tools simultaneously, design engineers can take advantage of the full
material potential, resulting in lighter weight designs. Simulation results can be compared to FEA results to determine if potential casting defects are near high stress regions, potentially
resulting in lower than expected casting performance. In
addition, specific geometries that improve manufacturability
and component integrity can be incorporated into the design
in the early stages of development. By using this type of
approach, design engineers can take full advantage of the
castings true potential.
Process Simulation
Process simulation tools, when used properly, are an effective method of evaluating potential casting integrity, establishing process settings, predicting residual stresses, and determining optimal gating and die cooling configuration.
Alluminio e leghe
Memorie >>
s
Fig. 4
s
Fig. 3
Example of P2000TM squeeze cast knuckle.
Esempio di snodo prodotto con la tecnica di squeeze
castingP2000TM.
Example of an aluminum bearing cap that was
converted from GPM to the P2000TM squeeze cast
process. All secondary machining operations were
eliminated.
Esempio di una calotta in alluminio prodotto mediante
squeeze casting (P2000TM) anzichè in gravità in
conchiglia. Tutte le operazioni secondarie di lavorazione
sono state eliminate.
Solidification simulations are used mainly to predict shrink
porosity and evaluate directional solidification. Fill simulations are used to identify potential fill related issues such
as laminations due to merging flow fronts, turbulence, and
improper venting.
Tooling design engineers rely on these tools when optimizing
gating size and location, cooling line placement, cooling media
and temperature, die configuration, and process development.
New process simulation techniques are now being used
to predict the microstructure at various locations throughout the casting. Since strength and ductility are influenced by
the microstructure, this tool can be used to predict mechanical properties at various locations throughout the casting. This
information can then be used when interpreting FEA results.
Other new developments allow for the prediction of residual
stresses induced during the casting and heat treating process. Most commercially available FEA software does not
consider residual stress. High residual stress can result in
lower than expected component performance and dimensional
capability.
minimizing secondary machining operations. Fig. 3 shows an
example of an aluminum bearing cap that was converted from
gravity permanent mold to squeeze casting. Due to the near net
shape capability of the squeeze casting process, all secondary
machine operations were eliminated. The use of precision
cores with minimal draft (less than .5º per side) eliminated
the need for a secondary drilling operation. The flatness
and surface finish requirements were achieved in the ascast condition, eliminating the milling operation.
For applications that require high mechanical stiffness, design engineers must consider both the modulus of elasticity
and section modulus. Modulus of elasticity is a function of the
stiffness of the alloy itself and is fairly similar for most aluminum casting alloys. Section modulus is a function of stiffness
from the casting geometry. Increasing the section modulus
through design can offset issues with a lower modulus of
elasticity. Complex geometries such as ribs, pockets, and
u-shaped sections can be used to increase section modulus. It
is recommended to avoid drastic wall thickness changes and
isolated thick sections. By avoiding localized thick sections and
drastic wall thickness changes, the tendency to form shrink
porosity is greatly reduced. Isolated thick sections can also induce stress concentration points and cause casting defects such
as hot tears and heat sinks.
Material Selection
One important advantage of the squeeze casting process is that
it is can be used with various alloy/ heat treat combinations
Design Recommendations
that can be tailored to meet design requirements. Primary alThe squeeze casting process is capable of producing comloys, such as A356 (AlSi7Mg) are used in the T6 condition for
plex geometries with high dimensional accuracy and repeatapplications that require high strength and ductility such as
ability. This allows designers to create near-net shapes, thus
control arms, steering knuckles, and suspension links. Secondary alloys such as
ADC12 (AlSi11Cu3Fe) are used in the
Alloy-Temper Yield (MPa) Tensile (MPa) % Elongation Hardness (HBN)
as-cast, T5, and T6 conditions for apA356- T6
220-260
9-15
85-100
290-340
plications that require high strength,
ADC12-F
140-170
2-3.5
95-105
200-270
pressure tightness, and wear resistance. Typical mechanical properties
ADC12- T5
230-260
1-3
110-130
280-320
are shown in Tab. 1.
ADC12- T6
290-320
2-5
120-140
344-380
s
Tab. 1
indici di prestazione dei materiali candidati.
Materials indexes for candidate materials.
P2000TM APPLICATIONS
Fig. 4 shows an example of a squeeze
cast front steering knuckle. In this
3
Alluminio e leghe
<< Memorie
ideally suited for this type of component due to the superior
physical and mechanical properties, dimensional capabilities,
and prior success with similar applications. The annual requirement of 400,000 castings is achieved using a two-cavity
die, ADC12 alloy, and a T6 temper.
CONCLUSION
s
Fig. 5
P2000TM rack & pinion housing for a full size
truck application.
Alloggiamento pignone per camion prodotto con la
tecnologia P2000TM.
application, a direct conversion was made from cast iron to
a much lighter-weight, near net-shape aluminium squeeze
casting. Since steering knuckles are considered safety critical
components, rigorous testing is required prior to shipment. Examples of tests include material property measurements, component strength and fatigue testing, dimensional checks, x-ray,
and ultrasonic inspection. The P2000TM squeeze casting process
was able to meet, and in some cases, exceeded all customer requirements and expectations with A356.2 alloy and a T6 temper [3]. This high-volume knuckle (120,000 parts annually) has
been in production for several years.
Fig. 5 shows an example of a squeeze cast rack and pinion
housing for a high-volume full size truck. The integrity of the
cast housing is critical to the overall function of the hydraulic
steering system. Leakage of hydraulic fluid from any pressurized area of the casting can create a drop in hydraulic pressure, thereby creating a potential malfunction of the vehicle
steering system. Through the years rack and pinion housings
have primarily been made via the conventional HPDC process.
For this particular application, the customer required higher
mechanical properties and burst requirements than the HPDC
process could deliver. The P2000TM squeeze casing process was
Squeeze casting is an established shape-casting process that
is capable of producing lightweight, high integrity, automotive components that can be used for structural applications. The squeeze casting process has many advantages
over other casting processes including high mechanical
properties, near-net shape capability, minimal gas and shrink
porosity, and the ability to heat treat.
Even with the many advantages of the squeeze casting
process, desired quality level cannot be guaranteed without
proper design and upfront engineering. Carefully planned
casting geometries and tooling designs can offset issues
with manufacturability and casting performance. Advanced
computer aided engineering software such as solid modeling,
process simulation, and finite element analysis are powerful
tools that can be used to assist with product development,
tooling design, and process engineering. The use of these
tools, combined with the design and casting engineers
knowledge and experience, can result in lighter weight
casting designs that meet or exceed all performance and
cost targets. The use of lightweight castings will assist the
automobile manufacturers in improving fuel economy and reducing vehicle emissions.
REFERENCES
1) S. CORBIT. and R. DASGUPTA, Squeeze cast automotive applications and squeeze cast aluminum alloy properties (199901-0343). SAE International Congress and Exposition (1999).
2) D. APELIAN and M. MAKHLOUF (eds.), High integrity
aluminum die casting: alloys, processes, & melt preparation.
North American Die Cast Association (NADCA), Rosemount,
Illinois (2004).
3) R. DASGUPTA, C. BARNES, P. RADCLIFFE, and P. DODD,
Squeeze casting of aluminum alloy safety critical components
for automotive applications. World Foundry Congress (2006).
ABSTRACT
APPLICAZIONI DI “SQUEEZE CASTING” NEL
SETTORE AUTOMOBILISTICO E CONSIDERAZIONI
PER LA PROGETTAZIONE
Parole chiave: alluminio e leghe, pressocolata, processi
Per la crescente enfasi sulla riduzione del peso nei veicoli, continua ad
aumentare la domanda di componenti automobilistici più leggeri. Lo
“squeeze casting” è un processo che permette di produrre componenti
leggeri e ad alta integrità che possono essere impiegati per applicazioni
strutturali sugli autoveicoli. Negli ultimi anni il processo di “squeeze casting” è stato utilizzato con varie leghe di alluminio per la produzione di
componenti “near-net-shape” che richiedono alta resistenza meccanica,
4
duttilità, tenuta a pressione o alta resistenza all’usura [1]. Il processo di
“squeeze casting” si è dimostrato un processo economico per applicazioni
ad alti volumi di produzione ed offre ai progettisti una alternativa rispetto
ai processi convenzionali, come la colata a gravità in conchiglia (GPM), la
colata in bassa pressione (LPDC), la colata in sabbia di alluminio / ghisa,
e la pressocolata convenzionale (HPDC).
Il presente documento descrive la tecnologia di squeeze casting (P2000TM)
sviluppata dalla Contech; fornisce anche esempi di alti volumi di produzione per componenti di autoveicoli fabbricati presso tale azienda. Il documento presenta anche delle considerazioni relative alla progettazione
dei prodotti, una panoramica delle tecniche di simulazione del processo, il
confronto delle proprietà meccaniche, alcuni studi di casi per componenti
automobilistici specifici.
Alluminio e leghe
Memorie >>
CASTABILITY MEASURES FOR
DIECASTING ALLOYS: FLUIDITY,
HOT TEARING, AND DIE SOLDERING
B. Dewhirst, S. Li, P. Hogan, D. Apelian
Tautologically, castability is a critical requirement in any casting process. Traditionally, castability in sand
and permanent mold applications is thought to depend heavily on fluidity and hot tearing. Given
capital investments in dies, die soldering is a critical parameter to consider for diecasting. We discuss
quantitative and robust methods to insure repeatable metal casting for diecasting applications by investigating these three areas. Weight reduction initiatives call for progressively thinner sections, which in turn
are dependent on reliable fluidity. Quantitative investigation of hot tearing is revealing how stress develops
and yields as alloys solidify, and this has implications on part distortion even when pressure-casting
methodologies preclude hot tearing failures.
Understanding the underlying mechanism of die soldering presents opportunities to develop methods
to avoid costly downtime and extend die life. Through an understanding of castability parameters,
greater control over the diecasting process can be achieved.
KEYWORD: castability, die soldering, fluidity, hot tearing, part distortion, residual stress
INTRODUCTION
Over the years, castability has been addressed through various angles and perspectives. However no matter what has
been accomplished, it is fair to state that at the present there
is not a single method that the community can point to as a
means of defining an alloy’s castability in terms of measurable quantitative parameters. It is critical that means for
controlling the casting process be developed. Without robust
measures, one will not be able to control the casting process.
It is the latter that is the motivating force behind this project.
Hopefully, the investigative techniques being developed in
this research will become standardized so that an accepted lexicon and methodology is practiced throughout the
casting community.
This paper will focus on three parallel lines of research with
applicability to light metals diecasting: fluidity, hot tearing
(as it relates to stresses developing within solidifying metals as a function of chemistry and microstructure), and die soldering. Each of these three areas of research has the potential
to positively benefit the HPDC industry, either directly or as
B. Dewhirst, S. Li, P. Hogan, D. Apelian
Metal Processing Institute - WPI, 100 Institute Road Worcester, MA 01609 USA
Paper presented at the International Conference High Tech
DieCasting, Montichiari, 9-10 April 2008, organised by AIM
an accompanying benefit to research conducted for other
purposes.
Vacuum fluidity testing allows for the evaluation of various alloys and process modifications in a laboratory setting under rapid solidification conditions, but suffers from
a poor reputation and, as a consequence, has principally been
used for qualitative experimentation. Hot tearing, a consequence of stresses developing during feeding until the casting
tears itself apart, is not found in alloys used in HPDC, but the
investigative techniques being applied to understand hot
tearing are providing a window into how these stresses
develop. Die soldering is important because, in improperly
designed castings, soldering can be a significant problem that
can severely inhibit productivity.
FLUIDITY
Fluidity is a material’s ability to flow into and fill a given cavity, as measured by the dimensions of that cavity under specified experimental conditions, and fluidity is heavily dependent on heat flow during solidification.
Investigations into the impact of foundry variables such as
mold coatings, alloying additions, head pressure, and especially superheat have been investigated and correlated with
mechanisms. For sand and permanent mold castings, it is
abundantly clear that increasing solidification range results
in decreasing fluidity (all other factors being equal). Specific investigations are often alloy or metal/mold/coating
specific in scope, but very subtle influences of minor varia-
37
Alluminio e leghe
tions in alloy purity can be detected. There is some question as
to whether these trends transfer over to die casting, and that
question will be the focus of our discussion.
Thanks in large part to the work of Ragone in developing his
vacuum testing apparatus, which Flemings et al. built upon,
fluidity has seen great advances since Ragone’s 1956 doctoral
thesis [1-6]. Over a period of 8 years, Flemings and collaborators produced the fluidity equations and solidification
mechanisms which are at work in linear castings during
standard fluidity tests.
Ragone demonstrated that the influence of viscosity or a change in viscosity on (casting) fluidity was minimal, and while
the equations he presented did include a viscosity term, subsequent formulations correctly dropped it as insignificant as
compared with other sources of experimental error [1].
The fluidity equation from Flemings [3], for metal with some
superheat ΔT and a mold which conducts heat rapidly is given
below as Equations 1 and 2.
(1)
(2)
Where:
Lf
a
k
c’
To
ΔT
ρ’
Vo
H
h
T
Tm
T’
λ
final length, fluidity
channel radius
critical solid concentration
specific heat of liquid metal
ambient environmental temperature (room temperature)
superheat
density of metal
velocity of metal flow
heat of fusion of metal
heat transfer coefficient at mold-metal interface
the time average melt temp in the fluidity test
metal melting temperature
temperature of superheated metal entering flow channel
critical solid concentration required to stop flow in
‘mushy’ alloys
Flemings reports that the critical solid concentration is
between 0.2 and 0.3 fraction solid, and Campbell gives 0.50.6 using slightly different criteria [4,7,8]. This is the fraction solid where, as will be discussed under flow stoppage mechanisms, the flow is choked off. Attempts to tie this choking
off to dendrite coherency by Dahle, as explored by Backerud,
were inconclusive. He did not find an unambiguous impact of dendrite coherency measurements on fluidity [9-11].
The specific fraction solid at which this takes place varies
with alloy composition and solidifying phase morphology.
This critical fraction solid is likely to be higher for die casting
due to the increased pressure involved, but the extent of
increase is likely to depend on alloy-specific morphology
characteristics. Much work on the relevant solid fractions
where flow is possible has been carried out in the area
of SSM, both in terms of alloy rheology and thermodynamics, and this may have much to contribute in understanding how this factor changes according to the specific casting
and alloy conditions [12].
Past work in the field has focused on maximizing fluidity,
however we believe that decreasing the variations in fluidity is as important as determining under which conditions
fluidity is maximized. There are two main aspects to variation
38
<< Memorie
in fluidity:
→ One is the standard deviation of test methods used in
the lab to determine fluidity.
→ The other is the range over which fluidity values will
vary in a real casting environment where alloy chemistry,
temperature controls, etc. vary within some range.
Given the high part numbers involved in die casting,
questions of repeatability are especially important. Thin
sections are desirable for a variety of reasons, and can be
achieved with increased mean fluidity, but if that increase is
coming at the expense of increased fluidity variation, this will
have the undesirable effect of increasing scrap rates.
Often, the factors which can be adjusted to improve fluidity
have other impacts on the casting process, and so a careful tradeoff must be achieved between insuring there is
enough fluidity (and a margin of safety) without causing deleterious side-effects. Greater fluidity is often achieved by increasing melt superheat, but as will be discussed below, this has
negative implications for die soldering. Mold coatings can
decrease the heat transfer coefficient, and thus increase fluidity, but this may have a small negative impact on cycle time.
While minor alloy additions often have little impact on
fluidity, the secondary alloy components (specifically, their
heat of fusion and morphology) do contribute to fluidity.
Our work to improve the laboratory testing of vacuum fluidity
measurements is largely focused on improving the repeatability of measurements by controlling the various experimental parameters. After a controlled volume of melt is collected, a
thermocouple is inserted into it. When the metal cools to a preset temperature, it is elevated such that the end of a borosilicate tube is immersed in the melt, and vacuum is applied.
The measurement of that length is then made before the
pyrex tube is removed from the experimental setup, as the
rapid fracturing of the glass and other factors otherwise make
it difficult to determine the ‘zero point.’ Through repeated
measurements under controlled experimental conditions we
are establishing the reliability of the test.
A continuing trend in all of engineering, including metal
casting, is the application of modeling software to problems
of interest. These codes, in the case of casting intended to predict filling, hot spots, etc. are no more reliable than the data
upon which they are built. It is hoped that increased precision
of fluidity testing will have a positive impact on these modeling codes by allowing direct comparison of simple geometries in both simulation and the laboratory. Since these
codes do not include direct fluidity calculations, accurate
experimental tests of fluidity would seem to be a good independent check.
HOT TEARING AND INTERNAL STRAIN
Though hot tearing is a casting phenomenon that occurs in
sand castings and processes where the solidification rate is
slower than in die-castings, the mechanism of stress distribution during solidification is appropriate for discussion in
high integrity castings.
This is true more so now than ever now that we can measure and quantify stresses during solidification. Material
behavior during solidification is what matters.
Campbell [7] defines a hot tear as a uniaxial tensile failure,
which results in cracks on the surface or inside the casting.
Alloys having a wide freezing range have a higher tendency
to hot tear. Variables that influence hot tearing include alloy
composition and processing variables [13,14].
Alluminio e leghe
Memorie >>
s
Fig. 1
Cast Iron Mold designed to detect the onset of
the hot tearing. Commercial cast alloy 713 and 518
were evaluated; the former is known to be sensitive to
hot tearing, and the latter has good resistance to hot
tearing. The pouring temperature was set at 60˚C
above the melting point of the alloy during this effort.
The mold temperature was maintained around 200˚C.
Stampo in ghisa progettato per rilevare l’insorgenza di
cricche a caldo. Sono state valutate le leghe commerciali
713 e 518; la prima è risaputa essere sensibile alla
criccatura a caldo, mentre la seconda ha una buona
resistenza verso tale fenomeno. In questa prova la colata
è stata eseguita a una temperatura di 60˚C al di sopra
del punto di fusione della lega.La temperatura dello
stampo è stata mantenuta intorno ai 200˚C.
Hot tearing susceptibility of alloys is greatly influenced
by solidification behavior of molten metal in the mushy
zone. Solidification can be divided into four stages [15]: (i)
Mass feeding where the liquid and solid are free to move; (ii)
Interdendrtic feeding when the dendrites begin to contact
each other, and a coherent solid network forms; (iii)
Interdendritic separation. With increasing fraction solid, the
liquid network becomes fragmented. If liquid feeding is not
adequate, a cavity may form. As thermal contraction occurs,
strains are developed and if the strain imposed on the network
is greater than a critical value, a hot tear will form and
propagate. Lastly, in stage (iv), Interdendritic bridging or
solid feeding occurs. Simply stated, hot tearing occurs if the
solidification shrinkage and thermal deformation of the solid
cannot be compensated by liquid flow.
Measuring the development of strains and the evolution
of hot tearing during solidification is not trivial. The
Metal Processing Institute is a member of the Light Metals Alliance, and we have teamed up with our alliance partner
CANMET to address hot tearing in aluminum alloys. The
constrained bar mold used in this study was developed at
CANMET Materials Technology Laboratory (MTL) and designed to measure load and temperature during solidification. Fig. 1 shows one of the mold plates and testing setup.
The mold is made of cast iron and coated with insulating mold
wash. The test piece has two arms. One test arm (12.5mm) is
constrained at one end with heavy section (22.5mm) to keep
the bar from contraction, so the tension will be developed
and hence cracking could be induced during solidification. The
other arm is for load and temperature measurement with
one end connected to a load cell. This opened end of the
mold is closed with a graphite cylinder block which can move
freely in horizontal direction. The block is connected to the solidifying material on inner side with a screw and on external
s
Fig. 2
(a) Temperature-load-time curves of alloy 713;
(b) Derivative of Load vs. time curve.
(a) Curve temperatura-carico-tempo per la lega 713;
(b) andamento della derivata del carico vs. tempo.
side with a load cell. Two K-type thermocouples are used for
the temperature measurement. One is positioned at the riser
end and the other at the end of the bar as shown in Fig. 1.
After pouring the melt into the mold, the temperature and
load were recorded with a computer data acquisition system.
Fig. 2 and 3 show the measured temperatures and load recorded during casting as a function of time for alloy 713 and 518
respectively. The load represents the tension force developed
in the casting during solidification. The cooling curve T1 was
recorded with thermocouple tip positioned at the riser end
and T2 with thermocouple tip at the end of the bar as shown in
Fig. 1. A rapid rise in temperature (both curves) was observed
immediately after pouring and the temperature started falling
shortly. It’s noticed that negative loads (compressive forces)
were developed shortly after pouring for the tests, probably
due to the pressure head of the melt [16]. When the rod begins
to solidify but cannot contract freely, the tension force increases. Fig. 2(b) and 3(b) are derivatives of load vs. time curve to
determine onset of hot tearing. An obvious change in the rate
suggests that cracking might occur there.
From Fig. 2b, load began developing at proximately 9 seconds
and the solidification temperature was around 617˚C (Fig.2a),
then increased rapidly. It is shown that the rate changed
abruptly to zero at 16.5 seconds, suggesting a severe tear occurred there.
Hot tearing occurred at around 530˚C, corresponding to 94%
solid, according to Pandat Scheil solidification calculation.
The technique developed to measure hot tearing tendency is
a valuable tool to differentiate between alloys and to use it to
optimize alloys for high integrity castings.
39
Alluminio e leghe
<< Memorie
s
Fig. 4
Main effects plot of the effect various alloying
elements on die soldering. Iron, Manganese and Titanium show strong positive effects on reducing soldering, while Nickel promotes soldering [17].
Mappatura dei principali effetti dei vari elementi di lega
sull’incollaggio allo stampo. Ferro, manganese e titanio
mostrano forti effetti positivi sulla riduzione della adesione, mentre il nichel promuove tale fenomeno [17].
s
Fig. 3
(a) Temperature-load-time curves of alloy 518;
(b) Derivative of Load vs. time curve.
(a) Curve temperatura-carico-tempo per la lega 518;
(b) andamento della derivata del carico vs. tempo.
Fig. 3 shows the temperature-load-time curves of alloy 518.
The load started to develop at 10 seconds, and then increased smoothly with time. No abrupt change of rate was
observed, suggesting no crack would occur during solidification. The difference between the load curves of alloy 713
and 518 reveals different hot tearing susceptibility between the
two alloys.
DIE SOLDERING
Die soldering occurs when the cast aluminum alloy comes into
contact with die steel.
Due to the natural affinity of iron and aluminum, a reaction
occurs at the surface which results in the formation of intermetallic phases. Over a series of shots, a significant amount
of aluminum becomes stuck to these phases at the die surface,
and the resulting cast part can begin to miss critical tolerances
or to lose integrity. At this point, the die must be shut down
and cleaned, which is an expensive process when it occurs too
frequently. It is estimated that 1 to 1.5% of variable overhead is
directly attributed to die soldering in casting plants.
With such a large economic effect on the casting process, it is
clear why die soldering needs to be controlled. There are several ways in which this can be achieved. These can be broken down into three groups, which will be discussed further
below: melt chemistry, process conditions and the die surface
condition.
The chemical composition of an alloy can have a dramatic effect on soldering behavior.
The importance of alloy chemistry was shown at WPI’s Me-
40
tals Processing Institute by Sumanth Shankar [17]. In his experiments, he dipped H13 steel pins in 380 alloy and rotated
them to simulate the drag force experienced at the surface of
the die during injection of the metal. After dipping, the thickness of the intermetallic layers that had formed on each sample
was analyzed as a measure of soldering tendency. His results
showed that small additions of Sr and Ti (0.004% and 0.125%,
respectively) had a much greater effect on soldering tendency
than the time of dipping (30 to 75 seconds) or the temperature
of the melt (1150 to1250F).
To further expand on this discovery, Shankar performed another set of experiments to test the effects of a much wider range
of alloying elements. The main effects are shown in Figure .
Not surprisingly, iron had the greatest effect of any alloying
element in the study on reducing die soldering. Iron has long
been added to die casting alloys in order to reduce the die soldering tendency of alloys. It is well known that alloys with insufficient iron content (<0.8-0.9%) will solder readily to the die
under the right conditions. A look at the phase diagram in Fig.
5 shows that the solubility of iron in aluminum with 10% silicon at typical casting temperatures is quite low, around 2-3%.
At temperatures where the melt is likely to be in contact with
the die, this solubility drops even lower. Therefore, even at low
concentrations the presence of iron in the melt reduces the chemical potential gradient of iron from the steel to the melt significantly and slows the reactions that occur at the surface.
Of the other alloying elements, strontium also has the potential
to help control die soldering, in addition to its common use as
a eutectic modifier. In industrial trials a small strontium addition was shown to reduce die soldering by more than 20%. The
effect is not apparent in the main effects plot above because
both of the levels selected were at or above the critical concentration.
The mechanism behind this reduction has to do with the effect strontium has on the viscosity and surface tension of the
alloy. As Fig. 6 shows, the addition of strontium changes the
apparent viscosity and subsequently the surface energy of the
alloy. This causes a reduction in the ability of the alloy to wet
Alluminio e leghe
Memorie >>
s
Fig. 5
Phase diagram of Aluminum-10% Silicon and
low solubility of Fe.
Diagramma di stato dell’ alluminio-10% silicio con bassa
solubilità del ferro.
the die surface and reduces the contact area and the reaction
between the two.
High temperatures and high melt velocity are process conditions which lead to soldering.
Of the two, high temperature is the most important to avoid in
order to prevent soldering.
This can most effectively be done through careful design of
the die. By configuring the part and optimizing the design of
the die cooling system, the potential for soldering can be greatly reduced. It is very important to consider this during the
design phase of a die because once a die is manufactured it
is very difficult to reduce any hot spots. Other potential solutions include using additional spray in the high solder areas to
reduce temperature or the use of inserts with high conduction
coefficients.
Impingement velocity is important to control as well. The die
surface should be coated with lubricants and is likely oxidized
from prior treatment. A high impingement velocity can wash
these protective coatings off of the die surface, exposing the
die steel to the aluminum alloy and begin erosion of the die
surface. Both of these effects will promote the beginning of die
soldering.
SSM processing can help to reduce both the temperature and
velocities apparent in the casting system, and should help reduce die soldering [12].
Die coatings can be useful as a diffusion barrier between the
steel in the die and the aluminum in the cast alloy. An effective coating must be able to withstand the harsh conditions at
the surface of the die, however. Coatings which are sometimes
used include CrN+W, CrN, (TiAl)N and CrC [19]. Additionally, surface treatments such as nitriding and nitro-carburizing
can help to strengthen the surface and prevent erosion, which
accelerates the soldering process by roughening the surface
and creating local temperature excursions at the peaks of the
die surface which solder very quickly.
Accurate modeling of the casting process during the design
phase is very important to an effective control against die soldering. All of the previously mentioned controls require ad-
s
Fig. 6
Change in viscosity of an Al-Si alloy with the
addition of 230ppm Sr [18].
Variazione della viscosità di una lega Al-Si con l’aggiunta
di 230 ppm di stronzio. [18].
ditional cost during the design and manufacturing of the die,
and it must be understood how badly soldering will affect the
process before the costs of any of those controls can be justified.
CONCLUSIONS
Though these three alloy characteristics seem tangentially
related, they are factors that influence castability. In order to
control these castability indices, it is necessary to develop experimental methods until robust quantitative analysis is possible. Once quantitative data can be extracted, the improvement
in our understanding will occur. In the case of die soldering,
multiple possible avenues to reduce the problem have been
identified. Even when the initial intention was to resolve problems occurring in sand and permanent mold castings, such as
hot tearing, the information gleaned about how stresses develop in liquid metal has wider applicability. Though die casting
usually assures good fluidity through the use of pressure, if
fluidity (and the factors which influence its variation) are well
understood, it is possible to operate within tighter processing
windows.
REFERENCES
1] D.V. RAGONE, C.M. ADAMS, H.F. TAYLOR, AFS Trans. 64,
(1956), p.640.
2] D.V. RAGONE, C.M. ADAMS, H.F. TAYLOR, AFS Trans. 64,
(1956), p.653.
3] M.C. FLEMINGS, Brit. Foundryman 57, (1964), p.312.
4) M.C. FLEMINGS, Solidification Processing. McGraw-Hill,
New York (1974).
5] M.C. FLEMINGS, E. NIYAMA, H.F. TAYLOR, AFS Trans. 69,
(1961), p.625.
41
Alluminio e leghe
6] J.E. NIESSE, M.C. FLEMINGS, H.F. TAYLOR, AFS Trans. 67,
(1959), p.685.
7] J. CAMPBELL, Castings. Butterworth-Heinemann, Oxford (1993).
8] A.K. DAHLE, L. ARNBERG, Materials Science Forum, 217222, (1996), p.259.
9] A.K. DAHLE, L. ARNBERG, Materials Science Forum, 217222, (1996), p.269.
10] L. BACKENRUD, E. KROL, J. TAMMINEM, Solidification
Characteristics of Aluminum Alloys Volume 1: Wrought Alloys. (1986).
11] L. BACKENRUD, G. CHAI, J. TAMMINEN, Solidification Characteristics of Aluminum Alloys Volume 2: Foundry Alloys. (1986).
12] Science and Technology of Semi-Solid Metal Processing.
<< Memorie
North American Die Casting Association, (2001).
13] G.K. SIGWORTH, AFS Trans. 104, (1996), p.1053.
14] A.S. METZ, M.C. FLEMINGS, AFS Trans. 78, p.453.
15] D.G. ESKIN, K.L. SUYITNO, Progress in Materials Science,
49, (2004).
16] G. CAO, S. KOU, Met. Trans. A. 37A, (2006), p.3647.
17] S. SHANKAR, A Study of the Interface Reaction Mechanism Between Molten Aluminum and Ferrous Die Materials,
Ph.D. Worcester Polytechnic Institute, (2000).
18] S. SHANKAR, M.M. MAKHLOUF, Internal ACRC Report,
May 2005.
19] J. Wallace, A Guide to Correcting Soldering. North American Die Casting Association, (2006).
ABSTRACT
MISURE DI COLABILITÀ PER LEGHE DA
PRESSOCOLATA: FLUIDITÀ, CRICCABILITÀ A CALDO
E INCOLLAGGIO AGLI STAMPI
Parole chiave: alluminio e leghe, pressocolata
Tautologicamente si può affermare che la colabilità rappresenta un requisito fondamentale in ogni processo di pressocolata. Tradizionalmente, si
pensa che sia la colabilità in sabbia sia le applicazioni in stampi permanenti dipendano dalla fluidità e della criccabilità a caldo. Se si considerano poi
gli investimenti in capitale per gli stampi, anche i fenomeni di incollaggio
diventano un aspetto critico da non sottovalutare. Nella presente memoria si discutono metodi quantitativi ed efficaci per garantire la ripetibilità
42
delle colate di metallo per applicazioni in pressocolata, analizzando i tre
aspetti. La necessità di ridurre il peso richiede di conseguenza l’impiego
di sezioni sempre più sottili, che a loro volta dipendono da livelli affidabili di fluidità. L’analisi quantitativa della criccabilità a caldo permette di
comprendere come si sviluppano e si manifestano le tensioni quando la
lega solidifica, con implicazioni sulle distorsioni del pezzo, anche quando
le tecniche di pressocolata sono in grado di evitare rotture da criccatura
a caldo.
La comprensione del meccanismo alla base dell’incollaggio della lega agli
stampi permette di sviluppare metodi per evitare costosi tempi morti e di
prolungare la vita degli stampi stessi. Attraverso la comprensione dei
tre parametri di colabilità è possibile ottenere un maggiore controllo sul
processo di pressocolata.
Magnesio e leghe
Memorie >>
A PROBABILISTIC APPROACH FOR
MODELLING OF FRACTURE IN
MAGNESIUM DIE-CASTINGS
C. Dørum, D. Dispinar, O.S. Hopperstad, T. Berstad
Quasi-static tensile tests with specimens cut from a generic cast component are performed to characterise the
mechanical behaviour of the high-pressure die cast magnesium alloy AM60. The experimental data is used
to establish a probabilistic methodology for finite element modelling of thin-walled die castings subjected to
quasi-static loading. The cast magnesium alloy AM60 is described using an elastic-plastic constitutive model
consisting of a high-exponent, isotropic yield criterion, the associated flow law and an isotropic hardening rule.
A novel probabilistic approach for modelling of fracture in thin-walled magnesium die-castings using finite
element analysis is developed. The Cockcroft-Latham criterion for ductile fracture is adopted with the fracture
parameter assumed to follow a modified weakest link Weibull distribution. Comparison between the experimental and predicted behaviour of the cast magnesium tensile specimens gives very promising results.
KEYWORDS: magnesium, die-castings, ductile fracture, weibull distribution, finite element analysis
INTRODUCTION
With increased focus on environmental issues, structural designers in the transport industry are forced to search for light-weight
solutions. New materials are considered for vehicle design if
they provide benefits at an affordable cost. The cold-chamber
high pressure die casting method is an important production
method for aluminium and magnesium castings; particular suitable for fully automatic, high productivity, high volume production of complex near net shape parts.A major challenge with this
production method is to optimise the process parameters with
respect to the part design and the solidification characteristics
of the alloy in order to obtain a sound casting without casting
defects. Unbalanced filling and lack of thermal control can cause
bifilms, porosity and surface defects due to turbulence and solidification shrinkage. Consequently, the fracture behaviour of
cast components can be of stochastic character.
To be efficient in the development of new products it is necessary to use finite element (FE) analysis to ensure a structural design that exploits the material. In order to be able to obtain a reliC. Dørum
SINTEF Materials and Chemistry, No-0314 Oslo, Norway
Structural Impact Laboratory (SIMLab), Centre for Research-based
Innovation, No-7491 Trondheim, Norway
D. Dispinar
SINTEF Materials and Chemistry, No-7465 Trondheim, Norway
O.S. Hopperstad
T. Berstad
SINTEF Materials and Chemistry, No-7465 Trondheim, Norway
able prediction of the structural behaviour using such analyses,
an accurate description of the material behaviour is essential.
Hence, a reliable failure criterion is also required, that enables
the designer to exploit the potential of the cast material. This
work presents a new probabilistic approach for finite-element
modelling of the structural behaviour of thin-walled cast magnesium components.
Fig. 1 shows the geometry of the generic AM60 component in-
s
Fig. 1
Illustration of generic cast component: Length = 400 mm, thickness = 2.5 mm, width = 80 mm,
height = 40 mm.
Illustrazione di un generico componente pressocolato:
Lunghezza = 400 mm, spessore = 2.5 mm, larghezza =
80 mm, altezza = 40 mm.
51
Magnesio e leghe
s
Fig. 2
Geometry of tensile test specimen.
Geometria del provino per la prova di trazione.
vestigated in this study, together with the corresponding gating system. The length of component is 400 mm and the wall
thickness is approximately 2.5 mm. In previous studies [1, 2],
the AM60 material was characterized using uniaxial tensile
tests, uniaxial compression tests, and plate bending tests. The
results from the uniaxial tensile tests showed that the scatter
in elongation at fracture is quite large. The poorest area is the
outlet side, where values of effective plastic strain at fracture
as low as 2-3% were measured. The best areas were found to be
the 80 mm flange in front of the gates, where values of effective
plastic strain at fracture as high as 22% were measured [2].
TENSILE TESTS
The tensile tests were carried out in a hydraulic testing machine under displacement control. Force and displacement/
strain were continuously measured. The displacement rate
was adjusted to obtain a strain rate approximately equal to
2×10-3 s-1. All tests were carried out at ambient temperature.
Uniaxial tension specimens were cut from the inlet wall and
the outlet wall in the longitudinal direction, and from the 80
mm web of the casting in both the longitudinal and transverse
direction. The geometry of the tensile specimens is shown in
Fig. 2. The strain in the length direction was measured by an
extensometer with 25 mm gauge length. Cauchy stress versus
logarithmic plastic strain curves are provided in Fig 3. for different parts of the component.
The figure shows the strong variation of the strain to fracture
with position in the casting and between duplicate tests. It is
further observed that there is also variation in the flow stress
level between the various tests, but this variation is less significant.
METALLURGICAL CONSIDERATIONS
The mechanical properties of an alloy depend on the defects
that may be present in the matrix. These defects could be point,
line, surface or volume defects. Among these defects, volume
defects (porosity, secondary phases or inclusions) are known
to be the most significant ones and may affect the mechanical
properties dramatically.
Inclusions, basically oxides, play an important role in casting
operations. Particularly in high pressure die casting operations,
with casting speeds of minimum 15 m/s up to 40 m/s, the liquid metal advances into the mould in jets that introduces the
surface oxide to become incorporated into the melt. However,
the oxide inclusions can not exist in melts as a single, because
the only way they can become incorporated into the liquid is
by entrainment action [3]. During such a simple folding action,
the two non-wetted oxide surfaces come in contact to form a
bifilm that acts as a crack in the casting. Therefore, in high
pressure die castings, the casting will have a spatial distribution of casting defects. The size and population of these defects
are critical since they act as the initiation points for porosity
nucleation and also as stress risers. As seen from Fig. , pictures
using scanning electron microscope (SEM) show that the fracture surface has a high density of crack-like pores. By closer
examination, it can be confirmed that the crack-like pores are
all oxides.
It is well known that in the presence of defects or stress risers,
the components may fracture at stresses far away from their
nominal theoretical limits. Fig. 3 is a perfect example to such
phenomena. Even within the groups (longitudinal and transverse direction, outlet and inlet side) there is a huge scatter
of elongation and maximum stress values. It is important to
note that even the proof stress changes considerably. Here, the
oxides (or bifilms) act as a way of strengthening mechanism
in the matrix which is very similar to the behaviour of metal
matrix composites. It is also interesting to note that there is one
sample in Fig. 3 that fractures even before reaching the proof
stress. This pre-mature fracture is another example of the presence of defects (most probably bifilms) in the casting. The flow
lines observed on the surface of the samples (see Fig. 5) are the
proof of unstable flow of the liquid magnesium in the mould
cavity. It is so interesting to observe bifilms (folded oxides) in
such sizes on the outer surface of casting part.
MATERIAL MODELLING
s
Fig. 3
Cauchy stress vs. logarithmic plastic strain
curves for AM60.
Tensione Cauchy vs. curve logaritmiche di deformazione
plastica per l’AM60.
52
<< Memorie
The cast magnesium alloy AM60 is modelled using an elastoplastic constitutive model including a high-exponent, isotropic
yield criterion, the associated flow law and isotropic hardening. Fracture is modelled by element erosion when a fracture
Magnesio e leghe
Memorie >>
s
Fig. 5
s
Fig. 4
SEM images from the fracture surface of
samples showing the crumpled oxides.
Immagini SEM della superficie di frattura dei campioni
che mostrano gli ossidi.
Picture of test specimens with flow lines on the
surface.
Immagine dei campioni di prova con linee di scorrimento
sulla superficie.
criterion is reached. The model has been implemented in explicit finite element code LS-DYNA [4].
The high-exponent isotropic yield criterion [5, 6] is written in
the form
bution [10] of the local strength of cohesive elements. Thus, the
probability of introducing a weak cohesive element increases
with the cohesive element size. Inspired by this idea, the fracture parameter of a finite element is assumed to follow a modified weakest-link Weibull distribution in the current study.
The Weibull distribution gives the fracture probability P (σ) of
a material volume to under effective tensile loading, i.e.
(1)
where σ1 and σ2 are principal stresses in plane stress and k is
a material parameter. The flow stress σy is defined by the isotropic hardening rule
(2)
where εe is the effective plastic strain, σ0 is the proportionality
limit, and Qi and Ci are hardening parameters. Using a least
squares method, the hardening parameters were determined
from the Cauchy stress versus logarithmic plastic strain curves
in Fig. . Any variation in flow stress with position in the casting
was not accounted for in the FE simulations, and thus a mean
hardening curve was applied.
In the present model, a criterion of ductile fracture proposed
by Cockcroft and Latham [7] is added. The fracture criterion is
coupled with the element-erosion algorithm available in LSDYNA [4]. As the fracture criterion is reached in an element,
this element is removed (eroded) from the finite element model. The fracture criterion can be expressed as
(3)
where σ1 is the maximum principal stress and Wc is the critical
value of the integral W. Hence, fracture occurs when W = Wc.
Henceforth, Wc will be referred to as the fracture parameter,
while W will be denoted the Cockcroft-Latham integral. It is
seen that fracture cannot occur when the maximum principal
stress is compressive and that neither stresses nor strains alone
are sufficient to cause fracture. Furthermore, the fracture strain
increases with decreasing stress triaxiality (in the shear tests,
the stress triaxiality is significantly reduced compared to the
uniaxial tension test).
The uniaxial tensile test specimens failed before the point of diffuse necking for the AM60 alloy, and, accordingly, the stress and
strain field are uniform up to fracture. Hence, the fracture parameter is obtained as the area under the work-hardening curve.
Zhou and Molinari [8, 9] propose a micro-cracking model for
brittle materials (ceramics) considering the stochastic distribution of internal defects. The model introduces a Weibull distri-
(4)
where V is the volume, V0 is the scaling volume, σ0 is the scaling stress, and m is the Weibull modulus. Since cast magnesium is not a brittle material, the use of a critical fracture stress
is not justified. Instead, the Cockcroft-Latham ductile fracture
criterion is adopted, and the fracture probability of a material
volume is recast as
(5)
where Wc0 is the scaling value of the fracture parameter. By
using a random number generator and inverse sampling, this
Weibull distribution of fracture parameters can then be assigned to the integration points in the FE mesh. With this approach, a small element in the FE model will most probably be
given more ductile material properties than a larger element.
Fig 6. compares the numerical predictions with the experimental results from the tensile tests on cast magnesium AM60.
Here, the uniaxial tension test specimens were modelled by
720 shell elements (i.e., a characteristic element size equal to
1.0 mm) It is seen that the observed experimental scatter is well
reproduced numerically.
CONCLUDING REMARKS
The quasi-static behaviour of high-pressure die cast magnesium
alloy AM60 has been studied through tensile tests. The specimens were taken from various positions in the cast profile. The
experimental data were used to develop a probabilistic method
for finite element modelling of thin-walled die castings subjected
to quasi-static loading. The ductility of the specimens cut from
the castings depends on the position in the casting. There are also
significant variations in ductility when comparing the measured
characteristics of specimens cut from different castings that were
cast under equal casting conditions. Thus, as a result of unstable
flow of the liquid magnesium in the mould cavity, the mechanical properties of the casting are of stochastic nature. By combin-
53
Magnesio e leghe
<< Memorie
to establish a deterministic-stochastic approach that can model
both the variations in ductility depending on the material’s position in the casting as well the stochastic aspects.
ACKNOWLEDGEMENTS
The authors would like to acknowledge the support through the
EU project NADIA, and Hydro Aluminium and the Research
Council of Norway for their support through SIMLab, the Centre
for Research-based Innovation.
REFERENCES
s
Fig. 6
Experimental and numerical engineering stress
vs. engineering strain data.
Valori di tensione sperimentali e calcolati numerica vs.
valori di deformazione.
ing the Cockcroft-Latham fracture criterion and the Weibull statistical distribution function, the fracture parameter was defined
as a stochastic Weibull distributed parameter. Repeated finite
element simulations of the tensile tests were carried out, giving
predictions very similar to the experimental behaviour. Accurate
numerical prediction of the mechanical capacity (especially in
terms of ductility) of castings requires that the inhomogeneous
distribution of defects is included. A coupling of die-casting process simulations and the current approach should be investigated
1] Dørum, C., Hopperstad, O.S., Lademo, O.-G., Langseth, M..
Aluminium and magnesium castings – experimental work and
numerical analyses, Int. J. Crash. 8 (2003) 455-470.
2] Dørum, C., Hopperstad, O.S., Lademo, O.-G., Langseth, M.
Numerical modelling of the structural behaviour of thin-walled
cast magnesium components using a through-process approach.
Materials and Design 28 (2007) 2619-2631.
3] Campbell, J., Castings, 2nd Ed., Oxford, UK, ButterworthHeinemann, 2003.
4] Hallquist, J.O. LS-DYNA, v.970, Livermore Software Technology Corp., 2003.
5] Hersey, A.V., Dahlgren, V.A. The plasticity of an isotropic aggregate of anisotropic face-centered cubic crystals. J. Appl. Mech.
Phys. Solids 76 (1954) 241-249.
6] Hosford, W.F. A generalized isotropic yield criterion. J. Appl.
Mech. 39 (1972) 607-609.
7] Cockcroft, M.G., Latham, D.J. Ductility and the workability of
metals. J. Inst. Metals 96 (1968) 33-39.
8] Zhou, F., Molinari, J.-F. Stochastic fracture of ceramics under
dynamic tensile loading. Int. J. Solids and Struct. 41 (2004) 65736596.
9] Zhou, F., Molinari, J.-F. Dynamic crack propagation with cohesive elements: a methodology to address mesh dependency. Int.
J. Numer. Meth. Engng. 59 (2004) 1-24.
10] Weibull, W. A statistical distribution function of wide applicability. J. Appl. Mech. 18 (1951) 293-297.
ABSTRACT
APPROCCIO PROBABILISTICO PER LA
MODELLIZZAZIONE DELLA FRATTURA NEI
PRODOTTI PRESSOCOLATI IN MAGNESIO
Parole chiave: magnesio e leghe, pressocolata,
caratterizzazione, frattura
Nel presente lavoro sono state effettuate prove di trazione quasi-statiche su
campioni ricavati da un componente generico pressocolato al fine di caratterizzare il comportamento meccanico della lega di magnesio AM60 pressofusa ad alta pressione. I dati sperimentali sono stati utilizzati per stabilire
una metodologia probabilistica per la modellazione agli elementi finiti di
54
un pezzo a pareti sottili pressocolato, sottoposto a carico quasi-statico. La
lega di magnesio AM60 pressocolata è stata descritta mediante un modello
costitutivo elasto-plastico costituito da un criterio ad alto esponente di resa
isotropa; dalla legge di flusso associata da un regime di incrudimentro
isotropo. Mediante l’analisi agli elementi finiti è stato sviluppato un nuovo
approccio probabilistico per la modellizzazione della frattura di pressocolati
in magnesio a parete sottile mediante l’analisi agli elementi finiti. Per la
frattura duttile è adottato il criterio di Cockcroft-Latham assumendo che
il parametro di frattura segua una distribuzione di Weibull modificata ad
anello più debole. Per i campioni di magnesio pressocolato il confronto tra il
comportamento tensile sperimentale e quello previsto per i campioni tensili
di magnesio pressocolato ha fornito risultati molto promettenti.
Memorie >>
Acciaio inossidabile
THE EFFECT OF AUSTENITE VOLUME
FRACTION ON THE DEFORMATION
RESISTANCE OF 409 STAINLESS
STEELS DURING HOT-STRIP ROLLING
D. Chae, S. Lee, S. Son
The mill log data obtained from the hot-strip rolling of 409 stainless steels were analyzed in order to investigate the effect of the chemical composition on the deformation resistance. The results showed that the deformation
resistance depended sensitively on the austenite stabilizing capability of the material chemistry, suggesting
the austenite volume fraction as a dominant factor in controlling the deformation resistance. Deformation
resistance ratio (DRR) was defined as a ratio of the deformation resistance of a two-phase (ferritic+austenitic)
microstructure to that of a fully ferritic microstructure. The dependence of DRR on the austenite volume
fraction appeared to be linear, which was also observed by the plane strain compression tests performed on the
laboratory specimens with various austenite volume fractions. The implication of this result is that during the
hot-strip rolling of 409 stainless steels with a two-phase microstructure, these steels are likely to deform in an
equal-strain manner.
KEYWORDS: 409 stainless steel, mean flow stress, deformation resistance, two-phase material, austenite potential
and mill log analysis
INTRODUCTION
In hot-strip rolling, precise thickness control requires an accurate prediction of the roll force. The accurate prediction of
the roll force, in turn, depends on the accurate calculation of
the hot flow strength of the material because it significantly
affects the pressure at the interface between the work roll and
the rolled material. In order to calculate the hot flow strength
as a function of rolling parameters which are representative
of each rolling pass, an average flow strength, hereafter called
UNS
S40910
S40920
S40930
S40945
S40975
C
0.030
N
0.030
‘deformation resistance’ is defined over the total applied strain
during a rolling pass[1, 2]. Industrially, deformation resistance
is analyzed using mill log data to develop and refine rolling
mill models.
The 409 stainless steels are characterized by relatively low
carbon and nitrogen contents with approximately 11% chromium in their chemical compositions (Tab. 1). Titanium (and/
or niobium) is usually added enough to tie up carbon and nitrogen atoms. Due to the fact that titanium is a strong ferrite
former, these steels are fully ferritic in an annealed condition.
Composition percentage, max or range
Cr
Ni
Other elements
10.50.5
Ti 6(C+N) min, 0.5 max; Nb 0.17 max
11.7
0.5
Ti 8(C+N) min, Ti 0.15-0.5; Nb 0.10 max
0.5
Ti+Nb 0.08+8(C+N) min, 0.75max; Ti 0.05 min
0.5
Nb 0.18-0.40, Ti 0.05-0.20
0.5-1.0 Ti 6(C+N) min, 0.75 max
s
Tab. 1
Dongchul Chae, Soochan Lee,
Seunglak Son
Posco, Pohang, Korea
Paper presented at the 3rd International Conference Thermomechanical
Processing of Steels, organised by AIM, Padova, 10-12 September 2008
Chemical compositions of ferritic
stainless steel grades containing 11% chromium
in ASTM A 240/A 240M-00.
Composizione chimica dei gradi di acciaio inossidabile ferritico
contenenti 11% cromio negli ASTM A 240/A 240M-00.
la metallurgia italiana >> febbraio 2009
55
ID
A
B
C
D
F
C
0.007
~
0.010
Cr
9.5
~
11.5
Ni
0.1
~
2.0
Ti
0.2
~
0.3
<< Memorie
Ni-equivalent
0.57
0.80
0.93
1.02
2.46
Cr-equivalent
12.64
12.65
12.43
11.98
10.86
s
Tab. 2
Chemical compositions of five plates (wt%). Ni-equivalent and Cr-equivalent are defined as Ni+0.5Mn+30C+0.3Cu+25N
and Cr+2.0Si+1.5Mo+5.5Al+1.5Ti, respectively [10].
Composizione chinica di cinque lastre (wt%). Equivalenti di Ni e equivalenti di Cr sono definiti come Ni+0.5Mn+30C+0.3Cu+25N e
Cr+2.0Si+1.5Mo+5.5Al+1.5Ti, rispettivamente [10].
However, it has been reported that these steels may contain
austenite as a second phase at hot rolling temperatures [3].
Therefore, the present work aims at examining the influence of the presence of austenite on the deformation resistance of 409 stainless steels.
The industrial mill logs have been used to analyze deformation resistance as described elsewhere [4-7]. Most of the previous studies have used the mill log analysis on the carbon
steels, where the strain accumulation and recrystallization
between the rolling stands are important microstructural
phenomena to be analyzed for controlling the microstructure. However, there has been no study on the effect of the
presence of austenite on the deformation resistance of 409
stainless steels.
The prediction of the hot flow strength of the two-phase
material is difficult because the partitioning of stress and
strain is influenced by many factors [8]. The principal intrinsic and extrinsic factors which can affect the hot flow
strength of the two-phase material are the phase morphology and the mode of deformation, respectively. The plane
strain compression results on the wrought duplex stainless
steels by Iza-Mendia et. al. [9] implies the equal-strain condition as an useful assumption, where the hot flow strength of the two-phase material is simply calculated from the
strength of the each constitutive phase and its volume fraction [8-9]. In the current study, an attempt has also been
made to check if the equal-strain criterion is a reasonable
assumption in predicting the deformation resistance of
409 stainless steels, where the cumulative alloy contents
are much smaller than the duplex stainless steels mentioned above. For this purpose, specimens were cast and hotrolled in a laboratory to produce various austenite volume fractions and the plane strain compression tests were
performed at a high temperature. Of particular interest is
whether the equal strain condition is supported by the mill
log analyses.
1050 °C for 1 hour and subsequently hot-rolled into 12mm
thick plates.
The plane strain compression specimens were machined
from the 12mm thick plates such that the applied compression loading axis corresponded to the thickness direction
of the plate. The plane strain specimen geometry used is
shown in Fig. 1. The hot plane strain compression tests were
performed at 1050°C and at a strain rate of 5/sec using a
Gleeble 3800. The relatively large strain (equivalent strain
of 0.7) response of the material was determined. After unloading the plane strain compression specimen, the width expansion was measured with a measuring microscope and
approximately 12% width expansion was observed.
Log data were collected from the POSCO’s seven-stand hot
strip mill, Pohang, Korea. In order to calculate true strains,
strain rates, and deformation resistances, the following
log data were used; strip width, strip entry and exit thicknesses, work roll diameter, work roll rotational speed, roll
force, and strip entry temperature predicted according to a
model for each pass. In addition, the Sims’s equation was
employed to calculate deformation resistances and the roll
flattening was also taken into account [4-7].
EXPERIMENTAL
The primary goal of the current study is to investigate the
effect of austenite volume fraction on the flow strength in a
plane strain compression mode of deformation. This can be
accomplished by preparing the materials with the various
austenite volume fractions. For this purpose, five materials
were cast in a laboratory. Tab. 2 lists the chemical composition of the materials. The 50 kg ingots were initially hot-rolled into 20mm thick plates. These had been heat-treated at
56
s
Fig. 1
Specimen geometry of plane strain compression testing.
Geometria dei campioni della prova di deformazione da
compressione del piano.
febbraio 2009 << la metallurgia italiana
Memorie >>
DISCUSSION
s
Fig. 2
s
The effect of chemical composition on the
austenite volume fraction. The two ingots were quenched
after the heat treatment at 950~1250°C for 1 hour.
Effetto della composizione chimica sulla frazione in volume
del austenite. I due lingotti sono stati temprati dopo
trattamento termico a 950~1250°C per 1 ora.
The Shaffler diagram represents the as-solidified microstructure after a rapid cooling as a function of Cr-equivalent and Ni-equivalent [10]. It is noteworthy that the
compositional range of 409 stainless steels is near the twophase (martensitic+ferritic) boundary in the Shaffler diagram.
Because the presence of martensite implies the thermal
transformation from austenite, it is likely that the material
may have an austenite phase at high temperatures. This
seems to be closely related to one of the significant features of the Fe-Cr equilibrium diagram, that is, the phase
boundary between austenite and ferrite fields, known as
the gamma loop. Based on the Fe-Cr equilibrium diagram,
approximately 11% Cr is near the nose of the gamma loop.
Therefore, austenite is likely to form at high temperatures.
In addition, it is also well known that the role of austenite
stabilizing elements such as carbon and nitrogen is to shift
the gamma loop to higher chromium content. Therefore,
the small fluctuation in the austenite stabilizing capability
of the material chemistry can possibly affect the phase balance between ferrite and austenite significantly.
It can be seen in Fig. 2 that, for two 409 ingots produced
in a laboratory, the high temperature microstructure consists of two phases, austenite+ferrite, and the maximum
amount of austenite occurs at about 1050 °C.
The microstructures of the 12mm thick hot rolled plates
exposed at 1050°C for 5 minutes
are shown in Fig. 3. The microstructure of the material A (in Tab.
2) is observed to be fully ferritic
as shown in Fig. 3(a). The region
of light contrast in Fig. 3(b) corresponds to the elongated martensite transformed from austenite.
As the Ni-equivalent of the materials increases, the banded array
of alternating layers of ferrite and
martensite (i.e., austenite) becomes
distinctive. The consequence of the
highest Ni-equivalent is the formation of a 100% martensitic microstructure (Fig. 3(e)).
Fig. 3
Light micrographs
quenched after the heat
treatment at 1050°C for 5
minutes in the case of (a) the
material A, (b) the material
B, (c) the material C, (d) the
material D and (e) the material
F in Table 2. Murakami etchant
was used for etching.
Micrografie di pezzi temprati
dopo trattamento termico a
1050°C per 5 minuti nel caso del
(a) materiale A, (b) materiale B,
(c) materiale C, (d)
materiale D e (e) materiale F
di Tebella 2. Per l’attacco chimico
è stato utilizzato il reagente di
Murakami.
57
<< Memorie
on the hot flow strength, the mean flow stress at a certain strain is defined from the stress-strain curve obtained
from the plane strain compression tests.
[1]
The mean flow stresses at an equivalent strain of 0.5 were
shown in Fig. 5(a) as a function of austenite volume fraction. The sensitivity of the MFS to the austenite volume
fraction appears to be linear. This linear relationship may
be explained by Equ. 2.
where, fα and fγ are the volume fraction of ferrite and the
volume fraction of austenite, respectively.
s
Fig. 4
Compressive stress-strain responses of the materials with the chemical compositions and the microstructures
shown in Table 2 and Fig. 3, respectively.
Risposta sforzo a compressione-deformazione dei materiali con
composizione chimica e microstruttura mostrati rispettivamente
in Tab 2 e Fig 3.
The plane strain compression tests were performed on the
specimens machined from the materials in Tab. 2 at a strain
rate of 5/s and at 1050°C. Austenite volume fractions were
also measured using an image analysis technique on the
specimens quenched before compression. The measured
values were, 0, 32, 51, 73, and 100% as shown in Fig. 4. As
the amount of the austenite increases, the higher degree of
strain-hardening is observed on a flow curve. The distinctively different strain hardening behaviours results in the
significantly different flow stress levels at large strains.
Thus, these results indicate that the austenite volume fraction plays a major role in increasing the flow stress.
In order to quantify the effect of austenite volume fraction
Deformation resistance ratio (DRR) is defined as a ratio of
the MFS of a two-phase (ferrite+austenite) microstructure
to that of a fully ferritic microstructure by adopting so called “the mixture rule” as indicated in Equ. 2.
[3]
The dependence of DRR on austenite volume fraction is
shown in Fig. 5(b). The deformation of coarse two-phase
microstructures depends on whether the material obeys
an equal-stress condition (in which strain can concentrate in one of the microconstituents) or if an equal-strain
condition applies to the constituents. Given the elongated
morphology of the microstructure as shown in Fig. 3, we
expect that the equal-strain condition to be obeyed for deformation with the plane strain compression axes along
the thickness direction. The linear dependence of DRR on
austenite volume fraction as shown in Fig. 5(b) implies
that the equal-strain condition is closely followed.
The deformation resistance was calculated using the fol-
s
Fig. 5
Dependences of (a) mean flow stress and (b) DRR on austenite volume fraction.
Dependenza di (a) sollecitazione di flusso media e (b) DRR su frazione di volume di austenite.
58
[2]
Memorie >>
Material
M1
M2
M3
M4
C
0.005
~
0.015
Chemical composition (wt%)
N
C+N
Cr-equivalent Ni-equivalent
0.68
0.005
0.010
12.95
0.75
~
~
12.48
0.93
0.010
0.025
12.20
0.48
13.03
Austenite volume
fraction before F1 (%)
~0
7
33
~0
s
Tab. 3
Chemical compositions of the four materials(wt%) and their austenite volume fraction observed before the entry to the
first pass, F1, of the hot–strip finishing mill. Ni-equivalent and Cr-equivalent are defined as Ni+0.5Mn+30C+0.3Cu+25N and
Cr+2.0Si+1.5Mo+5.5Al+1.5Ti, respectively.
Composizione chimica dei quattro materiali (% in peso) e loro frazione di austenite (in volume) rilevata dopo ingresso alla prima passata, F1, nel laminatoio di finitura. Ni e Cr equivalenti definiti rispettivamente come Ni+0.5Mn+30C+0.3Cu+25N e
Cr+2.0Si+1.5Mo+5.5Al+1.5Ti.
lowing equation:
[4]
where, P, B, Ld and Qp are roll
force, strip width, the projected
length of the arc of contact between the roll and the rolled material
and Sims geometrical factor, respectively. It should be reminded
that the MFS obtained from the
stress-strain curve can be compared to the deformation resistance
obtained from the rolling data. Thus, MFS in Equ. 2 were
replaced with Km. Then, the DRR is also defined from the
deformation resistance calculated from the mill log data.
[5]
where, Km,α and Km,γ are the deformation resistances of a
single ferritic microstructure and a single austenitic microstructure, respectively.
The dependence of deformation resistance on the entry
temperature was analyzed from the first rolling pass, F1,
to the last rolling pass, F7. Four materials were examined
in details. The differences in the composition and the microstructure are summarized in Tab. 3. Among the four
materials, it was observed that M2 and M4 had two phase
(ferrite+austenite) microstructures right before the entry to
F1. The austenite volume fractions of M2 and M3 materials
right before the entry to F1 were measured to be 7% and
33% as shown in Fig. 6, respectively while austenite formation was not found from the M1 and M4 materials.
The rolling conditions for the materials, M2, M3 and M4,
were also much similar to those of M1 except for the entry temperatures. It is a usual occurrence that if the rolling
temperature increases, then the deformation resistance decreases. This usual expectation is confirmed from the observation that the deformation resistance of the fully ferritic
M4 is lower at the same pass than that of the fully ferritic
M1 because the rolling temperature of M4 was higher than
that of M1. However, comparing the responses of M1 with
those of M2 and M3, a significant trend is observed (Fig.
s
Fig. 6
Microstructures of (a) the material M2 and
(b) M3. Martensitic phases (i.e., high temperature
austenitic phases) with dark contrast are elongated
along the rolling direction.
Microstrutture di (a) il materiale M2 e (b) materiale
M3. Con contrasto scuro le fasi martensitiche (ex fasi
austenitiche alle alte temperature) con contrasto scuro
sono allungate lungo la direzione di laminazione.
7(a)). The deformation resistances of the two-phase materials, M2 and M3 are higher than that of the fully ferritic
material, M1, even though the rolling temperatures were
higher in the case of M2 and M3.
This unexpected behaviour can be understood by the presence of the higher volume fraction of hard austenite in the
microstructure of the materials, M2 and M3.
An attempt was made to calculate the DRR values during
the first pass, F1, for the materials shown in Tab. 3. In order to calculate the DRR, the dependence of deformation
resistance on the entry temperature must be known for the
fully ferritic material. Therefore, additional mill logs were
analysed. Twenty seven materials with six different chemical compositions were selected for the investigation. The
applied strains were in the range of 0.45~0.50 and the strain
rates were between 7.5 to 9.5/sec. Thus, the rolling conditions except for the temperature were almost constant. The
samples for the microstructural examination were taken
right before the entry to the first pass, F1, and all the microstructures investigated using an optical microscope were
59
<< Memorie
s
Fig. 7
The deformation resistances of (a) the materials (Table 3) during rolling passes and (b) the materials with a single ferritic
microstructure during the first rolling pass, F1.
Resistenza alla deformazione dei (a) materiali (Tab. 3) durante i passaggi di laminazione e (b) i materiali con una singola microstruttura
ferritica durante il primo passaggio di laminazione, F1.
found to be ferritic. The calculated values of the deformation resistances during the first pass, F1, were analyzed together and Arrehenius type of temperature dependence of
deformation resistance was obtained as follows:
[6]
where, A and B are constants [12]. The values of the constants, A and B, are 7.9x10-4kg/mm2 and 11.36x103K, respectively. The Arrehenius type fit was compared with
the calculated deformation resistance data during the first
pass, F1 (Fig. 7(b)).
The DRR was evaluated by inserting Equ. (6) into Equ. (5).
The calculated DRR as a function of austenite volume fraction was shown in Fig. 8(a) for the materials, M1, M2, M3
and M4. These data imply that the DRR values calculated
from the mill log analysis are quite close to those obtained
from the plane strain compression results as shown in Fig.
8(a). This is understandable because the strains and the
strain rates considered were quite similar between the mill
log analyses during the first rolling pass, F1, and the plane
strain compression tests performed in this study. The influence of solute content on the ability to form austenite
at hot rolling temperature was predicted in terms of the
austenite potential based on thermodynamic calculations
using a commercial software package, ThermoCalc. Austenite potential (%) is defined as austenite mass fraction at
1000°C. As shown in Fig. 8(b), The DRR increases as the
austenite potential increases.
Therefore, it implies again that the DRR values depend sensitively on the austenite stabilizing capability of the material chemistry.
Finally, the DRR values during the first pass, F1, calculated from the twenty eight materials with different chemical
compositions were shown as a function of austenite potential (Fig. 9(a)). The rolling conditions such as the strains and
the strain rates were also very similar. It is interesting to
note that the DRR values approximately become one if the
60
austenite potential is equal to or lower than about 50%. At
these low levels of the austenite potential, the optical microstructures examined from the samples taken right before the entry to F1 were fully ferritic. The effect of the chemical composition on DRR was shown in Fig. 9(b). It is clearly
seen that the DRR increases as Ni-equivalent increases and
Cr-equivalent decreases. The implication of Fig. 9(b) is that
small fluctuations in Ni-equivalent and Cr-equivalent can
significantly change the phase balance between ferrite and
austenite at hot-strip rolling temperatures.
CONCLUSION
The deformation resistances of 409 stainless steels during
hot-strip rolling were investigated to understand the influence of the presence of austenite at high temperatures
on the rolling force.
Although the commercially cold-rolled and annealed sheet
products of 409 stainless steels are essentially ferritic in an
as-annealed state, these steels can have a two-phase microstructure (ferrite+austenite) during hot-strip rolling, which
depends on the chemical composition. The hot plane strain
compression tests were performed on the specimens, cast
and hot-rolled in a laboratory, with different austenite volume fractions. The tests clearly showed the significant role
played by the austenite volume fraction in controlling the
strain hardening behaviour of the two-phase microstructure. As a result, the mean flow stress at an equivalent strain
of 0.5 was found to be approximately a linear function of
the austenite volume fraction. The mill data from the hotstrip rolling were analyzed in order to calculate the deformation resistance developed in the roll gap of the rolling
stand. Based on the Sims’s model, the calculated deformation resistances were found to depend sensitively on the
austenite stabilizing capability of the material chemistry.
Accordingly, the influence of the solute content on the ability to form austenite at hot rolling temperatures was predicted in terms of the austenite potential which is defined
as the austenite mass fraction at 1000°C. The deformation
Memorie >>
s
Fig. 8
The dependence of DRR as a function of (a) the measured austenite volume fraction and (b) the calculated
austenite potential. The line fit in Fig. 5(b) was superimposed to compare the DRRs from the experiments with those
from the mill log analyses.
Variazione del DRR in funzione di (a) frazione in volume dell’austenite misurata e (b) potenziale di austenite calcolato. La
linea inserita in Fig. 5(b) è stata aggiunta per confrontare i DRR delle sperimentali con quelli dedotti dall’ analisi dei documenti di produzione.
resistance ratio (DRR) was defined as a ratio of the deformation resistance of a two-phase microstructure to that of a
fully ferritic microstructure. The DRR appears to vary linearly with the austenite volume fraction, thus implying that
the material is likely to deform in an equal-strain manner
along the hot rolling direction.
REFERENCES
1) William L. Roberts, Hot Rolling of Steels, Marcel Dekker
Inc., 1983, p.649
2) Vladimir B. Ginzburg and Robert Ballas, Flat Rolling
Fundamentals, Marcel Dekker Inc., 2000, p.199
3) Robert G. Nooning, Jr., Master Thesis, University of
Pittsburgh, 2002, p76.
4) T. M. Maccagno, J. J. Jonas, S. Yue, B. J. McCrady, R. Slobodian and D. Deeks, ISIJ International, Vol. 34, 1994, No.
11, p.917.
5) T. M. Maccagno, J. J. Jonas and P. D. Hodgson, ISIJ International, Vol. 36, 1996, No. 6, p.720.
6) F. Siciliano Jr., K. Minami, T. M. Maccagno and J. J. Jonas,
ISIJ International, Vol. 36, 1996, No. 12, p.1500.
s
Fig. 9
The dependence of DRR during a first rolling pass, F1, as a function of (a) the calculated austenite potential
and (b) the Cr-equivalent & Ni-equivalent. Three dimensional data have been projected on the two dimensional planes
in Fig. 8(b).
La dipendenza del DRR durante il primo passaggio di laminazione, F1, in funzione di (a) potenziale di austenite calcolato e (b)
Cr e Ni equivalenti. I dati tridimensionali sono stati proiettati sui piani bidimensionali di Fig. 8(b).
61
7) Fulvio Siciliano, Jr. and John J. Jonas, Metallurgical and
Materials Transactions A, Vol. 31A, February (2000), p.511
8) Lode Duprez, Doctoral Thesis, Universiteit Gent, March
(2003), p.97
9) A. Iza-Mendia, A. Pinol-Juez, J. J. Urcola, and I. Gutiér-
<< Memorie
rez, Metallurgical and Materials Transactions A, Vol. 29A,
February (1998), p.2975
10) George Krauss, Steels, ASM International, 2005, p.495
11) Shigeru Shida, Journal of the Japan Society for Technology of Plasticity, Vol. 9, No. 85, 1968-2, p.127
ABSTRACT
L’EFFETTO DELLA FRAZIONE IN VOLUME
DELL’AUSTENITE SULLA RESISTENZA ALLA
DEFORMAZIONE DEGLI ACCIAI INOSSIDABILI 409
DURANTE LA LAMINAZIONE A CALDO DI NASTRI
Parole chiave: acciaio inossidabile, laminazione a caldo
Nel presente lavoro sono stati analizzati i dati raccolti dal laminatoio per la
produzione a caldo di nastri in acciaio inossidabile 409 al fine di esaminare
l’effetto della composizione chimica sulla resistenza alla deformazione. I risultati hanno dimostrato che la resistenza alla deformazione dipende sensibilmente dalla possibilità di stabilizzare l’austenite tramite la composizio-
62
ne del materiale, suggerendo quindi che la frazione in volume di austenite
sia il fattore dominante nel controllo della resistenza alla deformazione. Il
rapporto di resistenza alla deformazione (DRR), è stato definito come un
rapporto tra la resistenza alla deformazione di una microstruttura bifasica
(ferritica + austenitica) rispetto a quella di una microstruttura pienamente
ferritica. La dipendenza del DRR dalla frazione in volume di austenite si
è dimostrata lineare, come è risultato anche dalle prove di deformazione
da compressione del piano effettuate su campioni di laboratorio con diverse
frazioni in volume di austenite. L’implicazione di questo risultato è che,
durante la laminazione a caldo di nastri di acciaio inossidabile 409 con
microstruttura a bifasica, questi acciai risultano predisposti alla deformazione in modo uniforme nella direzione di laminazione.

2 marzo 2009 << la metallurgia italiana

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